[ { "image_filename": "designv11_61_0001349_105994906x83385-Figure12-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001349_105994906x83385-Figure12-1.png", "caption": "Fig. 12 Drawing of a notch tensile test specimen", "texts": [ " Microvoids are formed as a result of particle-matrix decohesion or the cracking of second-phase particles. The process of microvoid growth involves considerable localized plastic deformation and requires the expenditure of large amounts of energy (Ref 18). Therefore, finer dimples in laser-welded specimens are indicative of lower ductility of these specimens with respect to BM. Notch tensile tests were performed with the objective of comparing the notch sensitivity of the FZ of LCW specimens with respect to BM. Figure 12 presents the drawing of the test specimen. Notch tensile testing of LCW and BM specimens magnified the difference between their ductility. Careful observation of the specimens during the course of testing revealed that in BM specimens gross plastic deformation was introduced (causing crack tip blunting) at an earlier stage than that in the LCW specimens. In these specimens, the crack extended with two zones of plastic deformation (appearing in the form of relief on the surface of the specimens under testing), inclined at about 45\u00b0 with respect to the direction of crack extension" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000201_iemdc.1999.769043-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000201_iemdc.1999.769043-Figure7-1.png", "caption": "Fig. 7. Displacements mused by the 5\" harmonic (1250 Hz) ofthe magnetic forces. (a) X axis. (b) Y axis.", "texts": [ " Figures 5 and 6 show, respectively, the calculated and measured accelerations as a function of the frequency. The ANSYS mechanical sohare was used to obtain these results. Observing the calculated and measured resonant peaks the main peaks correspond to odd harmonics of the one can notice that: magnetic forces. 0 the calculated acceleration peak values agree well with the measured ones. The displacements in X and Y directions caused by the 5* harmonic (1250 Hz) of the magnetic forces are presented in Fig. 7(a) and 7@), respectively. V. CONCLUSIONS In this work the forced response of a Switched Reluctance Motor is obtained by the Modal Superposition Methd The local magnetic forces in the stator teeth are calculated by a method based on the Maxwell Stress Tensor. Experimental Modal Analysis method is employed to obtain viscous damping. A good agreement between the calculated stator accelerations and the m e a s d ones is obtained. REFERENCES [l] C.G.C Neves, R. Carl-, N. Sadowski, J.P.A. Baaos,; S.L" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000443_s00170-005-0009-x-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000443_s00170-005-0009-x-Figure3-1.png", "caption": "Fig. 3 The local coordinates of spatial parallel mechanism with 4-PUU", "texts": [ " 1 shows, where z-axis is perpendicular to the guide plane P1P2P3P4, the origin is on the midline of the two guides, x is superposed with the midline of the two guides and y-axis is perpendicular to the two guides. According to the method presented above, the local coordinate system oc xc yc zc is shown in Fig. 2, where zc-axis is perpendicular to the plane of the manipulator M1M2M3M4, the origin is superposed with the geometric center of M1M2M3M4, xc and yc axes are parallel to the two orthogonal sides of the manipulator. Firstly, we will analyze the DoF of the manipulator as Fig. 1 shows. According to [9], we can find the dimension of the constraints spaces that all of the reciprocal screws, shown in Fig. 3, can be spanned is: d \u00bc dim span $rB1P1M1 $rB2P2M2 $rB3P3M3 $rB4P4M4 8>< >: 9>= >; \u00bc 2: (11) Therefore, F \u00bc 6 d \u00bc 6 2 \u00bc 4: (12) So, the manipulator shown in Fig. 1 has four DoFs, including three orthogonal translational movements and one rotational moment around zc-axis. Now we can select (xc yc zc) and the rotational angle around zc-axis, \u03b2, as the stance parameters. If we presume the length ofM1M2 to be 2a and the length ofM1M4 to be 2b, the local coordinates of the four vertexes of the manipulator can be obtained: rLM1 \u00bc b a 0\u00bd T rLM2 \u00bc b a 0\u00bd T rLM3 \u00bc b a 0\u00bd T rLM4 \u00bc b a 0\u00bd T : 8>>< >>: (13) The transform matrix from the local coordinate system to the absolute one is: A \u00bc cos sin 0 sin cos 0 0 0 1 2 4 3 5: (14) With Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003445_9781118361146.ch7-Figure7.30-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003445_9781118361146.ch7-Figure7.30-1.png", "caption": "Figure 7.30 Diagram showing the stator and rotor of an induction motor", "texts": [ " However, as we have seen, alternating current can easily be generated using an inverter, and in fact the inverter needed to produce the alternating current for an induction motor is no more complicated or expensive than the circuits needed to drive the BLDC motors or SRMs we have just described. So, these widely available and very reliable motors are well suited to use in electric vehicles. The principle of operation of the three-phase induction motor is shown in Figures 7.30 and 7.31. Three coils are wound right around the outer part of the motor, known as the stator, as shown in the top of Figure 7.30. The rotor usually consists of copper or aluminium rods, all electrically linked (short-circuited) at the end, forming a kind of cage, as also shown in Figure 7.30. Although shown hollow, the interior of this cage rotor will usually be filled with laminated iron. The three windings are arranged so that a positive current produces a magnetic field in the direction shown in Figure 7.31. If these three coils are fed with a three-phase alternating current, as in Figure 7.23, the resultant magnetic field rotates anti-clockwise, as shown at the bottom of Figure 7.31. This rotating field passes through the conductors on the rotor, generating an electric current. A force is produced on these conductors carrying an electric current, which turns the rotor" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002375_s11012-008-9143-5-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002375_s11012-008-9143-5-Figure5-1.png", "caption": "Fig. 5 Surfaces of loading: for i-cycle of loading (surface 1) and after following anneal (surface 2) (the plane (S1, S2) is shown); 3 surface of loading after (i \u2212 1)-cycle", "texts": [], "surrounding_texts": [ "In order to calculate the values of di via formula (49), it is necessary to calculate eM i (eM 1 = e1) on the ba- sis of formulae (48), (58) and (61). The calculation of eM i and ei is carried out with the following constant values: r = 4.44 \u00b7 1011 MPa2, \u03ba = 0.55, A0 = 20.5, B0 = 1.12 \u00b710\u22122, q = 5. The obtained results (denoted by superscript c) are put in Table 3. Let us calculate the steady creep rates of the specimens cut from the disks. The value \u03c3M P = HP M3 (\u03b1 = 0, \u03b2 = \u03c0/2, \u03bb = 0) from formula (63) gives the creep limit of the material of specimens cut from the R-disc. Then on the basis of formulae (64)\u2013(66), the steady creep rates \u03b5\u0307M R are calculated. Using the following constants, rP = 8.28 \u00b7 1010 MPa2, \u03baP = 0.31, C0 = 0.15, D0 = 29.9, the following results are obtained: \u03c3M P = 317 MPa, \u03b5\u0307M R = 1.55 \u00b7 10\u22124%/h (while the value obtained experimentally obtained is \u03b5\u0307R = 1.6 \u00b7 10\u22124%/h). The steady creep rate of the unhardened material determined from formulae (64)\u2013(66) at \u03c3P = 254 MPa is \u03b5\u0307U = 1.35 \u00b7 10\u22123%/h and \u03b5\u0307U /\u03b5\u0307M R = 8.71, which testifies to the considerable hardening of the material due to repeated mechanicalthermal treatment. For the S-disk, the distance to the tangential planes from formulae (53), (54) and (29) takes the value HP M1 = \u221a 2/3\u03c3P because W1 = 0, which in turn is true because exp(\u2212D1) \u2192 0 at \u03c3C1 = 838 MPa. Therefore, formula (20) gives the same result as for the case of the unhardened material. Consequently, one cycle of treatment under large stress at plastic deforming does not result in an increase of material strength concerning creep deformation, what agrees with the experimental results." ] }, { "image_filename": "designv11_61_0002802_s0018151x09010167-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002802_s0018151x09010167-Figure4-1.png", "caption": "Fig. 4. The exponential law of distribution of power of friction in the volume of surface layers.", "texts": [ " The heat release on friction contact is due, on the one hand, to breakdown of adhesion bonds in the actual areas of contact and, on the other hand, to deformation of microvolumes of roughnesses [11]. The second (deformation) component causes volume heat release in the surface layers interacting bodies. It has been proved [12] that the deformation component of heat release produces a significant fraction of total power of friction and significantly affects the temperature distribution in the surface layers. We assume that the volume density of power of heat release under friction is distributed in the contacting bodies by the exponential law (Fig. 4), (14) Here, h1 and h2 denote the thickness of surface layers of heat release characteristic of the contacting bodies and of the mode of friction. In view of (14), the problem on thermal conductivity for a system of two half-spaces takes the form \u03c91 x t,( ) \u03b1 h1 ----q t( ) x h1 ----\u2013 \u23a9 \u23ad \u23a8 \u23ac \u23a7 \u23ab ,exp= h1 0, x 0, t 0;> > > \u03c92 x t,( ) 1 \u03b1\u2013( ) h2 ---------------- q t( ) x h2 ---- \u23a9 \u23ad \u23a8 \u23ac \u23a7 \u23ab ,exp= h2 0, x 0, t 0.><> 128 HIGH TEMPERATURE Vol. 47 No. 1 2009 BELYAKOV, NOSKO (15) The solution of problem (15) in images of integral Laplace transform has the form (16) We can use the \u201coriginal-image\u201d tables [6] and find the originals \u03d5i(x, \u03c4) of the respective images \u03a6i(x, s) determined using relations (16)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000052_cdc.1999.832740-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000052_cdc.1999.832740-Figure6-1.png", "caption": "Figure 6: Shortest path to a line w.", "texts": [ " Parameter a for path type ( I X ) must be such that the robot orientation along the straight line segment Si is perpendicular to the line w (note that the same condition could be used to determine parameters a and b for path types ( V I I ) and ( V I I I ) ) . The determination of i allows to get, from (6), the expression of zf(C,Of,m) and yf(C,Bf,m) to put in Eq. (8) which implicitly defines the function [(e,). The extremals of this function will provide the e* o timal solution of the considered problem. The function [(e,) exists and is unique and C\u2019 within given closed intervals of variation of Of [22]. f. In Fig. 6 there is an example of the output produced by our software for the computation of the shortest path to a line. The starting configuration is the origin, the gray lines are the traces of the contact surfaces relative to each vertex of the robot on the plane ( x , y, 6';) and t , U, v are the parameters of the path. The closed curve centered at the origin is the set of the accessible positions reachable by paths of length l (0; ) given in left top corner of the figure with the orientation 0; also given in the figure" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001525_2006-01-0460-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001525_2006-01-0460-Figure3-1.png", "caption": "Figure 3 Simulation set up for FE model validation in bending (left) and shear (right) load at an impactor speed of 20 Km/h. Passive behavior of muscles is modeled by assigning a minimum value of 0.005 as initial activation level to each muscle.", "texts": [ " As a first step it was considered important to ascertain that the passive model validates against known experimental corridors. In-vivo passive response from a cadaver was modeled by setting the minimum activation level of 0.005 for each muscle and deactivating the reflex action. As the current FE model is an extension of Chawla\u2019s model, we have compared it with the simulation results reported by Chawla et al., (2004) and the experimental results of Kajzer et al., (1999). Simulations for the validation were performed using PAM-CRASHTM, an explicit dynamic solver. Figure 3 shows the set up used to perform the simulation to validate the FE model. The sacrum and two locations of the femur were fixed (as shown in Figure 3) and a pre-load of 400 N representing body weight was applied at the top of the femur. The impactor force, lower and upper tibia displacements at locations (P1 and P2) and the ligament forces were compared with the simulation results of Chawla et al. (2004) and test results of Kajzer et al. (1999). Comparison of passive loading cases Figure 4 shows the impactor forces in shear loading simulations with inactivated muscles and those reported by Chawla et al., (2004) and Kajzer et al., (1999) for the impactor speed of 20 km/h", " However, the impact force correlates well with the forces reported by Chawla et al., (2004) and Kajzer et al., (1999) (correlation of 0.91 and 0.95 respectively, as obtained using the \u201ccorrel\u201d function in Microsoft ExcelTM). Impactor Force in Shear 0 1000 2000 3000 4000 0.000 0.008 0.015 0.023 0.030 Time (sec) C on ta ct F or ce ( N ) Present Study Chawla et al. 2004 Kajzer et al. 1999 Figure 4 Comparison of Impactor force in shear loading. Figure 5 compares the lower and upper tibia displacements (at P1 and P2 in Figure 3) for shear loading. The lower displacements curves match with the displacement curves of Chawla et al. and Kajzer et al. shear test with correlations of 0.99 and 0.97 respectively. The upper tibial displacements show correlations of 0.99 and 0.91 respectively with respect to Chawla et al., (2004) and the experimental results respectively. The upper tibial displacements deviate slightly from the experimental results after about 15-20 ms. However, these values are very sensitive to the point chosen for recording the displacement as significant tibial rotations are observed during this phase" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000915_bfb0015079-Figure1.3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000915_bfb0015079-Figure1.3-1.png", "caption": "Fig. 1.3. Impact and switching on phase plane portrait", "texts": [ " When the system is in mode C then the swing limb will contact the ground and prompt a chain of events that may lead to stable progression. The importance of this contact event can be better understood if the motion is depicted in the phase space of the state variables. We simplify the present discussion by describing the events that lead to stable progression of a biped for a single degree of freedom system, however this approach can be generalized to higher order models. The phase plane portrait corresponding to the dynamic behavior that is described in the previous section is depicted Fig. 1.3a. The sample trajectories corresponding to each mode of behavior are labeled accordingly. The vertical dashed lines represent the values of the coordinate depicted in the phase plane for which the contact occurs. For the motions depicted in this figure, the only trajectory that leads to contact is C. The contact ew~nt for this simple model produces two simultaneous events: 1. Impact, which is represented by a sudden change in generalized velocities. 2. Switching due to the transfer of pivot to the point of contact. At this instant, the role of the limbs are exchanged, the old stance limb becomes the new swing limb and the old swing limb becomes the new stance limb. This exchange is reflected by sudden changes in the values of generalized positions and velocities. The combined effect of impact and switching on the phase plane portrait is depicted in Fig. 1.3b. As shown in the figure, the effect of the contact event will be a sudden transfer in the phase from point 1 to point 2, which is generally located on a different dynamic trajectory than the original one. If the destination of this transfer is on the original trajectory, then the resulting motion becomes periodic, Fig. 1.3c. This type of periodicity has unique advantage when the inverted pendulum system represents a biped. Actually, this is the only mode of behavior that this biped can achieve progression. The most striking aspect of this particular mode of behavior is that the biped achieves periodicity by utilizing only a portion of a dynamic trajectory. The impact and switching modes provide the connection between the cyclic motions of the kinematic chain and the walking action. We can clearly observe from the preceding discussion that the motion of a biped involves continuous phases separated by abrupt changes resulting from impact of the feet with the wMking surface" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001305_s00170-005-0257-9-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001305_s00170-005-0257-9-Figure1-1.png", "caption": "Fig. 1 Schematic diagram of the new SL system. 1-laser, 2-AOM, 3-beam expander, 4-galvanometric scanner, 5-F-Theta objective, 6-scraper, 7-elevator, 8-AC servomotor, 9-linear encoder, 10-vat, 11- rollers pump, 12-computer", "texts": [ " The thinner the layer thickness and the better the quality of the recoating process, the higher the resolution in the vertical direction. So the diameter of laser light spot on the scanning plane and the thickness of recoating process are two key parameters that determine the resolution of an SL system. In order to improve the resolution of the SL process, it is necessary to reduce the diameter of laser light spot in the scanning plane and the thickness of recoating process when the SL system is designed and built up. The principle components of the high-resolution SL system have been demonstrated in Fig. 1. The new RP system consists of a single mode He-Cd laser with TEM00 beam quality, an improved small laser light spot scanning system, a novel recoating system and a control system. The laser light spot with a diameter of 12.89 \u03bcm is focused on the focal plane using the optical system, and the resin layer with a thickness of 20 \u03bcm can be obtained by using a rollers pump to supply resin and control the position of resin surface in the recoating process. Therefore, the novel SL system has higher resolution than the conventional SL system, and is competent for fabricating small-size objects with intricate microstructures" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000278_1-84628-179-2_7-Figure7.4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000278_1-84628-179-2_7-Figure7.4-1.png", "caption": "Fig. 7.4. Side view of the T-Wing.", "texts": [ " Collective blade pitch control is still required to marry efficient high-speed horizontal flight performance with the production of adequate thrust on take-off, however, even this complication can be deleted, with little performance penalty, if high dash speeds are not required. \u2022 The current vehicle uses a canard to allow a more advantageous placement of the vehicle centre of gravity (CG). \u2022 Two separate engines are used in the current design though the possibility of using a single engine with appropriate drive trains could also be accommodated. The basic configuration of the T-Wing is presented in Figure 7.3 and Figure 7.4, with a diagram in Figure 7.5 showing some of the important gross geometric properties. The aircraft is essentially a tandem wing configuration with twin tractor propellers mounted on the aft main wing. The current T-Wing vehicle is a technology demonstrator and not a prototype production one. The aims of the T-Wing vehicle programme are to prove the critical technologies required of a tail-sitter vehicle before committing funds to full-scale development. The most important aspects of the T-Wing design that have to be demonstrated are reliable autonomous hover control and the ability to perform the transition manoeuvres between horizontal and vertical flight" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000740_robot.2004.1308050-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000740_robot.2004.1308050-Figure1-1.png", "caption": "Figure 1: Fixed base single m", "texts": [ " Then, in section 4, the mentioned self-coordination techpique is introduced, while showing how kinematic-independence can be obtained only at the cost of adopting iterative, multi-rate data exchange between the composing subsystems. Results concerning the effectiveness of the proposed self-coordination method are subsequently provided in section 5. Finally some concluding remarks are reported in section 6. . 11. FIXED BASE-SINGLE ARM CONTROL WlTH - SINGULARITY AVOIDANCE Let us consider a redundant (i.e. #dof %) fixed base single arm, for instance similar (without loss of generality) to the commonly used one reported in fig.1 (with\u2019 the wrist constituted by a 3-dof rotational joint, typically of Euler or Roll-Pitch-Yaw type). In such figure, frame represents the so called \u201cgoal : h e \u201d , which has to be reached (in position and orientation) by the so~called \u201ctool kame\u201d of the manipulator. Position and orientation of frames and

are respectively described 6y the Transformation matrices T, and X;T, with the former given as external reference input, and the latter real time computed by a suitable Direct-Geometry module, as the result of the product 07803-8232-3/04/$17" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003192_s11768-009-8102-6-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003192_s11768-009-8102-6-Figure3-1.png", "caption": "Fig. 3 Model of sonar sensor for line features.", "texts": [ " In contrast, false features produce sets of incoherent sonar arcs and thus can be easily spotted. The conclusion from this example is compelling and well known: using a careful sonar model is crucial to adequately interpret sonar returns. This work is restricted to indoor polygonal environments, and we will use line features to represent the walls. A wall, provided that its surface is smooth, will only produce a sonar return for sensor Sj when the perpendicular to the wall is contained in the sonar emission cone, as shown in Fig.3. \u03b8Sj \u2212 \u03b2 2 \u03b8R k \u03b8Sj + \u03b2 2 , (1) where \u03b2 is the angle of beam width. The distance actually given by the sensor \u03b2Sj will correspond to the perpendicular distance from the sensor to the line. Conversely, given the sensor location and the measured distances, all possible lines giving such a return are tangent to the arc depicted in Fig.3. In this paper, Hough transform is used to extract lines and points from the sonar data. Using the Hough transform for feature extraction, lines are extracted and represented in the robot coordinate frame R, located in the central robot position, using parameters \u03b8R s and \u03c1R s , defining the line ori- entation and its distance to the origin. With the definition of the Hough space, the basic voting algorithm for line features is shown in Table 1. The local maxima having a number of votes above a certain threshold are points corresponding to the line features in Cartesian plane" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000010_s0167-8922(03)80068-8-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000010_s0167-8922(03)80068-8-Figure6-1.png", "caption": "Figure 6. Description of the test arrangement: (A) microscope, (B, C) piezoelectric accelero-meter, (D) steel ball, (E) glass disc, (F) moving ball motor, (G) press screw, (H) clamped steel beam, (I) glass disc beating support, (J) rigid frame, (K) stem connected to the shaker.", "texts": [ " This method has been validated using a mineral oil with known viscosity and pressure viscosity coefficient. In the case of the clamped beam system, a shaker is used to normally excite the contact. A dynamic force can then be applied superimposed to the static load induced by the press screw. For this end, a signal generator and a power amplifier are used. Two piezoelectric accelerometers are mounted on the support of the ball allowing measuring both the vertical and tangential dynamic responses of the contact. The test arrangement is shown in Fig. 6. Finally, a piezoelectric force transducer is mounted between the shaker's stem and the support of the ball in order to measure the excitation force. With the main characteristics of the materials in contact given in table 1, the static contact compression fi, the maximum hertzian pressure P and the linearised contact stiffness k, function of the applied static load N can be evaluated considering a single ball plane contact. With a static load of 25 N, the maximum static contact compression 5 is 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002229_s026357470700389x-Figure9-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002229_s026357470700389x-Figure9-1.png", "caption": "Fig. 9. (a) Configuration leading to forward motion and (b) Configuration leading to backward motion.", "texts": [ " The optimal feasible direction dopt can be expressed in a vector form (doptx , dopty ) and the corresponding orientation angle \u03b8opt can be calculated as \u03b8opt = arctan(doptx , dopty ) (17) where \u2212\u03c0 < \u03b8opt \u2264 \u03c0 . We denote the angle between the current heading direction d and the optimal feasible direction as \u03b8r \u2208 [\u2212\u03c0, \u03c0] and define it as follows: \u03b8r = \u23a7\u23a8 \u23a9 \u03b8opt \u2212 \u03b8 if \u2016\u03b8opt \u2212 \u03b8\u2016 \u2264 \u03c0 \u03b8opt \u2212 \u03b8 \u2212 2\u03c0 if \u03b8opt \u2212 \u03b8 > \u03c0 \u03b8opt \u2212 \u03b8 + 2\u03c0 if \u03b8opt \u2212 \u03b8 < \u2212\u03c0 (18) If \u2016\u03b8r\u2016 \u2264 \u03c0 2 , the robot\u2019s heading has a positive projection on the optimal feasible direction, that is dT optd > 0. In this case the robot will move forward, and we define the actual direction dm = d (Fig. 9(a)). If \u2016\u03b8r\u2016 > \u03c0 2 , the direction opposite to the robot\u2019s heading \u2212d has a positive projection on negative gradient direction dopt, that is dT opt(\u2212d) > 0. In this case, the robot will move backward, and we define the actual moving direction dm =\u2212d (Fig. 9(b)). The control law of linear velocity for the robot is then given by: q\u0307 = k dm |dm| (19) where k is a constant used to determine the robot\u2019s speed. The control law for angular velocity is a simple linear control law intended to turn the robot as fast as possible towards the desired heading. The desired angular velocity is defined by: \u03b8\u0307 = k\u03b8\u03b8r (20) where k\u03b8 is a constant. In Fig. 10, we use a task of target reaching with obstacles to illustrate the performance of Newton method and Gradient method using the geometric transform" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003786_ramech.2011.6070458-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003786_ramech.2011.6070458-Figure2-1.png", "caption": "Fig. 2. Alternative paths in the presence of an obstacle.", "texts": [ " (13) 68 2011 IEEE 5th International Conference on Robotics, Automation and Mechatronics (RAM) We also use the derivative of (8) and (9) with respect to u as 2 321 32)(' uauaaux , (14) 2 321 32)(' ububbuy , (15) where ]1,0[u is a path parameter . B. Collision-Free Path Analysis In the presence of an a priori known obstacle, a procedure of placing a via-point is proposed. A property of Bezier path is used, that is, if the enveloping polygon of the path does not have intersection with any obstacle, the path is collision-free. As shown in Fig. 2, the obstacle is modeled as its environment, i.e., a circle with radius brrj , where jr and br represents the radius of the obstacle and the radius of the vehicle, respectively. According to the property, a path from the center of the j -th obstacle, i.e., jO , to the goal point, i.e., G , generated by only used 1 Bezier path cannot guarantee the path to be collision-free. At least a combination of two Bezier paths is needed to solve the problem. We need to construct paths between any two vehicles such that their distance is as far as possible. By such the way, any collision of any two vehicles, especially at the intersection point of their paths, can be avoided, as shown in Fig. 3. According to a property of three-degree Bezier curves, that is, the entire parts of the path are interiors of the area of polygonal area bounded by lines connecting its control points. It implies that if there is no obstacle intercepting the area, the vehicle\u2019s path is guaranteed to be collision-free. As shown in Fig. 2, two via-points P and Q are placed as connectors for four Bezier-paths which perform two path alternatives. Let 'r be the distance of jO and one of the via points , i.e., P or Q . 1P is defined as the intersection point of iPOb, and the boundary of the j-th obstacle, and 2P is the intersection between PG and the obstacle. Let }10,)1(),(|),{( ,b b, GOyxyxL iGO i , (16) }10,)1(),(|),{( 21 21 AAyxyxL AA , (17) and jOC be the closed set of the j-th obstacle. The decision to make new via-points is established if the following condition is satisfied. {} b, ji OGO CL , (18) {} 21 jOAA CL . (19) In the case of new via-points generation phase, the via-point candidates P and Q (see Fig. 2) are generated by solving the solution of an intersection point of the extension of lines 1b, PO i and 2G PO . 1P and 2P are the tangents to the boundary of the obstacle passing through the center of the vehicle and the goal point G , respectively. We assume that the environment is not extremely narrow. Therefore, the selected via-point is the one with closest distance to the goal point G . The obstacle-polygons intersection checking can be used to figure out the maximum distance of the free control points 0A and 3A to the initial and goal points, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003145_2009-01-3012-Figure14-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003145_2009-01-3012-Figure14-1.png", "caption": "Figure 14 \u2013 VRA Example", "texts": [ " [6] Visual Recurrence Analysis (VRA) \u2013 The software used in this study is VRA 4.9 [17]. Examples of squeal data analysis using VRA is illustrated below. More in-depth analysis and application of VRA will be published separately. Squeal analysis using VRA (example) \u2013 Below is a VRA analysis example, using the vibration signal from Figure 1 (a) \u2013 Test 1, Example 1, which has clear vibration amplitude growth over time. The time-delay used in the phase space plots was 3 data points (corresponding to approximately 39 microseconds at 25.6 kHz sampling frequency). Figure 14 (a) \u2013 shows a section of the vibration signal, including the transition range: From Figure 14 (b) we can see evolving patterns. Below are more detailed discussions of the \u201cphase space plot\u201d in Figure 14 (b), dividing it into several zones (cumulatively through time): 1. Random Zone (Figure 15): This is the center portion of the phase space plot in Figure 14 (b), which corresponds to the start of the time signal. The data points are random, without any order or pattern. The amplitudes of the signal are very small, as well. 2. Transition Zone 1 (Figure 16): This is the main chaotic transition range, where both determinism and patterns are visible. While the vibration amplitude increases (but not at constant speed), the patterns also evolve over time, implying ongoing complex dynamic interactions. Notice also that there is a sudden change in the spiral-out pattern, with much quicker amplitude increase, before it starts to level (at the end of this time period)", " Transition Zone 2 (Figure 17): There seems some amplitude / phase modulation, which could be due to changing force or force interactions. It could also be due to either frequency modulation (two modes) or frequency fluctuation (one single mode). This zone is perhaps less critical than Transition Zone 1, as it is likely too late \u2013 the \u201cpatient\u201d is already \u201csick\u201d. 4. Transition Zone 3 (Figure 18): After some more amplitude / phase modulation, it eventually settles into simple circles (periodic orbits) \u2013 only the initial outer circles are shown in Figure 18. The outer rim continues to grow (see Figure 14 (b) for whole outer rim), indicating continual growth of vibration amplitude. However, there is no more evolving pattern, hence nonchaotic. This is the least important zone. The \u201cpatient\u201d is indeed getting even sicker, but the dynamic disease has started much earlier. This paper attempts to discuss the complex \u201cfriction excitations\u201d, and propose some possibilities of squeal triggering mechanisms. We need to look beyond (not ignore) , in order to understand the complex \u201cfriction excitations\u201d. It does not mean that is not involved in friction, or that friction level is not involved in \u201cfriction excitations\u201d" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002236_2007-01-3740-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002236_2007-01-3740-Figure7-1.png", "caption": "Figure 7b: Single Cavity, Single Regime IVT", "texts": [ " The transmission is designed for a typical \u201centry level\u201d vehicle with circa 30 hp / 22 kW and 45 Nm in a vehicle with a gross vehicle weight (GVW) of circa 850 kg. Performance comparison of this concept to a 4 speed manual transmission in the target vehicle with identical vehicle weights and over the fuel economy test cycle provides the following results:- Hence, due to the ability of the variable drive transmission to optimise engine operation, the IVT delivers improved vehicle performance versus a manual transmission without compromising fuel economy. The transmission has now been realized in hardware and is under test (figure 7a and 7b). SECTOR VEHICLES The low-cost compact IVT has potential for the lowest power and torque vehicle applications. For higher power and torque applications, due to both efficiency and package considerations, the single cavity design becomes less attractive compared to twin cavity designs. Therefore when considering sub-A, A & B sector vehicles, a twin cavity Variator is viewed as superior in performance. Maintaining the focus on cost effective systems, the two roller per cavity Variator design is retained and the Variator is now incorporated into a CVT arrangement rather than an IVT arrangement" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002518_j.mee.2008.07.007-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002518_j.mee.2008.07.007-Figure7-1.png", "caption": "Fig. 7. Experimental transmittance curves of various LC cells for parallel (Tp) and the perpendicular (Tc) configurations used in the Soutar and Lu method.", "texts": [ " Nonetheless, using the experientially-obtained value [15], that is, K22 6.18 10 12 N, the magnitude of W/ can be estimated, thus obtaining W/ 4.08 10 6 J/m2. The applicability of polymer nano-grooves to LC alignment was compared with that of the conventional rubbing process. The surfaces of PA-coated substrates were rubbed with a conventional rubbing machine. Using a LC cell consisting of rubbed PA surfaces, the light intensity transmitted by the system in Fig. 5 was then measured. From the transmittance curve shown in Fig. 7(rub- bing/no grooves), the cell parameters and the azimuthal anchoring energy were estimated using the preceding method as / = 1.68 rad ( 96 ), b = 3.63 rad, wD = 0.78 or 0.78 rad ( 45 or 45 ), and W/ 5.03 10 6 J/m2. Upon comparison with the value for the PA nano-grooves it is found that the anchoring energy obtained for the rubbed PA is of the same order. Therefore, it can be suggested that the present nano-grooving approach can replace the conventional rubbing process for inducing LC alignment. Additionally, nano-grooving and conventional rubbing were simultaneously conducted. Thus, two different LC cells, one of which was rubbed parallel to the groove direction (rubbing || grooves) and the other perpendicular to the groove direction (rubbing \\ grooves) were prepared. In the transmittance curves shown in Fig. 7, an increase and decrease in the amplitude, which varies depending on the molecular twisting and phase retardation, were observed for rubbing || grooves and rubbing \\ grooves, respectively. This can be explained from the viewpoint of the LC anchoring stability on the surfaces. In the case of the rubbing || grooves, the artificial nano-grooves existing parallel to the rubbing direction synergically affect LC alignment, resulting in stronger anchoring. In contrast, on the surface of the rubbing \\ grooves, LC molecules might be mutually disturbed as they orient either to the rubbing direction or to the nano-groove direction, and therefore the anchoring stability would decrease" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001704_20050703-6-cz-1902.02304-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001704_20050703-6-cz-1902.02304-Figure3-1.png", "caption": "Fig. 3. The 'Mechatron II' educational platform in its basic \"stripped\" form, before the addition of sensors, actuators and interface circuitry.", "texts": [ " Each course is designed as a series of assignments (problems) of increasing complexity. Any of these problems can be performed using the same basic equipment: a small wheeled platform. The students are asked to impart a specific behaviour or function to the robotic platform that is fully maintained by the Laboratory (Chamilothoris, and Voliotis, 2003; Gopalakrishnan, et al., 2004). The features of the educational platform and the process of its development have been reported in (Chamilothoris, 2002; Chamilothoris, and Voliotis, 2003). The current form of the device, depicted in Fig. 3 and termed the 'Mechatron II', includes a differential wheel drive, power conditioning circuitry, a solder-less electronics board and a multitasking computer-on-a-chip device. Students have access to the equipment and the facilities of the Laboratory during normal course hours and also during 'open doors' periods, to the extent of approx. 12 hours per week. Also, student teams that reach an advanced stage of work towards developing a solution are given a 'private' Mechatron unit (Fig. 4) and accessories for exclusive use by the team up to the final examination" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001606_iros.2004.1389974-FigureI-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001606_iros.2004.1389974-FigureI-1.png", "caption": "Fig. I. Six DOF mmipulmr.", "texts": [ " Given that the damping control seeks only to prevent any movement of the links which is not directly determined by the task space control, measuring the differences in joint angles generated by the two types of control demonstrates which joints are most affected by the use of the gradient control. A contour of the capability with respect to these joints gives a clear picture as to the behaviour of the gradient with respect to the capability topography. Refemng to Fig. 5 , the two joints with the highest average variation between the two controls, q2 and 96, are chosen as the contour variables. Figs. 6, 7, and 8 show the contours of acceleration capability taken from Fig. 4 for a variety of highly varying joint angles. The one point common to both paths in each figure is the starting position. The paths outlined indicate that in all cases the damping control stays firmly within low-capability areas, whereas the gradient control path begins ascending the nearer peaks and attempts to raise the capability. Also, in every case the gradient path slowly descends from the higher capability areas and begins approaching the lower contours. The descent of the gradient path back toward the low- capability areas is caused by increasing end-effector proximity to the workspace boundary, an effect dictated by the task space control" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003372_978-3-642-25486-4_28-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003372_978-3-642-25486-4_28-Figure1-1.png", "caption": "Fig. 1. Scheme of parallel indexing cam mechanism", "texts": [ " Then by applying unilateral contact method clearance model and dynamic model of parallel indexing cam with clearance are developed. Numerical integration is applied to obtain the responses. Through the numerical results we analyze effects of clearance on dynamics response and the reasons for impact vibration on turret. It provides basis for obtaining the fault characteristic of indexing cam and implementing dynamic model-based fault diagnosis. The basic structure of parallel indexing cam mechanism is shown in Fig. 1. Profile of parallel indexing cam can be generated by the method of Exponential Product Formula. The general frame is built in Fig. 2, in which origin of frame is located in center of turret, axis z is the rotation direction of turret and axis x is defined as line from center of turret to cam center. In order to describe the motion of parallel indexing cam, skew is defined by rotation vector and coordinate of any point in rotation axis. The skew parameters are shown as table 1. where a is the center distance between cam axis and turret axis; r is the radius of roller in turret" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003445_9781118361146.ch7-Figure7.26-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003445_9781118361146.ch7-Figure7.26-1.png", "caption": "Figure 7.26 A 100 kW, oil-cooled BLDC motor for automotive application. This unit weighs just 21 kg. (Photograph reproduced by kind permission of Zytek Ltd.)", "texts": [ " They are very widely used in computer equipment to drive the moving parts of disc storage systems and fans. In these small motors the switching circuit is incorporated into the motor with the sensor switches. However, they are also used in higher power applications, with more sophisticated controllers (as of Figure 7.21), which can vary the coil current (and hence torque) and thus produce a very flexible drive system. Some of the most sophisticated electric vehicle drive motors are of this type, and one is shown in Figure 7.26. This is a 100 kW, oil-cooled motor, weighing just 21 kg. These BLDC motors need a strong permanent magnet for the rotor. The advantage of this is that currents do not need to be induced in the rotor (as with, for example, the induction motor), making them somewhat more efficient and giving a slightly greater specific power. Permanent magnet synchronous motors which are a type of BLDC motors are increasingly used in electric vehicles. Modern electronics allow the supply frequency to be continuously varied so that it can be used to control the motor speed and hence the vehicle speed" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000226_eurbot.1999.827617-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000226_eurbot.1999.827617-Figure4-1.png", "caption": "Figure 4: Coordinate systems of PRIAMOS and SPIKE", "texts": [ "100mrn] and cy,P,y E ( - 2 5 O . . . 2 5 O ] . The gradient of the spindle is s = 2mm. The number of revolutions of the motors is U = 3000 6 = 50 3 . Therefore the velocity of the servos is Vma, = 100 y. This velocity is reached in At = 0.5s, so the maximum 2Stewart-Platform of the IPR KarlsruhE acceleration of the servos is i 7 . s At - - a m a x - m mm = 200- = 0.2- S2 S 2 3.3 Combination The combination of PRIAMOS and SPIKE is a highly dynamic system, since both parts have no restrictions in their kinematics. Figure 4 shows a possible application carrying a glass filled with liquid. The coordinate systems will be described in section 4. The mobile robot generates only accelerations in 2,- y, and &-direction. The Stewart Platform can generate accelerations in all six directions. If sensors are used instead of odometric calculations, the motion of the vehicle can also be described by six accelerations. The parameters of the robot position are z, y and cy, the parameters of the platform position are z, y, z , j , p and P with 3 corresponding to cy", " The algorithm of the filter is like this: The parameters of the vehicle movement X , y , ? , j , p , r , j , p , + , j , p , r are transformed to the real coordinates of the platform acceleration X P , i i P , Z p , bxP,wyP,wzP , The washout filter calculates the position and orientation of the platform Z N , YN, Z N , P x P , PyP, ( p e p , PxL, PyL, Y'tL These values serve for computing a 4x4 matrix describing the actual position and orientation of the platform. The inverse kinematics deliver the length of the legs. Figure 4 shows the relationship between all coordinate systems. The relationship between W and V determines the movement of the robot in the real world, that between N and P the movement of the platform. If the acceleration vector in C referring to n/ is the same as in V referring to W except for the sign, the vector in L referring to W is zero. In this case the payload is not affected by any translational accelerations. The original filter was based on experiences of Baarspul [Baa89, Baa771, but with only four degrees of freedom" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001766_ichr.2004.1442673-Figure13-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001766_ichr.2004.1442673-Figure13-1.png", "caption": "Fig. 13. Fhtation Angle of Waist and Pitch Angle of Swing Leg", "texts": [ " In order to effectively cancel the stance foot torque by the waist rotation, the waist should be counterbalanced to the largest factor, i.e., the leg motion. The angle of the waist rotation is obtained as follows coupling with the pitch angle of the hip joints: Waist fixed Propose Antiphase (Eq- (2)) 0 c t- t t where k=O.l75[tad](lO.O[deg]) is the maximum amplitude of the waist rotation which is used to obtain the same amplitude of the standard walk. 8 ~ ( t ) and O,(t) are pitch angle of both swing leg shown in Fig.13, 8,,,=0.5[rad](28.6[deg]) is the maximum of ( B R ( ~ ) - t9~( t ) ) in one walking cycle. By using (2), the antiphase rotation is generated for the proposed walk compared with the standard walk as shown in Fig.14. The stance foot torque due to the acceleration of the swing leg is canceled by the waist angular momentum. 4.1. Comphson of Stance Foot Torque The effect of the waist rotation for reducing the stance foot torque is confirmed both by simulation and experiment. Fig.15 shows the body posture at 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001932_s00170-007-1295-2-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001932_s00170-007-1295-2-Figure1-1.png", "caption": "Fig. 1 The principle of selective laser sintering (SLS)", "texts": [ "eywords Selective laser sintering . Hot isostatic pressing . Cold isostatic pressing . Relative density Selective laser sintering (SLS) is a kind of rapid prototyping (RP) technology, whose principle is shown in Fig. 1. Any desired shape can be fabricated by SLS without the application of any tooling, so it has been often used for manufacturing complex metal parts [1\u20133]. In manufacturing metal parts by SLS, indirect SLS is usually adopted, in which high-melting-point metal powders are often bonded by binder powders so that their relative densities and mechanical performance are lower. Consequently, SLS metal parts are degreased, vacuum sintered, and infiltrated with a low-melting metal for improving their performance [4]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001054_bfb0015074-Figure2.1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001054_bfb0015074-Figure2.1-1.png", "caption": "Fig. 2.1. Two robots holding an object", "texts": [ " the constraint forces/moments on the object derived in Sect. 2. are elaborated; they are parameterized by external and internal forces/moments. In Sect. 4. a hybrid position/force control scheme that is based on the results in the previous section, is presented, before load-sharing control being discussed. Advanced topics in Sect. 5. are mainly those of research in the author's laboratory. This chapter is concluded in Sect. 6. 2. D y n a m i c s o f M u l t i r o b o t s a n d C o o p e r a t i v e S y s t e m s Consider the situation depicted in Fig. 2.1 where two robots hold a single object. The robots and the object form a closed kinematic chain and, therefore, equations of motion for the system is easily obtained. A point here is that the system is an over-actual;ed system where the number of actuators to drive the system is more than the number of degrees of freedom of the system. Therefore, how to deal with the constraint forces/moments acting on the system becomes crucial. Here, we formulate those as the forces/moments that the robots impart to the object" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000619_j.triboint.2003.11.011-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000619_j.triboint.2003.11.011-Figure1-1.png", "caption": "Fig. 1. Experimental apparatus.", "texts": [ " The technique has been further improved by Jolkin and Larsson [13]. The aim of this paper is to develop a Fast Fourier Transform (FFT)-based technique for relatively simple and fast determination of pressure distribution from measured EHD lubricant film thickness. This approach is validated on dry Hertzian contact and tested out using experimental results for oil-lubricated EHD contact. The experimental apparatus is classical and have been already described elsewhere [14]. It comprises three main parts as Fig. 1 shows: a conventional optical tests rig, a reflected light microscope with colour video camera and a personal computer. The film thickness was measured in a concentrated contact formed between a transparent disk and a highly polished steel ball of 25.4 mm in diameter. Depending on the expected pressure range, the transparent disk could be made of glass or sapphire. The composite roughness of such a contact is 3 or 4 nm depending on the disk material. For measurement purpose, the transparent disc is coated on its underside with a thin semi-reflective chromium layer" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003633_s10846-012-9726-1-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003633_s10846-012-9726-1-Figure4-1.png", "caption": "Fig. 4 Reference frames and UAV Euler angles. a Earth frame and UAV body frame. b Roll, pitch and yaw [16]", "texts": [ " The aileron input can be defined as the substraction of the deflections of left and right ailerons, \u03b4a = 1 2 (\u03b4al \u2212 \u03b4ar). (1) Consequently, the input vector reduces to \u03b4 = [\u03b4t \u03b4a \u03b4e \u03b4r]T . To describe the attitude and motion of a 6-DOF (degree of freedom) UAV in 3D space, the 12 state variables are 3 position states : P = [x y z]T 3 velocity states : V = P\u0307 = [u v w]T 3 angular position states : = [\u03c6 \u03b8 \u03c8 ]T 3 angular velocity states : = \u0307 = [p q r]T where (x, y, z) are the coordinates of the aircraft in earth-fixed inertial reference frame, F0, whose origin is at the defined home location on map, Fig. 4a. The physical meaning of \u2212z is the flying height. (u, v, w) are the velocities of the UAV along x, y, and z axis respectively. (\u03c6, \u03b8, \u03c8) are the roll, pitch and yaw angles, illustrated in Fig. 4b. (p, q, r) are the roll, pitch and yaw rates. The modeling objective is to find the state space representation, Eq. 2, or transfer function representation, Eq. 3, of the relationship between control inputs \u03b4 and the state variables X = [x y z u v w \u03c6 \u03b8 \u03c8 p q r]T . where A and B are the coefficient matrix of linearized differential equations of motion. X = G(s)\u03b4 (3) where G(s) = \u23a1 \u23a2\u23a2\u23a2\u23a3 G11 \u00b7 \u00b7 \u00b7 G14 G21 \u00b7 \u00b7 \u00b7 G24 ... . . . ... G121 \u00b7 \u00b7 \u00b7 G124 \u23a4 \u23a5\u23a5\u23a5\u23a6 is the matrix of transfer functions for each inputoutput channel" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001747_4-431-31381-8_12-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001747_4-431-31381-8_12-Figure1-1.png", "caption": "Fig. 1. Self-excited mechanism with bent knee and cylindrical foot", "texts": [ " Therefore, we try to apply a bent knee angle to the support leg in order to make it walk faster. In the next section, the analytical model and its basic equations of locomotion will be introduced. In section 3, we show the typical simulated results of stable biped locomotion on level ground with and without a bent knee angle and foot and the convergence characteristics of the self-excited walking in relation to the bent angle. Next, we present the calculated results of the effect of the knee bent angle with and without a foot radius on the walking performance. Figure 1 shows the biped mechanism that walks with a bent knee. We consider the biped walking motion on a sagittal plane. The biped model consists of only two legs and does not have a torso. The two legs are connected in a series at the hip joint through a motor. Both legs have a thigh and a shank that are connected at the knee joint. We assume that the biped has knee brakes so that the knee can be locked at any bent angle after the knee collision of the swing leg. The support leg does not extend fully but retains some flexion during the stance phase" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002270_978-3-540-88513-9_19-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002270_978-3-540-88513-9_19-Figure2-1.png", "caption": "Fig. 2. Schematics of 6-SPS Parallel Manipulator", "texts": [ " (1) could be changed into its homotopy function as ( , ) ( ) (1 ) ( ) 0H x t tF x t G x= + \u2212 = where [0,1]t \u2208 Therefore, we have two following boundary conditons ( ,0) ( ) 0H x G x= = ( ,1) ( ) 0H x F x= = Our objective is to solve ( , ) 0H x t = instead of ( ) 0F x = by varying parameter t form 0 to 1. Hence, Eq. (2) could be presented as follows ( ) ( ) ( ) ( ) ( ) ( ) 1 1 1 1 2 2 1 2 , , , , , , , , , , , , n n n n n n n n n n n n n n n n H x y H x y x y x x H x y H x y H x y y y H x y x y + + \u2202 \u2202\u23a1 \u23a4 \u23a2 \u23a5\u2202 \u2202\u23a2 \u23a5 \u2212 \u2212\u23a1 \u23a4\u23a1 \u23a4\u23a2 \u23a5\u2202 \u2202 \u23a2 \u23a5\u23a2 \u23a5\u23a2 \u23a5 \u2212 = \u2212\u23a2 \u23a5\u23a2 \u23a5\u2202 \u2202\u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5\u23a3 \u23a6 \u23a3 \u23a6\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 To avoid divergence, Wu [16] provided some useful homotopy function. By appropriately choosing the auxiliary homotopy function, we can obtain the solutions to Eq. (1). Fig. 2 shows a general 6-SPS parallel manipulator, constructed by connecting a hexagon mobile platform to a base with 6 SPS limbs. All the spherical joints Ai (i=1, 2, \u2026 6) on the base platform are coplanar, as well as all the spherical joints Bi (i=1, 2, \u2026 6) on the mobile platform are coplanar. The mobile frame O1-X1Y1Z1 is attached to the mobile platform, where O1 is the circumcenter of the mobile platform and axis O1Y1 through the midpoint of the line B2B3. The reference frame O-XYZ is attached to the based platform, where O is the circumcenter of the mobile platform and axis OY through the midpoint of the line A2A3", " Based on the homotopy theory, the author write the algorithmic program of the forward kinematics for 6-SPS parallel manipulator. In the following, numerical example will verify the feasible of the solution. In order to verify the proposed forward kinematics method, we first perform inverse kinematics to get the limb lengths for an actual pose. This pose will be serve as the desired pose for the forward kinematics. A general 6-DOF parallel manipulator is selected as an example for numerical simulation. Just as shown in Fig. 2. The radius of circumcircle the base platform and mobile platform are Ra=200 Rb=100, respectively. And 0 / 6\u03b8 \u03c0= , / 2\u03b8 \u03c0= . The initial pose of parallel manipulator is [10,30,180, , , ]12 12 12 T\u03c0 \u03c0 \u03c0 , and 6 limb lengths is 0 [229.5248,196.0392, 244.6751, 251.3970,229.8890,194.0281]TL = . The pose that we desired is [50, 30,230, , , ]6 3 4 T\u03c0 \u03c0 \u03c0\u2212 , and corresponding 6 limb lengths is 1 [195.9674,242.9199,371.4124,383.4943,272.9773, 200.9865]TL = . Our objective is to find a unique actual position and orientation ( 1 [ , , , , , ]TO x y z \u03b1 \u03b2 \u03b3= ) of the parallel manipulator under the input variables 1L " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002638_s10043-008-0026-8-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002638_s10043-008-0026-8-Figure8-1.png", "caption": "Fig. 8. Principle of movement of the body driving type of the optically driven actuator. (a) Initial state of the actuator before laser irradiation. (b) After laser irradiation. (c) After stopping laser irradiation. (d) Completion of movement.", "texts": [ " Figure 6 shows a photograph of an actuator on an acrylic surface; the unit is 5mm wide and 10mm high. When we irradiate from the left in Fig. 6, the actuator moves in the opposite direction. A He\u2013Ne laser with 10mW of power is used as a light source. Figure 7 shows the moving displacement of the actuator irradiated at frequencies of 1 and 2Hz. The velocities are 33.3 mm/s at 1Hz and 76.7 mm/s at 2Hz. The unit can be moved by direct driving using the PVDF cantilever. We have developed another type of actuator using the PVDF cantilever as a body. Figure 8 shows the principle of this body driving type of optical actuator. This type is easy to control, because irradiation can be done at the top of the actuator. Another advantage is that the driving direction can be controlled by changing the irradiated position of the cantilever. Figure 8(a) shows the actuator\u2019s initial position of the actuator. It consists of eight optical fibers that form the legs joined to a PVDF cantilever body. The legs are called \u2018\u2018w\u2019\u2019, \u2018\u2018x\u2019\u2019, \u2018\u2018y\u2019\u2019, and \u2018\u2018z\u2019\u2019 at positions of \u2018\u2018a\u2019\u2019, \u2018\u2018b\u2019\u2019, \u2018\u2018c\u2019\u2019, and \u2018\u2018d\u2019\u2019, respectively in Fig. 8. In Fig. 8(b), we irradiated by laser at the top of the body and the unit bends due to the pyroelectric and inverse piezoelectric effects. The front fibers, \u2018\u2018y\u2019\u2019 and \u2018\u2018z\u2019\u2019, are caused to slide on the base because the rear fibers, \u2018\u2018w\u2019\u2019 and \u2018\u2018x\u2019\u2019, are fixed for a friction force. Displacement of the front fibers can be controlled by changing the irradiated power. We can control the moving direction of the actuator by the irradiated position. After laser irradiation in Fig. 8(c), the front fibers indicated \u2018\u2018y\u2019\u2019 and \u2018\u2018z\u2019\u2019 stop at the point indicated by \u2018\u2018c0\u2019\u2019 and \u2018\u2018d0\u2019\u2019 which is in front of the initial point \u2018\u2018c\u2019\u2019 and \u2018\u2018d\u2019\u2019 because the frictional force of the leg exceeds the restorative force of the leg. The other legs slide on the base because the frictional force of the front legs is less than the restorative force of the actuator. Finally, in Fig. 8(d), the actuator can be moved for one cycle. Figure 9 is a photograph of the body driving type actuator on an acrylic surface; the body is 5mm wide and 20mm long. The legs are 500 mm in a diameter and 5mm long. We irradiate it from the top of the actuator; a 10mW He\u2013Ne laser is used as the light source. Figure 10 illustrates the displacement of the actuator by an irradiated frequency of 1Hz; the velocity is 20 mm/s. We investigated optical driving of an actuator constructed from a PVDF cantilever" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003117_ichr.2009.5379588-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003117_ichr.2009.5379588-Figure5-1.png", "caption": "Fig. 5. Artificial Potential Field forces and torques applied on the robot.", "texts": [ " The well-known relation that characterized this motion planning technique is: Pint a (Cp S + c; ~) (5) Tint Finty iP; - X r ) - Fintx (Py - Yr) (6) a = {I if interaction o otherwise where D and D or are dissipative factors used for the asymp totic stabilization of the system; Tint is the torque applied to the robot, depending on the interaction force and on its application point; l is the distance between P and H; a is a switching variable that turns to 1 when the humanoid robot and the door are close enough. Note that the motion of the robot depends also on the force Pap! and on the torque Tap! applied by the Artificial Potential Field (APF) that will be described in Section III-A. (8) (9) (10) U where Ci represents the position of one of the four corners. The repulsive profile, therefore, pushes the robot far away from obstacles and let the robot use its orientation to pass through 4 T rep = L Fix (Cix - X r ) - Fiy(Ciy - Yr) (11) i=O Concerning the orientation, see Fig.5, we assume that an attractive torque T att forces the robot to look always towards the subgoal first, and the goal after. T att depends on the position and velocity gains k op and k ov, in accord with the spring/dumper model. Also a repulsive torque T rep, due to the forces Fo, F I , F2, F3 applied in the four corners by the repulsive profile, influences the orientation of the robot; its computation is done as follows: where U is an adequate potential function, continuous, differ entiable and, usually, depending on the position of the robot" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000025_ecc.2003.7085301-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000025_ecc.2003.7085301-Figure4-1.png", "caption": "Figure 4: Bell-Hiller system with angular displacements", "texts": [ " The system consists of a so-called flybar (a teetering rotor placed at a 90o rotation interval from the main rotor blades and tipped on both ends by aerodynamic paddles) and a mixing device that combines the flybar flapping motion with the cyclic inputs to determine the cyclic pitch angle applied to the main rotor blades. The flybar and main rotor flapping motions are governed by the same effects, namely the gyroscopic moments due the helicopter roll and pitch rates. However, unlike the main rotor, the flybar is not responsible for providing lift or maneuvering ability. Thus, it can be designed to have a slower response and provide the desired stabilization effect. The notation used to describe the Bell-Hiller system is presented in Figure 4, where the mechanical arrangement for the X-Treme helicopter is reproduced. Due to the geometric constraint introduced by the mixing lever, the flybar flapping and rotor blade pitching motions are effectively combined. The equations of motion for the main rotor blade pitching can be written as [ \u03b8\u03081c \u03b8\u03081s ] +\u03a9A\u03b8\u0307 [ \u03b8\u03071c \u03b8\u03071s ] +\u03a92A\u03b8(\u03bc) [ \u03b81c \u03b81s ] = \u03a92B\u03b4(\u03bc) [ \u03b41c \u03b41s ] +\u03a92B\u03c9 [ p\u0304 q\u0304 ] +\u03a92B\u03bb(\u03bc) \u23a1 \u23a3 \u03bcz \u2212 \u03bb0 \u03bb1c \u03bb1s \u23a4 \u23a6 . (8) The blade pitching motion, in particular its response to helicopter shaft rotations, depends on the physical parameters of the Bell-Hiller system, namely the lever arms l1, and l2, the flybar radii R1 and R2, and the flybar Lock number defined as \u03b3f = \u03c1cfa0f ( R4 2 \u2212 R4 1 ) / I\u03b2f , (9) where \u03c1 is the air density, cf the paddle chord, a0f the paddle lift curve slope, and I\u03b2f the flybar moment of inertia" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003215_iecon.2008.4758119-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003215_iecon.2008.4758119-Figure4-1.png", "caption": "Fig. 4. Fault #3, deformation of the protective shield.", "texts": [ " Second, since in other researches one of the most investigated fault is the hole in the outer race [6] [9], we have reproduced also this single-point defect in order to have a comparison with the results obtained in the literature . In fact, a \u201cperfect\u201d fault like that shown in Fig. 3 cannot occur during the working of a bearing, but a similar defect can be caused by circulating currents, so this hole represents a magnification of defects that can really happen in the bearings. Then, we have investigated a fault which has not be considered in previous researches, i.e. a deformation of the protective shield (Fig. 4). This fault can be produced by errors during the assembly and can be considered as a cyclic fault, even if it does not produce effect like air-gap eccentricity. So, it is expected not to show particular changes in the current spectrum. Finally, we have produced a corrosion of the bearing, which can be caused by humidity of the environment and can be considered as a generalized roughness (non-cyclic fault, see Fig. 5). The experimental set-up consists of a 2.2 kW three-phase induction motor with two pole pairs, fed by the mains (400 V, 50 Hz) and coupled with a brake [13]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002320_978-3-540-77457-0_2-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002320_978-3-540-77457-0_2-Figure7-1.png", "caption": "Fig. 7. Simulation model", "texts": [ " Consequently, the force model in the parallel distributed model agrees with the experimental results. Note that in the radially distributed model, the force magnitude remains constant and the component perpendicular to the plate behind the fingertip reaches its maximum rather than its minimum at \u03b8p = 0. That is, the radially distributed model does not agree with the experimental results. Based on the parallel distributed model, we simulated grasping and manipulation by a pair of 1-DOF fingers with soft fingertips. Figure 7 shows a simulation model. Let \u03b8l and \u03b8r be rotational angles of the left and right fingers. Assume that the two fingers have the same dimensions. Let L be the length between the center of the hemispherical fingertip and the joint of the finger. A pair of fingers pinches a rectangular object of width Wobj. Let (xobj, yobj) be the positional vector and \u03b8obj be the orientation angle of the pinched object. The relative angle between the object and the left finger is given by \u03b8r \u2212 \u03b8obj, while the angle between the object and the right finger is given by \u03b8l + \u03b8obj" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002685_rnc.1378-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002685_rnc.1378-Figure6-1.png", "caption": "Figure 6. The motorcycle model.", "texts": [ " In this example, these bounds may appear very large, but this is justified by the fact that they must apply to any trajectory whose acceleration is bounded by 2.2m/s2 which is \u2016\u0308\u2016\u221e (Figures 4 and 5). 3. THE EXACT OUTPUT TRACKING PROBLEM FOR THE NON-HOLONOMIC MOTORCYCLE MODEL Consider the simple non-holonomic motorcycle model presented in [7]. The model kinematics are given by x\u0307 = v cos y\u0307 = v sin \u0307 = v (6) where (x, y) is the position of the motorcycle, is its leading angle with controls given by the speed v and the curvature (see Figure 6). Let \u2208C3(R,R2) be a periodic curve of period T such that \u2016\u0307(t)\u2016>0, \u2200t \u2208R; in order to track exactly this periodic trajectory the control must be v=\u2016\u0307\u2016 and =\u2016\u0307\u2016\u22121 d arg(\u0307) dt (7) Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc The problem is to show that there exists a value 0 of the roll angle such that starting with that value, the motorcycle does not overturn when it follows , with the controls given by (7). As shown in [7], the roll angle has to satisfy this equation: \u0308=h\u22121 [ g sin + ( h ( d arg(\u0307) dt )2 cos ) sin \u2212 ( \u2016\u0307\u2016d arg(\u0307) dt +b d2 arg(\u0307) dt2 ) cos ] (8) where g is the gravity acceleration, h is the height of the center of mass (when the vehicle is vertical), b is the distance of projection of the center of mass to the ground from the contact point of the real wheel" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003210_6.2009-1833-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003210_6.2009-1833-Figure3-1.png", "caption": "Figure 3. Wind disturbance direction", "texts": [ " American Institute of Aeronautics and Astronautics 5 The basic idea is to counter the wind disturbance by adding a feedforward control signal ff to the signal fb from the feedback controller before it enters the plant. This is shown in Fig 2. Since is the control input needed to keep the helicopter in the equilibrium state related to the wind disturbance trim condition, provided by the feedback controller from the given body velocity must be subtracted. This is illustrated in Fig. 2, where the plant is disturbed by a wind velocity as depicted in Fig. 3. Thus, effective values of feedforward control inputs would be given by ff ( 5 ) where is the trim condition based on disturbance velocity vector \u00a0 \u00a0 and is the trim condition based on current body velocity . For better understanding of the trim condition based on , Fig. 4 illustrates the flight condition that results from the application of . Now, assume a hover flight condition and that the wind disturbance is coming from the front of the helicopter. These assumptions do not cause loss of generality since the feedforward control concept is independent of wind direction, and trim values can be computed for any equilibrium flight condition" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002168_isie.2007.4374753-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002168_isie.2007.4374753-Figure7-1.png", "caption": "Fig. 7. Rotating magnetic field, eddy currents and antagonistic torques", "texts": [], "surrounding_texts": [ "In order to construct the proposed prototype rotor, a radically different process from the construction of a conventional squirrel cage rotor, is followed. It has been indicated that in those newly rotor, the sheets don't have any slot, so that in order to have them perfectly joined, and not leave air spaces between them, its shape must have a perfect definite profile. This profile is reached by the shaping with a tool or matrix to deform every piece or sheet, giving the precise shape that is described as follows. If we consider a rotor of interior radio R, formed by n sheets of a thickness e, being this thickness very small compared to R, and with an outside radio RE, we will have the trigonometric ratios in the figure 8." ] }, { "image_filename": "designv11_61_0000700_12.571089-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000700_12.571089-Figure7-1.png", "caption": "Fig 7. Laser deposited Inco 718 sample 9a) and ground specimen for tensile test (b)", "texts": [ "59 Table 1 show the detailed results of the composition of the Inco 718 plate, the Inco 718 powder, two different vanes (vane 1 and vane 2) and the intersection between the back wall and the vanes. The composition of laser deposited samples basically corresponds to the inconel 718 powder and two vanes have the basically identical composition, which means that laser deposition demonstrate good metallurgical and compositional homogeneity. For tensile test, three wall-shape samples were deposited with dimension of 700mm*30mm*5mm (Fig.7). The deposition parameters are: laser power of 1500W, beam diameter of 5mm, scanning speed of 0.2m/min, powder feeding rate of 6.8g/min and layer thickness of 0.25mm. The laser deposited samples were first cut by EDM to the following shape and then ground to the final tensile specimens. The ground specimens were dense Proc. of SPIE Vol. 5629 63 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 05/16/2015 Terms of Use: http://spiedl.org/terms 6 and bright alloy ones free of porosities and cracks (Fig 3)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001695_1-4020-3169-6_36-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001695_1-4020-3169-6_36-Figure2-1.png", "caption": "Fig. 2. A static FEM computation at no load for a q = 0.3636 motor of a) surface magnet motor and b) embedded magnet motor.", "texts": [ " The FEM computations were carried out for different rotor structures and the results are shown in Table I. The results given for the surface magnet motors show that the 28-pole machine generates the highest torque and the 26-pole machine the lowest. The difference between the machines is anyway small. To compare embedded magnet motors with surface magnet motors some analytical calculations of the parameters were done. One of the motors studied was q = 0.3636, 24 slots and 22 poles. The results of static FEM computations at no load situation are shown in Fig. 2. On the left side is the motor designed with 22 surface magnets and on the right side with 22 embedded magnets. (There are 12 flux lines going through each magnet in both pictures.) For the surface magnet motor the fundamental value (from the Fourier spectrum) of the flux density normal component was 1.01 T and for the embedded magnet motor 1.17 T. The r.m.s values were 0.738 T and 0.92 T. With the same amount of magnet material \u2013 10.3 kg \u2013 the embedded magnet solution gives clearly higher flux density values at no load" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000390_05698190500313478-Figure12-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000390_05698190500313478-Figure12-1.png", "caption": "Fig. 12\u2014Project area Dxi and Dxa.", "texts": [ " [3], the following equation is obtained: dmbx = \u00b5p (d/2) \u221a (d/2)2 \u2212 x2\u221a (d/2)2 \u2212 x2 \u2212 z2 dx dz [5] By subtracting the frictional moment 2 \u222b\u222b Dxi dmbx +2 \u222b \u222b Dxa dmbx(in the case that only the exposure areas are covered with polymer lubricant) from the frictional moment 2 \u222b\u222b Dx dmbx(in the case that the entire surface of the ball is covered with polymer lubricant), the moment mbx (caused by the friction between one ball and the polymer lubricant in the x axis) is written as mbx = 2 \u222b \u222b Dx dmbx \u2212 2 (\u222b \u222b Dxi dmbx + \u222b \u222b Dxa dmbx ) [6] where Dx is the projection of the ball surface on the xz plane and is given by Dx = {(x, z) | \u2212d/2 \u2264 x \u2264 d/2, \u2212 \u221a (d/2)2 \u2212 x2 \u2264 z\u2264 \u221a (d/2)2 \u2212 x2} [7] In Eq. [6], Dxi and Dxa are the orthographic projections on the xz plane of the exposure areas at the inner and outer race sides, respectively (as shown in Fig. 12). For convenience in the analysis, the shape of the exposure area at the inner race side is simplified as a belt (length li and width si ), and that at the outer race side is simplified as a belt (length la and width sa). Under this simplification, Dxi and Dxa are written as Dxi = {(x, z) | \u2212 \u221a (d/2)2 \u2212 z2 \u2264 x \u2264 \u2212 \u221a d2 \u2212 l2 i /2, \u2212 si/2 \u2264 z \u2264 si/2} [8] Dxa = {(x, z) | \u221a d2 \u2212 l2 a/2 \u2264 x \u2264 \u221a (d/2)2 \u2212 z2, \u2212 sa/2 \u2264 z \u2264 sa/2} [9] Equations [6]-[9] show that mbx increases as Dxi and Dxa decrease and as \u00b5 and p increase" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000841_icsmc.2004.1401038-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000841_icsmc.2004.1401038-Figure1-1.png", "caption": "Figure 1. The sketch of the Tricept-like parallel manipulator", "texts": [ " In this paper, the kinematics of a Tricept-like parallel robot is detailed analyzed. Firstly, mobility analysis is performed to determine the number and characteristics of degrees of freedom. Then the inverse kinematics problems are described in closed forms. By calculating Jacobian matrix, dexterity analysis of the new parallel robot is performed. Thirdly the workspace analysis is studied. Finally the kinematic characteristics of the robot and the Tricept robot are compared. 2 Structure of the Tricept-like robot As shown in Fig.1, the proposed robot has a fixed platform and a moving platform connected by three legs. One leg connects the fixed platform and the moving platform by a universal joint and a prismatic joint respectively. The other two legs are UPS chains or SPS cha i i . The three prismatic joints in the robot are all actuated joints. It can he transformed from the Tricept 0-7803-8566-7/041$20.00 0 2004 ZEEE. 531 2 robot by taking off a leg and adding an actuator to the central leg. Considering the manipulator mobility, let F be the degrees of freedom, n the number of links, j the numbe~r of joints, f; the degrees of freedom associated with the i\" joint, and 1 = 6 , the motion parameter. Then, the degrees of freedom of a mechanism is generally govemed by the following mobility equation: For the manipulator shown in Fig.1. n = 7 , j = 8 . Applying equation ( I ) to the manipulator produces: F = 6 ( 7 - 8 - 1 ) + 2 ~ 3 + 3 ~ 1 + 3 ~ 2 = 3 . Hence, the manipulator is a 3-DOF mechanism. Due to the arrangement of the links and joints, the leg 3 constraints rotation about the z axis, and the manipulator has IWO translational degrees and one rotational degree as Tricept robot. 3 Inverse kinematics As shown in Fig.2, the reference frame 0 - x y z is attached to the fixed base at point 0, located at the center of the fixed platform" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001517_j.physb.2006.04.007-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001517_j.physb.2006.04.007-Figure1-1.png", "caption": "Fig. 1. Symmetrical accommodation loops traversing between x \u00bc70.7.", "texts": [ " In this section the model will be demonstrated on examples selected from the literature and used by some of the other known models. This gives us the opportunity to show the efficiency and speed of the model (see later) and we can compare the result to that of other models. Although an attempt was made to compare numerically the results from various models but the lack the availability of vital data in the literature prevented this move. In all these calculations right through the paper, for simplicity, normalized, dimensionless values will be used. With reference to Fig. 1, let us consider the case when, after cycling on the major loop, the field is stopped at xr \u00bc 1:2, (point A) reversed and a new cycling field applied oscillating between the peak values of 70.7. We can write, in general terms, the normalized set of equations for the nth f+ ascending and the f descending branches of the loops in the following form [9,10]: f \u00fen \u00bc f 1\u00f0ax a0\u00de \u00fe Cn, (1a) f n \u00bc f 1\u00f0ax\u00fe a0\u00de Cn, (1b) where Cn \u00bc f 1\u00f0axmn \u00fe a0\u00de f 1\u00f0axmn a0\u00de =2. (2) Here f1 represents the hyperbolic tangent transformation, a0 is the normalized coercivity, xmn is the nth maximum normalized field amplitude and a is a constant shape factor" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000882_acc.2004.1384686-FigureI-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000882_acc.2004.1384686-FigureI-1.png", "caption": "Fig. I . Observer-identifier-eonvoller system diagram,", "texts": [], "surrounding_texts": [ "Parameters used in friction model (2.2) are as follows: a, = 0.0 1, 0, = 0.1, F, = IO, F, = 20, v, = 0.1, a, = 0.02 for F(b1), F(8,) and u2 = 0.056 for F(8,), F ( 8 , ) .\nParameters used in deadzone model (2.3) are as follows:\nParameters used in backlash model (2.4) are as follows:\nExtemal disturbances that appear in (2.5) are as follows:\nd. = -0.1, d+ = 0.1, A , ( u ) = (U -d . ) , &(U) = (U -d+).\nd- =-0.1, d\u2019 =O.l, m = 1.\nd,, = dO4 =[0.0lsin(8,4), arctan(8,8,)lr., dn2 = dn3 = [0.01randn(2,1)]r,\nwhere randn represents white noise. To simulate payload changes, all elements in the matrices M,V above are multiplied by a factor of 3 during simulation periods from 1 to 6 s and 11 to 16 s.\nIt can be verified that this system satisfies all required assumptions of the proposed control scheme. Only the output, y =[S, , 4lr, is measured and all nonlinear functions are assumed unknown. The control objective is to guarantee that (i) all closed-loop signals remain bounded, and (ii) the output y follow the desired trajectory generated by passing a square wave of amplitude IO and 5, for 8, and 8, respectively, with zero mean, and 20-s period into the filter l/(s + 2)\u2019 . Number of hidden-layer nodes is 3. The observer and controller are as in Section 111. All design parameters are as follows:\nc i = 1 5 , u ufi =a 6 - u WE, . = U vgt .=a ki - 0 1 \u2019 9 V i = 2 4 3 ,\n~ = 0 . 1 , Lj=[16,91,216,180]T, Vj= l ,Z .\nSampling period is 1 ms. Saturation limit of control inputs is set at *50 Nm . All initial values are set to 0.1. Simulation results are as shown in Fig. 6 to 8. We can see from Fig. 6 and 7 that the controller achieves good tracking performance of both S, and 8, even if payload is changed during the two time intervals from 1 to 6 s and 11 to 16 s. This results from good observer and controller performance as seen in Figures 7 and 8, parts (c) and (e), respectively. Actual control inputs are bounded as shown in Fig. 8.\nTo have a brief idea of how much each component in the control system contributes to the overall tracking performance, several plots are provided in Fig. 5. We compare four situations: a) when the observer is not used, b) when the identifier is not used, c) when both observer and identifier are not used, and d) when both observer and identifier are used. We can see that the tracking performance of this controller is comparable to that obtained in the ideal case.\nv. CONCLUSION\nAnalysis as well as simulation using a rather complete model of a two-link flexible-joint manipulator shows the\neffectiveness of the proposed backstepping neural network observer-controller design scheme. It should be noted that this design scheme can handle uncertainties very well both from additive disturbances and from unknown nonlinear fimctions, which can be time-varying or contain unmodeled dynamics.\nREFEKENCES\n[ I ] M. C. G d , L. M. Sweet and K. L. Stmbel, \u201cDynamic models for control system design of integrated robot and drive systems,\u201d Journal o f b n a m i c Svstem Measuremen1 & Control. vol. 107. no. 1, pp.53-9, March 1985. I21 F.L. Lewis, J. Campos and R. Selmic, Neuro-Furry Control of .. Industrial Sysrem -with A~tuafor Nonlinearilies. Philadelphia: SIAM, 2002. [3] B. Brogliato, R. Ortega and R. Lozano, \u201cGlobal tracking controllers far flexible-joint manipulators a comparative shrdy,\u201d Automatico, vol. 31, no. 7, bp. 941-956. July 1995. [4] S.S. Ge, T.H. Lee and C.J. Hams, Adoplive Neural Network Control o/ Robolic Manipulolors. Singapore: World Scientific Publishing, 1998. [ 5 ] J. Hernandez and J. P. Barbot, \u201cSliding observer-based feedback control for flexible joints manipulator,\u201d Aulomalico, vol. 32, no. 9, pp. 1243-1254, September 1996. [6] H. D. Taghirad and G. Bakhshi, \u201cCompositc-H- controller synthesis for flexible joint mbots,\u201d in Proc. 2002 IEEE/RsI Inrl. C m 1 on Intelligent Robots and System, Switzerland, 2002, pp. 2067-2072. C. W. Park and Y. W. Cho, \u201cAdaptive tracking control of flexible joint manipulator based on fupy model reference approach,\u201d in IEE Pro=. Control Theory and Applicofions, 2003, pp. 198-204. F. Abdollahi, H. A. Talebi and R. V. Patel, \u201cA stable neural network obseNer with application to flexible-joint manipulators,\u201d in Proe. pk Internotianal Con1 on Neural In/onnotion Processing, 2002, pp. 1910-1 914. [9] W. Chatlatanagulchai, and P. H. Meckl, \u201cRobust observer backstepping neural nchvork contml of nonlinear systems in strict feedback form,\u201d in Proc. Americon Control Conference. Boston, 2004, to be published [ IO] H. Nho, \u201cAn experimental and theoretical sudy of various control approaches to flexible-joint robot manipulator undergoing payload changes,\u2019\u2019 Motion and Vibration Control Lab. Report, School of Mechanical Engineering, Purdue Univenity, 2003. [ I l l M. W. Spong, \u201dModeling and control of clastic joint robots,\u201d Journol qf Dynamic Sys tem Measurement & Control-Tronsacrians ofrheASME, vol. 109, no. 4,pp. 310-319, December 1987. [ I21 G. Tao and P. V. Kokotovic, Adaplive Control o / S y ~ l e m with Actuator and Semor Nonlineorities. New York Wiley, 1996. [ I31 C. Canudas, H. Olsson, K. J. Astmm and P. Lischinsky, \u201cA new model for contml of systems with friction,\u201d IEEE Trans. Automat. C o n k , vol. 40, no. 3, pp. 419-425, March 1995. [ I41 K. Homik, M. Stinchcombe and H. White, \u201cMultilayer feedforward networks are universal approximaton,\u201d Neural Networkr, vol. 2, pp. 359-366. [ I S ] S . S. Ge, C. C. Hang, T. H. Lee and T. Zhang, Stable Adaptive Neurol Network Conrrol. The Netherlands: Kluwer, 2002, ch. 2. [7] [E]", "are not used (measured states, known plant). (d) when both observer and identifier are used (unmeasured states, unknown plant)." ] }, { "image_filename": "designv11_61_0001836_978-3-540-71967-0_2-Figure2.6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001836_978-3-540-71967-0_2-Figure2.6-1.png", "caption": "Fig. 2.6. \u2018Morphious\u2019, the Virginia Tech morphing wing wind tunnel simulator: a cruise configuration, b attack configuration, c wing twist (Photos courtesy of the designer David A. Neal, III)", "texts": [ " and in camber, twist, and asymmetric planform changes for flight control motivated by predator birds such as a hawk [13]. This bio-inspired direction for morphing aircraft structures has lead to numerous research projects span- ning flight dynamics, aerodynamics, structural mechanics, and control. The most common motivating example is the desire to have an unmanned aircraft that can morph from a long aspect ratio, straight winged plane for efficient loitering flight into a highly maneuverable short, swept wing aircraft that is effective in attack (Fig. 2.6). The second common example is the design of high altitude long endurance (HALE) aircraft that can take off and land on their own. Extremely long, highly flexible wingspans are required for long endurance and such wings tend to hit the ground during take off and landing. A morphing solution would be to fold or otherwise morph such wings into shapes more favorable for take off and landing. Structural health monitoring (SHM ), condition-based maintenance (CBM ) and birth-to-retirement refer to the capability of using sensors throughout the life or an adaptronic structure to monitor its state of health and act ac- cordingly" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003046_1.3650515-Figure13-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003046_1.3650515-Figure13-1.png", "caption": "Fig. 13 Experimental equipment used by Litton Systems", "texts": [ " Load and speed were varied during the tests, and the minimum gap was measured. 3. Results. The test results are shown in graphical form in Fig. 12. The temperature of the bearing varied from 75 F to 100 F during the tests, resulting in a 3 percent change in gas viscosity. This probably had little effect on the results obtained if no other, such as distortion of the bearing surfaces, occurred. B. Litton Systems, Inc. 1 Description of Test Equipment. The Litton thrust bearing test fixture, shown in the diagram of Fig. 13, consisted of a synchronous motor-driven rotor, restrained from radial motion by an externally pressurized gas journal bearing and loaded axially by a gas-pressurized plenum beneath the rotor. The experimental thrust bearing was mounted on a gimballed pad which positioned the bearing parallel to the rotor face, separated by the gas film. Gas film thickness was measured by an electromagnetic pickoff with a maximum error of 8 microin. over the operating range. The ranges of the parameters which can be simulated by the fixture are tabulated below: As constructed, the rotor was a closed cylindrical sleeve with a flat face and functioned as the rotating member of the thrust bearing system" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000709_j.jsv.2003.12.016-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000709_j.jsv.2003.12.016-Figure1-1.png", "caption": "Fig. 1. Simplified model of launcher and finite element mesh.", "texts": [ " Vibration of the launcher is accompanied by fluctuations in the liquid pressures, especially at the fuel feed inlets in the propulsion unit, which induces pressure and feed rate oscillations, and consequently, fluctuations in the rocket engine thrust, together with oscillatory forces at the anchor points of structure lines of the launcher. These forces can develop to increase the launcher vibrations, and hence lead to an instability known as the \u2018\u2018Pogo\u2019\u2019 effect. This study focuses on the fluid\u2013structure interaction by using the simplified model for partially-filled cylindrical shells under a thrust as shown in Fig. 1. In this, finite element meshes consist of cylindrical shell elements, fluid solid elements, and interface elements. To obtain a symmetric finite element formulation, we describe the fluid by pressure \u00f0P\u00de and displacement potential \u00f0j\u00de; respectively. The motion of an isothermal, inviscid fluid with small disturbance is governed by Continuity equation: \u2019rF \u00fe rF vi;i \u00bc 0 in RF ; \u00f01\u00de Euler equation: rF \u2019vi \u00bc tij;j in RF ; \u00f02\u00de Barotropic equation: \u2019P \u00bc B rF \u2019rF \u00bc 0 in RF ; \u00f03\u00de where rF is the fluid density, vi is the velocity component, and tij is the Cauchy stress tensor given by tij \u00bc Pdij; where P is the pressure, dij is the Kronecker Delta and B is the bulk modulus", " If lf is near lp \u00fe lq and fpq fqp > 0; lf \u00bc lp \u00fe lq7 e1 2 Lpq \u00fe O\u00f0e21\u00de; Lpq \u00bc fpq fqp lplp 1=2 \u00f032\u00de and if lf is near lp lq; fpq fqpo0; and lp > lq; lf \u00bc lp lq7 e1 2 Lpq \u00fe O\u00f0e21\u00de; Lpq \u00bc fpq fqp lplp 1=2 : \u00f033\u00de From Eqs. (32) and (33), it is easily seen that the sum-type and difference-type combination resonances cannot exist simultaneously for any pair of natural frequencies lp and lq: In this study, we used the finite element method to numerically test a simplified model of a liquid-filled structure, as shown in Fig. 1(a). And as shown in Fig. 1(b), the model of the finite element meshes is composed of the shell structure elements, the 8-node interface elements, and the interior liquid elements. And the liquid element consists of the uniform three-dimensional 20-node cubic and 15-node tetrahedron elements. To check the validity of the results, we compared numerical data with those of previous work. Table 1 summarizes the material and geometric data for the cantilevered cylindrical shell model. Table 2 shows the numerical results of the present study and the theoretical and experimental results [21,22]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002498_icems.2009.5382880-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002498_icems.2009.5382880-Figure6-1.png", "caption": "Fig. 6. Comparisons with the effective value, the average value, the electromagnet current waveform, and duty factor of the proposed motor and Minato motor.", "texts": [ " 5, the analysis of the action of Minato motor is shown. In the proposed motor, \u201cstrong repulsive force\u201d is not used and \u201cweak repulsive force\u201d and \u201cweak suction force\u201d are used. Therefore, the permanent magnet can not be demagnetized. For not demagnetizing the permanent magnet, the permanent magnet, of which pole face is wide, can be used. A magnetic flux quantity of the permanent magnet is large, because the pole face is wide. Therefore, generated electromagnetic torque is big, the on-period is big, and, d is also big in the proposed motor. Figure 6 shows comparisons with the effective value, the average value, the electromagnet current waveform, and duty factor of the proposed motor and Minato motor. In Fig. 6.a, the electromagnet current waveform, the effective value, the average value, and duty factor of Minato motor are shown and in Fig. 6.b, the proposed In the analysis of the copper loss of Minato motor, an angle of on-period in Minato motor is a and an angle of on-period in the proposed motor is na. The current average values of the proposed motor and Minato motor are same. The current effective value of the proposed motor compared to the current effective value of Minato motor is n/1 . Therefore, the copper loss of the proposed motor is 1/n of the loss of Minato motor, when both coil resistances of the electromagnet are same. The copper loss of the proposed motor becomes 20% of the loss of Minato motor, if 5=n , n is the ratio of on-period. In Fig. 6, \u03b8 is a phase angle and t\u03d6\u03b8 = . In the general DC brushless motors, an alternating current flows. In the general DC brushless motors, armature current, i, consists of quadrate-axis current, iq, and direct-axis current, id. The source voltage is increased in order to increase iq. The equivalent weak field control is carried out, when source voltage becomes biggest and the source voltage can not be increased. The equivalent weak field control is carried out in order to increase iq. In other words, iq is increased by making id to be a negative value" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003954_iros.2010.5650000-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003954_iros.2010.5650000-Figure3-1.png", "caption": "Fig. 3. Definition of sub-contacts: (a) the vertex-triangle contact and (b) the edge-edge contact.", "texts": [ " The e1 consist of two vertices of v1 and v2, and the four edges of e1, e2, e3, and e4 are arranged in counter-clockwise order on the normal vector (u1) to make the f1. The adjacent relation between elements implies the connection by an edge in case of two vertices, by a vertex in case of two edges, and by an edge in case of two faces. The adjacent relations are used to generate another contact states. In our framework, all faces of the polyhedral objects are triangulated to employ the proposed sub-contacts (SCs). The sub-contacts are defined as a vertex-triangle (v-t or t-v) contact and an edge-edge (e-e) contact between two objects in Fig. 3. The sub-contacts are denoted as (vi, tj) or (ti, vj) for the vertex-triangle contact and (ei, ej) for the edge-edge contact. Although the contacts of v-v, v-e, e-t and t-t can be presented theoretically, we do not regard them as the sub-contacts. It is because the v-v contact and the v-e contact are impossible in robotic operation and the e-t contact and t-t contact are multi-contact situations which can be described by two v-t contacts and three v-t contacts, respectively. B. Triangulation of Faces We propose a hierarchical structure of vertices and edges to triangulate faces" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003026_icelmach.2008.4799923-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003026_icelmach.2008.4799923-Figure4-1.png", "caption": "Fig. 4. Voltage and current phasor diagram for the winding 1 (a) using the complex phasors of current and voltage fundamental waveforms (b).", "texts": [ " A difficulty in the practical use of (13)-(14) is that the voltages vab1 and vab2 are expressed by (4)-(5) as functions of the time t from the beginning of the simultaneous commutation, while the parameters L11, L22, L12 are expressed by (15) as functions of the rotor position \u03b8. It is then necessary to express \u03b8 too as a function of time t, i.e. in the form: tt \u03c9\u03b8\u03b8 += 0)( (19) \u03b80 being the rotor position at the beginning of the simultaneous commutation event. For this purpose, the current and voltage phasor diagram of a motor winding (e. g. winding 1) can be considered as shown in Fig. 4a. The complex quantities 1abV , 1aV , 1bV , 1cV , 1aI , 1aE are the phasors associated to the fundamental components of vab1, va1, vb1, vc1, ia1, ea1 respectively (as exemplified for the current ia1 in Fig. 4b); ea1 denotes the inner e.m.f. due to the rotor excitation current only (open-circuit e.m.f.); the angle \u03b41 is the so called load angle of the motor. The same angles appearing in Fig. 4a are reported in Fig. 5 as phase shifts between the various waveforms obtained from a dual threephase LCI drive simulation [1], [4], [6]. From both Fig. 4 and Fig. 5 it is clear that the phase shift between cos(\u03b8) wave, corresponding to the d-axis phasor, and the fundamental of vab1 is \u03c0/2+\u03c0/6+\u03b41=2\u03c0/3+\u03b41. It is then possible to write vab1 as: ( )[ ] ( )[ ] ( )[ ]1 1 11 6/7)(sin2 2/3/2)(sin2 3/2)(cos2)( \u03b4\u03c0\u03b8 \u03c0\u03b4\u03c0\u03b8 \u03b4\u03c0\u03b8 ++= +++= ++= tV tV tVtv n n nab (20) Equalling (20) to (4) yields at last: 16/7)( \u03b4\u03c0\u03b1\u03c9\u03b8 \u2212\u2212+= tt , (21) i.e. the sought rotor position \u03b80 at the beginning of the simultaneous commutation is: 10 6/7 \u03b4\u03c0\u03b1\u03b8 \u2212\u2212= . (22) The load angle \u03b41 can be easily computed for any motor operating condition with standard methods, e.g. through phasor diagrams as shown in Fig. 4a. At this point, all the necessary elements are available for the numeric evaluation of (13)-(14), since vab1(t) and vab2(t) are given by (4)-(5) and coefficients L11, L22, L12 can be evaluated through (15) and (21) as functions of the time t as follows: ( ) ( ) ( )[ ]111 2cos)( \u03b4\u03b1\u03c9 \u2212+\u2032\u2032\u2212\u2032\u2032+\u2032\u2032+\u2032\u2032= tLLLLtL qdqd ( ) ( ) ( )[ ]\u03c4\u03b4\u03b1\u03c9 \u2212\u2212+\u2032\u2032\u2212\u2032\u2032+\u2032\u2032+\u2032\u2032= 122 2cos)( tLLLLtL qdqd ( ) ( ) ( ) ( )[ ] \u03c3\u03c4\u03b4\u03b1\u03c9\u03c4 LtLLLLtL qdqd \u0394+\u2212\u2212+\u2032\u2032\u2212\u2032\u2032+\u2032\u2032+\u2032\u2032= 2/2coscos)( 112 (23) Equations (13)-(14) are suitable for being represented in a circuital form as depicted in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000261_bjo.8.1.33-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000261_bjo.8.1.33-Figure1-1.png", "caption": "Fig. 1. (a) Standard appliance. (b) Screening appliance. (c) Reverse (Class Ill) appliance (after Baiters).", "texts": [ " It results in the instant enlargement of the enoral space, bringing the tongue forward in Class II cases and back in Class III. The acrylic bulk of the appliance is minimized to permit day-long wear and to maximize oral space. Incisor eruption is prevented, but an open bite is produced between the cheek teeth which will thus be free to erupt and achieve permanent bite opening and lowering of the mandible. The vestibular wire loops prevent intrusion of the cheeks into the interocclusal space. Appliance Types There are three Bionator variants (Fig. 1): I. The Standard Appliance for the treatment of Class II, Division I. 2. The Screening Appliance for the elimination of abnormal tongue activity in open bite cases. 3. The Reverse Appliance for the treatment of Class III. H. L. Eirew 1. The Standard Type (a) Construction Working bite If edge to edge incisal contact is possible, this should be registered by the working bite. With an exceptionally severe overjet, a more relaxed working position is used, based on correct canine relationships. In these cases, the lower incisors should be capped to prevent their further eruption and tilting" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001618_robot.2005.1570438-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001618_robot.2005.1570438-Figure4-1.png", "caption": "Fig. 4. Simplified analytical model of a two-link planar casting manipulator. The gripper is replaced by a mass point since we are not considering the orientation of the robot\u2019s end-effector.", "texts": [ " During this initial swinging phase, the gripper state is related to the instantaneous robot configuration by the following relation:\u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 x0 = xbase + a1 s1 + a2 s12 , y0 = ybase \u2212 a1 c1 \u2212 a2 c12 , \u03d5 = q1 + q2 + \u03c0 2 , x\u03070 = a1 c1 q\u03071 + a2 c12 (q\u03071 + q\u03072) , y\u03070 = a1 s1 q\u03071 + a2 s12 (q\u03071 + q\u03072) , \u03d5\u0307 = q\u03071 + q\u03072 , (1) Therefore, the gripper state at the throwing time is obtained by (1) evaluating the joint variables, q1 and q2, and their first temporal derivatives at the same time. By doing so, we assess also the gripper state at the beginning of its flight. According to our intention to focus in the first place only on reaching a target object in motion, the robot\u2019s end-tip orientation \u03d5 may not be considered. From now on, refer to figure 4 in which the gripper is replaced with a mass point. Similarly to what was obtained above, the mass point state at the throwing time is related to the manipulator state and is given by the relation\u23a7\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23a9 x0 = xbase + a1 s1 + a2 s12 , y0 = ybase \u2212 a1 c1 \u2212 a2 c12 , x\u03070 = a1 c1 q\u03071 + a2 c12 (q\u03071 + q\u03072) , y\u03070 = a1 s1 q\u03071 + a2 s12 (q\u03071 + q\u03072) . Once the mass point has attained enough energy, it is thrown towards the target object by releasing the string. Without any control action, the mass point goes through an arc of parabola subjected only to the gravity acceleration g" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001559_095440505x32814-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001559_095440505x32814-Figure1-1.png", "caption": "Fig. 1 Split roll surface: A, pull bar; B, fixed part; C, movable part; D, needle bearing; E, bearing; F, condenser; G, workpiece", "texts": [ " For friction measurements, the pin needs some space for tangential deflections, necessitating a gap between the roll and the pin. During the rolling process, material and lubricant may be extruded into this gap, leading to inaccurate measurements. Finally, the lateral deflections of the pin may create contact between the pin and bore which on the output of the transducer would be misleading. To avoid the problems related to the pin design, Wanheim and Zeuthen [9] invented a new concept, where the roll surface is split into two parts. One part is supported by needle bearings and fastened to the other by screws (Fig. 1). By use of a condenser, the distance between the needle-supported segment and the fixed segment could be determined. The output of the condenser expressed the forces between the two segments. This load represented then the integral of the frictional stresses acting at the needle-supported segment. By differentiating the signal, the variation in the frictional stresses through the contact zone could be determined. However, an inconvenient combination of the normal pressure and the material flow in the deformation zone caused the material always to flow uphill against the step created between the two segments, disturbing the measurements" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003840_ssrr.2013.6719339-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003840_ssrr.2013.6719339-Figure4-1.png", "caption": "Fig . 4 . Expected robot paths that are computed using control inputs.", "texts": [ " XI = xl_1 + V t sin 81_1 YI = YI_I + V t cos 81_1 81 = \ufffd_I + 0) t ( 1 ) I f a collision will occur on the expected path, the IRRO fmds the best robot pose to prevent this collision with an object near the expected path. This involves searching for the robot pose in the small area of the expected path. The operator can control the robot on the expected path, but there is no need to do so. If there is a narrow passage that the robot cannot pass through, and the operator commands the robot to pass through this narrow passage, the IRRO sends an alarm to the operator. However, the robot cannot override the command of the operator. Fig. 4 shows the expected paths. The solid robot model represents the current robot pose, and the red circles show the expected path. In the environment shown in Fig. 4, the operator may not be convinced that the robot can pass through the narrow passage, because of ambiguities related to the robot ' s size and width and the height of the confmed region. Therefore, the collision detection process has to be performed for the expected paths in advance. C. Expected Collision Detection The collision detection is performed using a box model robot. If a 3D point of an object shown on the 3D map is inside these boxes, the IRRO detects a collision with the environment" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000915_bfb0015079-Figure1.1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000915_bfb0015079-Figure1.1-1.png", "caption": "Fig. 1.1. Modes of oscillation of a simple pendulum", "texts": [ " The question arises from the inherent instability of inverted pendulums in upright positions. It is well known that one can change the structural stability of inverted pendulum systems by applying torques at various joints. Consider a single degree of freedom model that consists of one mass and a torsional spring at the pivot point. Typical modes of motion include oscillations about two static equilibria and a third mode where the pendulum undergoes cyclic motions. These three modes are depicted in Fig. 1.1 and labeled as A, B and C respectively. If the ground surface is included, the third mode of behavior is ruled out since this leads to the collapse of the penduhm. When we add a second link to the simple system that we have considered and coordinate the motion of various members by applying appropriate joint moments the dynamic behavior remains similar to we have observed in the single member case [16]. In the presence of the walking surface, the system may still operate in either mode A or B, as long as all the parts of the system remain above the ground, Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002930_icelmach.2008.4799943-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002930_icelmach.2008.4799943-Figure4-1.png", "caption": "Fig. 4. Different rotor topologies with interior magnets in the rotor", "texts": [ " ANALYSIS OF THE SALIENT POLE SYNCHRONOUS MACHINES WITH NON-UNIFORM AIR-GAP LENGTH In the previous section, during analysis of salient pole synchronous machines it is assumed that the air-gap length along the circumferential direction may be approximated with a sinusoidal function. However, this is valid only for an ideal optimised machine and which operates also under linear operation condition. Otherwise for the PM machines with inset or buried magnets in the rotor core the effective air-gap length depends on the magnet slot shape and also on the saturation condition of the machine. The following figure 4 shows two different rotor topologies with embedded magnets in the rotor core. For these machine types the variation of the air-gap length along the circumferential direction deviates far from the sinusoidal function if the reluctance component isn't good optimised or if the machine operates under high load condition (saturation effect). The variation of the effective air-gap permeance for a real salient pole machine is illustrated in the following figure 5. Referring to the figure 5 the variation of the effective air-gap length for the salient pole machines with non-uniform sinusoidal air-gap length along the circumferential direction can be described using the following function, g(O) =_--,,\"'," ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003095_s0022112070001490-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003095_s0022112070001490-Figure6-1.png", "caption": "FIGURE 6. Pressure coefficient for a subsonic wing ( M = 0.8, M I = 0.25, k = 3, y = 0, ct = 1) .", "texts": [ " In particular, along the intersection of the wing with the shock, x = MCt, Diffraction of shock waves by a moving thin wing 603 the disturbance pressure is constant outside the Mach cone of the equivalent wing and is the value along a ray from the vertex T' of the conical solution. At T' the pressure is not single valued. It ranges from the two-dimensional value behind the oblique shock attached to the leading edge and the conical values along the rays from the vertex T' to zero ahead of the leading edge. In figure 6, the wing is moving at subsonic speed (M, < 1, Bo < 1) . Outside the region G, the disturbance pressure is the subsonic steady solutionp*(x*, y, z ) corresponding to a wing moving at velocity (1 - x*M) C/(X* - M ) = U* in x-t variables. The pressure distribution behind the shock for t < 0 can be obtained from the present result by a translation of x co-ordinate, e.g. the pressure distribution at the instant to > 0 at x = - 2Ct0 is the same as that at x = - (2C - U*) to at the instant 1 = 0. (iii) Applications For a thin symmetric wing with an arbitrary planform and thickness distribution, the pressure disturbance behind the shock wave can be obtained directly from (4" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000717_kem.297-300.102-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000717_kem.297-300.102-Figure2-1.png", "caption": "Fig. 2 SED and CED of a uniaxial tension specimen and an equibiaxial tension specimen", "texts": [ " \u222b= = = \u03b5 \u03b5\u03c3 \u03b5\u03c3 \u03b5\u03c3 0 rr rr rr dW ddW rdrdW T (1) Here, \u03c3 is the stress tensor, \u03b5d is the strain increment tensor, \u03c3 r is the stress vector, \u03b5 r d is the strain increment vector, and r r is the normal vector. Integrating the CED increment during a loading history, CED can be obtained for a (virtual) crack at any location in a structure. It is mostly likely for a crack to initiate along the plane on which CED has its maximum [5]. SED and CED were calculated for a uniaxial tension specimen and for an equibiaxial tension specimen used for the fatigue tests [8, 9], and they are shown in Fig. 2. Note that CED for an equibiaxial tension specimen is a half of that for a uniaxial tension specimen even though SED is the same. Thus, it can be said that CED is in a better correlation than SED with the fatigue test data, which showed about four times longer fatigue life of an equibiaxial tension specimen. It is well known that J-integral is the same as SERR, but it can be calculated for a crack in a two dimensional structure by using elements which satisfy 2/1/1 r singularity. In order to calculate SERR for a crack in a three dimensional structure such as a tire, VCCT can be used" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003935_icems.2011.6073730-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003935_icems.2011.6073730-Figure8-1.png", "caption": "Fig. 8 Distribution of the flux line in an induction motor", "texts": [ " Obviously, the SVPWM scheme has less eddy current losses. All of these results indicate the SVPWM scheme has less harmonic components and eddy current losses and that is a tremendous advantage compared to the SPWM scheme. In this section, taking into account the nonlinear magnetic property of the iron core and in order to make the result more accurate, the time-stepping FEA method is used. An induction motor of 11kW driven by the PWM power supply is analyzed by the 2D nonlinear time-stepping analysis. Fig.8 shows results of the flux distribution at the time of 0.15s, and Fig.9 shows the curve of the eddy current losses varying with timevariation at no load condition. The following part will show the influences of carrier wave ratio and modulation index on eddy current losses. A. Influences of carrier wave ratio on eddy current losses under SVPWM supply Under the SVPWM supply, set the modulation index at a constant value M=0.9 and the fundamental frequency at f=50Hz, then change the switching frequency from 2kHz to 4kHz, a series of data are obtained in TABLE I" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002546_cjoc.200990207-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002546_cjoc.200990207-Figure1-1.png", "caption": "Figure 1 Arrangement of optical bench.", "texts": [ " In this photochemical laboratory a unique optical processor was constructed to meet the requirement of photochemical reaction.19 An optical processor consists of the following parts: (i) light source (a high pressure mercury vapour lamp), (ii) optical system for steady flow of radiation source, (iii) monochromatic system, (iv) double walled reaction cell, (v) magnetic stirring system, (vi) temperature controlling system, (vii) photo detectors and signal recording system, and (viii) Uddin\u2019s19 system for inert atmosphere (Figure 1) according to following scheme.5 All these systems were arranged with an optical processor which was fit to obtain the parallel beam of light as shown in Figure 1. Light from the mercury vapour lamp radiation source was made to pass through convex lenses L1, L2 and L3. Between lenses L2 and L3 light was passed through a hole (O) of 2 mm diameter. A red Kodak filter was used to obtain monochromatic wavelength. A plain reflector (R) was used for reflection of fraction of light to a reference photocell (P2). Remaining fraction of light was passed through reaction cell (C). A magnetic stirrer (M) was used to homogenize the reaction mixture. The photocell (P1) was used to detect the light intensity" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002953_ijsurfse.2009.026607-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002953_ijsurfse.2009.026607-Figure1-1.png", "caption": "Figure 1 Single flank gear-testing machine", "texts": [ " Our measurement policy is \u2022 with ultra-high precision \u2022 under high applied load \u2022 with high accuracy so as to observe not only peak-to-valley values but also the variation during one pitch period in detail. If it is achieved, we can identify not only manufacturing error, but also the influence of tooth deflection in the range of sub-micrometre order. As a result, it must also be possible to detect slight amplitude modulation in TE. However, since the measurement is limited under only quasi-static conditions, we have to give up the evaluation of high-speed transition phenomena as a reward of high precision. The instrumentation of the developed measuring system is shown in Figure 1. The measuring system consists of a back-to-back type gear-testing machine (centre distance: 156.2 mm) to investigate under high applied load of actual running conditions (up to 49 kN in the tangential force). A pair of optical rotary encoders is installed to the end of each shaft and they are connected through interpolating units to a personal computer. Sinusoidal waves from each rotary encoder are divided electrically and transformed into a series of pulses. The counter board detects the rotational angle of each shaft by counting those output pulses" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003678_tmag.2012.2201254-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003678_tmag.2012.2201254-Figure2-1.png", "caption": "Fig. 2. Mesh with antiperiodic boundary condition.", "texts": [ " The merit of this method is that on the two edges of the period boundaries, the local coordinates of the nodes do not need to be the same. That means, when the mesh is generated, it is not necessary to consider the periodic boundary conditions. This simplifies the process of mesh generation with acceptable accuracy. A mesh with such antiperiodic boundary conditions, from the Team Workshops Problem No. 30 [11] and which will be used to verify the proposed method as described in Section B of the following paragraphs, is given in Fig. 2. A simple example as shown in Fig. 3 is used to study the numerical errors arising from rotor rotation. The stator and rotor are made of iron and their conductivity is set to zero. There is a 100-turn coil carrying a dc current of 100 A in the airgap. Firstly a coarse mesh with 903 triangles is used (Fig. 4(a)). In this study the rotor rotates at a speed of 360 degrees/s. The time step size is 0.001 s. For the first test, there are 90 master nodes and 90 slave nodes on the sliding interface. Interpolationmethod is used to deal with the matching boundary condition [7]", " The same numerical experiment is applied to a refined mesh with 2626 elements as shown in Fig. 4(b). The comparison of numerical errors at each time step is shown in Fig. 6. The average error of all time steps using the traditional method is %, while that of using the proposed method is %. Again there is a significant reduction in numerical error using the proposed algorithm. The Team Workshops Problem No. 30 [11], which is a threephase solid-rotor induction motor (IM), is also taken as an illustrating example to verify the validity of the proposed method. With themesh shown in Fig. 2, the calculated phase voltage with the proposed method and the analytical result are compared in Fig. 7. It is shown that the differences between the analytical results and the proposed FEM method are acceptably small. Relative movement of objects is one of the main sources of numerical error in numerical modeling methods. The proposed formula for rotor position prediction has a third-order accuracy. To reduce the discontinuity of magnetic fields on moving interfaces, the number of slave nodes should be several times of that of the master nodes" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002928_icinfa.2009.5205046-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002928_icinfa.2009.5205046-Figure2-1.png", "caption": "Fig. 2. Mobile robot platform and elements.", "texts": [ " New sensor combinations should be used. Different supervision and control models should equally be used to carry out the mobile robot tasks. This paper presents the virtual environment implementation for project simulation and conception of supervision and control systems for mobile robots and focus on the study of the mobile robot platform, with differential driving wheels mounted on the same axis and a free castor front wheel, whose prototype used to validate the proposal system is depicted in Fig. 1 and Fig. 2 illustrate the elements of the platform. Suppose that the robot is at some position (x, y) and \u201cfacing\u201d along a line making an angle \u03b8 with the x axis (Fig. 3). Through manipulation of the control parameters ve and vd, the robot can be made to move to different poses. Determining the pose that is reachable given the control parameters is know as the forward kinematics problem for the robot. Because ve and vd and hence R and \u03c9 are functions of time, is straightforward to show (Fig. 3) that, if the robot has pose (x, y, \u03b8) at some time t, and if the left and right wheels have ground-contact velocities ve and vd 978-1-4244-3608-8/09/$25" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001645_icar.2005.1507416-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001645_icar.2005.1507416-Figure1-1.png", "caption": "Figure 1: General Tripedal Rotopod", "texts": [ " However, the mechanism has the potential to do continuous rotation, producing energyefficient legged locomotion by rolling in a manner similar to a single wheel device such as the Gyrover [14]. T 2230-7803-9177-2/05/$20.00/\u00a92005 IEEE A general tripedal rotopod mechanism consists of three legs arranged in a tripod configuration. Each leg is equipped with a single translational degree of freedom, so that the leg length for each leg can be varied between \u2113 and \u2113e = \u2113 + \u03b4. A rotating joint is placed at the apex of the tripod. The joint rotates an armature of length r at the end of which is placed a reaction mass mr. This general tripedal rotopod is illustrated in Fig. 1. When all three legs are extended at length \u2113e the plane of rotation of the apex joint is parallel to the ground plane. If one of the legs is retracted to length \u2113, then the plane of rotation of the apex joint is tilted. It tilts around the line joining the endpoints of the legs opposite the retracted leg. If the angle each leg makes with the vertical centerline is \u03b2, then the ground distance between the projection of the apex and a leg endpoint is \u2113eSin\u03b2. If all three legs are positioned equally around the apex, then the ground distance from a leg endpoint to the midpoint of the line joining opposite leg endpoints is 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000898_cira.2003.1222115-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000898_cira.2003.1222115-Figure7-1.png", "caption": "Figure 7 Forward kinematics from all links", "texts": [ " However, as a map we must choose one of them only and reject the other, basing on some other considerations such as the value of angle a, , or previous d u e of (x, y) in a dynamics environment. The map f23(azia3) considers the angles a2 and a3 only and completely ignores al . Similar procedures can be applied to links 1 and 2 as well as links 1 and 3. Therefore, two more forward kinematics maps flz(a1,az) and fia(a,,aa) can also be obtained. 5.2 Forward Kinematics Map f m ( a l , a2, a3) Another forward kinematics map can be obtained by considering all 3 links together as in Figure 7. The end-effector coordinates (x,y) is the common intersection of the three circles represented by the following equation set (X - Zb1)\u2019 f (Y Ybl)\u2019 = T1\u2019 (15) (5 - Xb2)\u2019 + (Y - Yb2)* = ?\u2018Z2 (16) (X - Zb3)\u2019 + (Y - Yb3)\u2019 73\u2019. (17) As before X M and yb, are defined as: Subtracting (16) and (17) from (15) yields the two equations of intersection straight lines between circle 1 and 2, and, circle 1 and 3: -2(Xb1 X b 1 ) X - 2(Ybl - Y b 2 ) Y (19) + ( l r b l I z - r 1 2 ) - ( ( 1 r b 2 1 2 - - T 2 2 ) = o -2(Xb1 - X b 3 ) Z - 2(Ybl - Yb3)Y ~ ( ~ ~ b 1 ~ z ~ ~ T 1 z ) ~ ( ( 1 ~ b 3 ~ 2 ~ ~ 3 2 ) ~ ~ lrb,lZ = xbiZ + ya,' for i = 1,2,3" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003779_epe.2013.6634393-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003779_epe.2013.6634393-Figure5-1.png", "caption": "Fig. 5: Vector plot of flux linkage", "texts": [ " 3, Output of flux linkage of the output windin wave rectifier, variation of flux l position is big by armature reac output from bidirectional variatio by armature reaction. Accordingl Fig. 6 shows the torque waveform a DCeRG system tput windings tional DCeRG f load characteristic at 3000 rpm. The efficiency garded. 100 % um output is 1396[W] and efficiency is reache fier, maximum output is 1170[W] and efficiency half-wave rectifier is bigger than full-wave rect gs in load current 5[A]. The flux linkage is unipo inkage is small. This case is because leakage flu tion like Fig. 5. In other words, full-wave rec n of flux linkage; on the other hand, variation of y, output is small. s and the torque ripple of half-wave rectifier is i is expressed by (1). (1) d to 88.4 %. On the is reached to 90.6%. ifier. Fig.4 shows the lar wave. In the fullx in the non-opposed tifier can extract the flux linkage is small ndicated as 150 %. (a) Half-wave recti Fig. 3: Load current characteristi Fig. 4: Flux linkage waveforms 0 200 400 600 800 1000 1200 1400 1600 0 5 10 O ut pu t[ W ], E xc ite d in pu t[ W ] M ec ha ni ca l i np ut [W ] Load current[A Output Mechanical input fier (b) Full-wave rec c maximum output characteristic 0 200 400 600 800 1000 1200 1400 1600 0 1 2 3 4 5O ut pu t[ W ], E xc ite d in pu t[ W ] M ec ha ni ca l i np ut [W ] Load curren Output Mechanical input 0 20 40 60 80 100 15 20 E ff ic ie nc y[ % ] ] Excited input Efficiency tifier 0 20 40 60 80 100 6 7 8 E ff ic ie nc y[ % ] t[A] Excited input Efficiency In the preceding chapter, we d armature reaction" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000285_cdc.2003.1271669-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000285_cdc.2003.1271669-Figure2-1.png", "caption": "Fig. 2. B Propeller motion", "texts": [ " INTRODUCTION Devil stick is a kind of juggling such that human beings manipulate the floating stick (the center stick) with the other sticks (the hand stick). Various stick motions are performed by jugglers. For example, they hit the center stick by two hand sticks altematively(see Fig. 1). This trick is called \u201cidling\u201d motion, and Schaal et al. realized hitting the center stick several hundred times on average in experiments[ 11. The purpose of this paper is to achieve the trick called \u201cpropeller\u201d motion. This motion is to rotate the center stick continuously by only pushing with one hand stick (see Fig. 2). To achieve this trick, we have to control a position and an angular velocity of the center stick. Though this is a difficult control problem because we can\u2019t pull the hand stick, it is a quite interesting one. This paper is organized as follows. In $U, we consider a model of the devil stick and make some assumptions, and derive a state equation of the devil stick. Control objective to rotate the center stick continuously is stated in $111, and based on it we try to control by output zeroing in SIV" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002602_s00707-007-0564-3-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002602_s00707-007-0564-3-Figure6-1.png", "caption": "Fig. 6. A model with four DOFs having C4V symmetry", "texts": [ " Therefore, the analysis of the system with six DOFs is reduced to four problems of single DOF: T \u00bc 1 1 1 1 1 1 0:5 0:5 1 1 0:5 0:5 1 1 1 1 1 1 0:5 0:5 1 1 0:5 0:5 2 6666664 3 7777775 : For k1 = 1,000, k2 = 750, m = 50 and P\u00f0t\u00de \u00bc f10; 0; 0; 0; 0; 0g sin ffiffiffi 5 p t; the calculations leads to: ~M \u00bc 300 0 0 0 0 300 0 0 0 0 150 0 0 0 0 150 2 664 3 775; ~K \u00bc 6000 0 0 0 0 2400 0 0 0 0 5250 0 0 0 0 9750 2 664 3 775; w \u00bc 0:0022 0:0004 0:0022 0:0011 2 664 3 775 sin ffiffiffi 5 p t) u \u00bc T:w) u \u00bc 0:0060 0:0023 0:0010 0:0007 0:0010 0:0023 2 6666664 3 7777775 sin ffiffiffi 5 p t: The results obtained from the group theoretic method are identical to those obtained by classic modal analysis. The symmetry group of C4V , is one of the specific groups which requires additional treatment. This group has eight symmetry operations as fe;C2;C4;C 1 4 ; 2rV ; 2rdg which can be identified for a structure with four DOFs, see Fig. 6. M \u00bc m 0 0 0 0 m 0 0 0 0 m 0 0 0 0 m 2 66664 3 77775; K \u00bc k1 \u00fe 2k2 k2 0 k2 k2 k1 \u00fe 2k2 k2 0 0 k2 k1 \u00fe 2k2 k2 k2 0 k2 k1 \u00fe 2k2 2 66664 3 77775: Here, the character table of the group C4V is presented and the calculations are performed by the formation of the matrix T, Table 8. Similar to the previous examples, and using the character table, the dimensions of the subspaces are calculated as n1 = 1, n2 = n3 = 0, n4 = 1, n5 = 2. Due to the existence of an irrep of dimension 2, there is a trivial subspace with two duplicate roots and there is no need to find both of them" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003829_j.jallcom.2010.07.142-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003829_j.jallcom.2010.07.142-Figure1-1.png", "caption": "Fig. 1. Illustration of the local rectangular coordinate system on specimens.", "texts": [ " he chemical composition of the alloy sheet was determined by X-ray fluoresence spectrometry (XRF). Pole figures were obtained using a PANalytical X\u2019Pert ro diffractometer using Cu radiation with a nickel filter in the diffracted beam. extural data were collected to a psi angle of 70\u25e6 and corrected for contributions rom the background and for defocusing. Data were then analyzed using a PANlytical X\u2019Pert texture software in order to produce complete pole figures for the rystallographic orientation type. The reference frame is illustrated in Fig. 1. The adius direction, tangent direction, and normal direction were denoted as RD, TD nd ND, respectively. Tensile tests were conducted at room temperature on as-deposited and sinered specimens. The sample size and shape in the tensile tests are shown in Fig. 2. pecimens were strained at a rate of approximately 0.02 s\u22121 with an extensometer lipped to the gauge length until failure. . Results and discussion .1. Microstructures A Ni\u201318Cr\u20130.6Al superalloy sheet strengthened with 0.4Y2O3 wt%) was deposited" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003117_ichr.2009.5379588-Figure12-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003117_ichr.2009.5379588-Figure12-1.png", "caption": "Fig. 12. Optimized trajectories for W t = 0.5 and W e = 0.5, when the goal is in a critic position for collisions because it is very close to the wall, G = (3.0, 2.0) . Analysis of two different approaching directions.", "texts": [ " We do not report the trajectories obtained for different priority on time and energy because they do not differ too much; in fact, for this specific position of the target, in the two cases, the optimal opening angle would not be significantly different and the value of energy would not make the difference in the computation of the fitness function. Finally we consider the configuration in which the goal is on the opposite side of the hinge of the door, G = (-3.0,2.0), and in which we set S = (-3.0, -2.0), Wt = 0.2 and We = 0.8. Note that the initial and target positions are symmetrical compared to the configuration in Fig. 12(b), but, because of the constraints imposed by the door on the direction of the motion, the resulting trajectory is completely different from the previous case. To let the reader better understand why and how we use the changing APF described in Section III-A, we report not only the optimized trajectory in Fig.14, but also the whole evolution of the potential profile when the robot is in the states mentioned in SectionIII-A, see Fig.l3. percentage equals to 25%; the number of generation evolved is 60; the population size is 100 individuals; the likelihood of crossover and mutation is 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002026_icma.2007.4303538-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002026_icma.2007.4303538-Figure2-1.png", "caption": "Figure 2. Hart Walker", "texts": [ " Locomat [1] by ETH and robotics stepper by NASA and UCLA [2] are good and only examples for the active walker which consists of treadmill and manipulator attached to the body. They are very sophisticated though, they are very expensive and cannot use in daily life. Although wheel chair is normally used for people who is gait disorder, it makes disuse syndrome (amyotrophia, arthrogryposis and impediment of the circulatory system) and therefore to keep upright position and walking are very crucial indeed. We have been developing the active walker by using the Hart Walker: HW (Fig.2) which consists of double upright knee ankle foot orthosis and 4-wheeled carriage with a stem located in the center of the carriage. Since waist part of the orthosis is attached to the top of the stem, there is no risk for falling, it is possible to keep the right posture, and both hands becomes completely be free. By applying the McKibben artificial muscle to the Hart Walker as the active walker, people can walk by health-people-like gait even if he/she has no muscular force at all. Thus the active walker is simple, inexpensive and possible to use in daily life", " To realize health-people -like gait by the active walker, we analyze human gait and acquire the ideal gait pattern for the active walker. Since it is very difficult to use people for experiment in this stage from ethical point of view, we apply a doll with the same kind of joints human has and weight. Feedback control method is applied to implement health-people-like gait and we find that our method is feasible and very flexible in weight and height change. 1-4244-0828-8/07/$20.00 \u00a9 2007 IEEE. 186 A hart walker shown in Fig.2 was developed in 1981 at England for mainly applying to infantile paralysis. 4100 children are using all over the world and more than 180 ones in Japan. It consists of knee-ankle-foot orthosis and carriage. The greatest asset of the hart walker is that user\u2019s hands become free and user can keep right posture. Load to the leg is controllable by modifying the position of connecting point between the stem and knee-ankle-foot orthosis. Moreover it is very easy to adjust the length of frames to the body as shown in Fig", " Also the doll can walk very naturally. While gait itself was unstable in various situation, i.e., not flexible in floor condition, weight change, and height change. Feedback control is applied as shown in next section. For controlling , actuator A and B in Fig. 6 is utilized. For , C and D. Control block diagram is shown in Fig.15. Variables in this figure describes as follows; Using a child-size doll (130cm height, 26kg weight) with the same kind of joints human has, walking experiment is undertaken. As mentioned in Fig. 2, HW has function to change the length easily. In case of 130cm height, we can change height in the range of +10cm to -10cm. Following by height change, weight should change in the range of +5kg to -5kg. Since we want to show how adaptive the active walker is in terms of height and weight change, we investigate experiments for several combinations with respect to height and weight as shown in Table 2. The most extreme combinations are 4 ones; +10cm and +5kg, -10cm and -5kg, +10cm and -5kg and -10cm and +5kg" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000381_j.cam.2005.03.092-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000381_j.cam.2005.03.092-Figure2-1.png", "caption": "Fig. 2. Vertex decimate in data simplification.", "texts": [ " Cararetta, Dahmen and Micchelli presented a non-stationary subdivision scheme for interpolating a set of given data points in 1991, which is a generalization of the four point subdivision scheme to the non-stationary case. If the initial data lie on a C2(R) function, then the limit function of the scheme approximates the original function quadratically. First, we sample the curve to obtain N source points {Si}Ni=0 in which the intrinsic definition of a planar curve is used [16], and then connect them orderly to form the initial polyline. We simplify it by deleting the points of \u201cless importance\u201d. The importance of a vertex is estimated by its directional angle i (demonstrated in Fig. 2). If the absolute value of the angle is less than a given threshold , delete the point Si , and record the lost arclength. Finally, we get the rest points and connect them to form a new polyline, which we call control polyline (denoted by {Pi}ni=0). The other data obtained after simplification is the arc-length reduction {Li}n\u22121 i=1 , where Li is the arc-length lost during the deformation between control points Pi and Pi+1. After simplification, the control polyline will reserve enough points in the high-detailed region and less data in the flat region" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002182_ecctd.2007.4529762-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002182_ecctd.2007.4529762-Figure7-1.png", "caption": "Fig. 7. \u03c7\u03032 \u2229 \u03c72 intersections showing the DC solution regions.", "texts": [ "33 \u2212 1000 (iD1 + iD2) (9) vD2 = e3 \u2212 e1 = 4 \u2212 1392iD1 \u2212 1096iD2 (10) where {e1, e2, e3} are nodal voltages. Notice that equations (9) and (10) can be rewritten into the form of equation (5) as follows: [ 1 0 0 1 ] [ vD1 vD2 ] + [ 1000 1000 1392 1096 ] [ iD1 iD2 ] \u2212 [ 3.33 4 ] = [ 0 0 ] (11) The graphic coordinates for the PWL curve in Fig.5 are collected in the set \u03c7n (for n = 1, 2). \u03c7n = { (\u22121500,\u22121\u00d7 10\u22129), (0.325, 0), (0.5, 0.014) } (12) By applying the graphical procedure reported in Section IV, the intersections \u03c7\u03032 \u2229 \u03c72 are obtained. Fig.7 shows three \u03c7\u03032 \u2229 \u03c72 intersections. The first occurs between the segment (1) of \u03c7\u03032 and the segment (2) belonging to \u03c72 curve. The second results from the segment (2) of \u03c7\u03032 and segment (2) of \u03c72 and finally the third intersection is related with the segment (2) of \u03c7\u03032 and the segment (1) of \u03c72 \u03c7\u03032\u2229\u03c72. As a result, it can be concluded that although in this example there are four DC element regions, only (21), (22) and (12) are the DC solution regions. According with the proposed denotation, k1 and k2 are variables that select a specific segment from the PWL curves vD1iD1 and vD2-iD2, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003954_iros.2010.5650000-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003954_iros.2010.5650000-Figure1-1.png", "caption": "Fig. 1. Contact states generated with (a) a vertex-face contact and (b) a vertex-triangle contact.", "texts": [ " First we triangulate all faces of a polyhedral object. If all faces are triangles with three vertices and three edges, we can apply a single rule to find neighboring contact states. A sub-contact is newly defined as a vertex-triangle contact. Then a contact state is analytically represented by using two sub-contacts of the vertex-triangle contact and the edge-edge contact. This method can distinguish contact situation more clearly than the vertex-face contact. For example, in case of the vertex-face contact of Fig. 1(a), we have to investigate 978-1-4244-6676-4/10/$25.00 \u00a92010 IEEE 4522 about all vertices, edges, and faces neighboring it to generate contact states. On the other hand, the search range is restricted to only three vertices, three edges, and three adjacent triangles for the vertex-triangle contact of Fig. 1(b). Therefore, in view of the contact state generation, the vertex-triangle contact can reduce the search range of contact states. This paper is structured as follows. Section II-A describes the definition of the sub-contacts and triangulation for the vertex-triangle contact is described in Section II-B. Adjacent sub-contacts from the current sub-contact are defined according to the type of adjacent triangles in Section II-C. Using the adjacent sub-contacts, the framework for automatic generation of the contact state graph is presented in Section III" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001447_ijvd.2005.006606-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001447_ijvd.2005.006606-Figure4-1.png", "caption": "Figure 4 Control parameters in a constant radius turn with an acceleration", "texts": [ " Longitudinal acceleration introduces the longitudinal weight transfer, in addition to the lateral weight transfer caused by the centrifugal force. The results of numerical simulation are shown in Figures 4 and 5. The sum of the active suspension forces of the three-wheeled vehicle also is incorporate into Figure 5. As the speed goes up, the active suspension force applied on the single wheel either on the front or on the rear-axle remains about the same value. Yet, one of the two active suspension force, paired on an axle will change its direction, such as Fz1 for the mode 1 and Fz4 for the mode 4, as shown in Figure 4(c) to (f). The phenomenon, active suspension forces changing the direction of its application while speeding-up, which appears in three-wheeled vehicle and modes 1 and 4 on a four-wheeled vehicle is the result of taking the dynamic weight transfer caused by both the inertia and centrifugal force. Therefore, the sum of the active suspension forces in modes 1 and 4 on a four-wheeled vehicle are less than that in three-wheeled vehicle, and that in a three-wheeled vehicle is less than that in modes 2 and 3 on a four-wheeled vehicle, as shown in Figure 5(c)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001083_sensor.2005.1496627-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001083_sensor.2005.1496627-Figure1-1.png", "caption": "Figure 1. Schematic view of the tunable liquid micro-lens system. The liquid droplet is positioned in an etched silicon cavitj and the remaining portion of the cavityfilled with a different liquid of the same density.", "texts": [ " This yields an relation known as the Lippmann equation [6], CEO p cos( e ) = cos( e,) + - Y L G ~ relating the initial contact angle 00. the dielectric constant of the dielectric layer E , the surface energy between droplet and surrounding YLG, the thickness of dielectric layer d and the applied voltage V . Because of the l/d dependence of the dielectric layer thickness, it is easy to see that using thin dielectric films will allow a reduction of the required applied voltage. STRUCTURE The system presented here and illustrated in figure 1 is based on a n-doped silicon wafer structured by KOHetching and covered with thermally deposited silicon oxide. While the 54 o sidewall V-grooves act as a dropcentering device, necessary to keep the lens in the desired position on the optical axis, the silicon dioxide functions TRANSDUCERS05 The 13th International Conference on Solid-state Sensors, Actuators and Microsystems, Seoul, Korea, June 5-9,2005 0-7803-8952-2/05/$20.00 @ZOOS IEEE. as the dielectric layer. Thermal silicon dioxide features high dielectric strength, conformal, defect-free covering at very thin film thicknesses and a moderate E which reduces the required applied voltage" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001931_j.tcs.2007.10.028-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001931_j.tcs.2007.10.028-Figure5-1.png", "caption": "Fig. 5. The substitution rule for the Ammann\u2013Beenker tiling (left), with prototiles S, L , and the star-dual substitution (right), with prototiles S?, L?. Even though the two substitutions look different on a first glance, they define very similar tilings, see Fig. 6.", "texts": [ " The tile substitution arising from (6) is well-known: It is the Tu\u0308bingen Triangle tile substitution (see Fig. 4, right). See [9] for more details and images of this substitution, as well as several others. Altogether we have established: Theorem 7. The ?-dual of the Penrose substitution is the Tu\u0308bingen triangle substitution, and vice versa. 3.2. The ?-dual of the Ammann\u2013Beenker tiling In a very similar fashion, one can compute the star-dual of the Ammann\u2013Beenker tiling [11]. The substitution rule of the Ammann\u2013Beenker tiling is shown in Fig. 5 (left). A part of an Ammann\u2013Beenker tiling is shown in Fig. 6 (left). The Ammann\u2013Beenker tiling is closely related to the cyclotomic field Q(\u03be8), where \u03be := \u03be8 = e 2\u03c0 i 8 . Its ?-dual can be calculated in a very similar fashion as in the case of the Penrose tiling. The substitution factor is \u03bb = 1 + \u221a 2 = 1 + \u03be + \u03be7 = 1 + \u03be + \u03be\u22121, its inverse \u03bb\u22121 = \u22121 + \u03be + \u03be\u22121. Let S be the triangular prototile, with vertices 0, i, 1+ i , and let L be the rhombic prototile with vertices 0, 1, 1+ \u03be, \u03be . The IFS for these prototiles is L = f1(L) \u222a f2(L) \u222a f3(L) \u222a f6(S) \u222a f7(S) \u222a f8(S) \u222a f9(S), S = f4(L) \u222a f5(L) \u222a f10(S) \u222a f11(S) \u222a f12(S), where f1(x) = \u03bb\u22121\u03be4x + 1+ \u03be, f7(x) = \u03bb\u22121\u03be3x + 1+ 2\u03be + \u03be2 + \u03be7, f2(x) = \u03bb\u22121\u03be x + 1+ \u03be + \u03be2 + \u03be7, f8(x) = \u03bb\u22121\u03be6x + 2+ \u03be + \u03be2 + \u03be7, f3(x) = \u03bb\u22121\u03be7x + 1+ \u03be + \u03be2, f9(x) = \u03bb\u22121\u03be7x + 1, f4(x) = \u03bb\u22121x + \u03be + \u03be2, f10(x) = \u03bb\u22121\u03be5x + \u03be + \u03be3, f5(x) = \u03bb\u22121\u03be2x + \u03be, f11(x) = \u03bb\u22121\u03be6x + 1+ \u03be + \u03be2, f6(x) = \u03bb\u22121\u03be2x + \u03be, f12(x) = \u03bb\u22121\u03be3x + 2\u03be + \u03be2. The functions f ] i of the dual IFS are easily obtained by considering the inverse functions f \u22121 i , and replacing each \u03be occurring by its Galois conjugate \u03be3 (or \u03be5 or \u03be7, the resulting dual tiling is always the same). The dual IFS reads: L? = f ] 1 (L?) \u222a f ] 2 (L?) \u222a f ] 3 (L?) \u222a f ] 4 (S?) \u222a f ] 5 (S?), S? = f ] 6 (L?) \u222a f ] 7 (L?) \u222a f ] 8 (L?) \u222a f ] 9 (L?) \u222a f ] 10(S?) \u222a f ] 11(S?) \u222a f ] 12(S?). Its solution are the two tiles shown in Fig. 5 (right): S? is an isosceles orthogonal triangle, and L? is a thin orthogonal triangle. Note that the original triangular prototile S can be dissected into two copies of S?, and the rhombic prototile L can be dissected into four copies of L?. In fact, unlike in the Penrose case, the ?-dual tiling of the Ammann\u2013Beenker tiling is closely related to the Ammann\u2013Beenker tiling itself, compare Fig. 6: dissecting each L into four tiles, and each S into two tiles, yields the ?-dual tiling of the Ammann\u2013Beenker tiling" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003829_j.jallcom.2010.07.142-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003829_j.jallcom.2010.07.142-Figure4-1.png", "caption": "Fig. 4. (1 1 1), (2 2 0), and (2 0 0) po", "texts": [ " le figu E t c t [ t v fi p a c fi c f a t c b g F m p A recent work suggests that the growing columns obtained by B-PVD have a certain crystallographic orientation. Vapor deposiion has also been reported to result in the preferential growth of ondensates [17]. This preferential growth causes crystallographic exture, which has been explained by \u2018evolutionary selection\u2019 17,18]. The theory of \u2018evolutionary selection\u2019 states that the crysal with the fastest growing direction orients itself towards the apor source. Fig. 4 shows the typical (1 1 1), (2 2 0), and (2 0 0) pole gures of the as-deposited alloy sheet. The (1 1 1) plane is nearly arallel to the substrate surface, but the poles of (2 2 0), and (2 0 0) re distributed around the apex of the figure in an approximately ircular fashion. This result suggests that a predominantly \u3008 1 1 1\u3009 ber texture is formed. A number of different textures in EB-PVD oatings, e.g., \u30080 0 1\u3009, \u30080 1 1\u3009, \u30081 1 3\u3009 and \u30081 1 1\u3009-type textures, were ound on rotated and stationary substrates [19\u201321]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001061_3-540-26415-9_112-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001061_3-540-26415-9_112-Figure1-1.png", "caption": "Fig. 1. Cross-section of ship showing two FPSO tanks in the middle and two ballast tanks on the sides.", "texts": [ " This cost rises to \u00a3150-200k to inspect 3 pairs of cargo tanks and 3-4 pairs of ballast tanks after ten years. These costs can be reduced substantially by sending a robot into the tank without first emptying it thereby saving the cost of cleaning and emptying the tanks. Weld cracks are caused by fatigue and are of two types. Low-Cycle fatigue is driven by panel deflection when filling and emptying tanks causes cracks at the toe of a bracket, generally in the secondary material. The drawing on the bottom right in figure 1 shows a bracket with cracks at its toe ends and the location of the brackets in the FPSO tank. High-Cycle fatigue is driven by wave pressure on the side and bottom shell of the tanks. It causes cracks at cut-outs where shell longitudinal strengthening plates connect to cut outs in the frames. Figure 1 is a drawing (bottom left) of cracks at cut outs and the location of longitudinal strengthening plates. It is also required to test for corrosion caused by coating breakdown on the tank bottom. Pits can develop at the rate of 2-3 mm/year and even faster at the rate of 5mm/year if more corrosive crude is present. The bottom plate is usually 18-25mm thick. Obtaining access to welds on strengthening plates on the walls and the floors of the tank is not easy. The environment is very cluttered so that a very large walking and climbing robot would be required to step over plates" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000247_iecon.2002.1182940-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000247_iecon.2002.1182940-Figure1-1.png", "caption": "Fig. 1. Dimensions and frequencies related to bearing faults signatures.", "texts": [ " The objective of this research is to identify hearing faults via machine vibration before catastrophic failure OCCUTS. 11. BEARING FAULT SIGNATURES Bearing faults take the form of mechanical defects or deformations in the raceways and rolling elements. Associated with these defects are the four characteristic fault frequencies and the rotor frequency. A defect on a given hearing surface will typically produce a frequency component at the corresponding characteristic fault frequency in the vibration spectrum. These frequencies are illustrated in Fig. 1 where FR = rotor (shaft) frequency FcF = cage fault frequency FlnF= inner raceway fault frequency 0-7803-7474-6102/$17.00 02002 IEEE FoRF = outer raceway fault frequency FBF = ball fault frequency DB = hall diameter Dp = pitch diameter NB = number of rolling elements SE = direction of force exerted by the rolling element on the outer raceway B= ball contact angle. The characteristic fault frequencies can he calculated using (1)-(4) [4], and a derivation of these equations is presented in [SI. F - L F R ( l - D, cos(e) Dp ) c F - 2 Bearing faults begin as localized defects on the raceways (or rolling elements), and, as the rolling elements pass over these defect areas, small collisions occur producing mechanical shockwaves", " The severity of this dcfcct is intended to rcpresent a more advanced stage fault. These results are illustrated in Fig. I 1 Duc to the advanced nature of this dcfcct, a peak near the outer raccway fault frequeiicy is clearly evident in thc power spectrum estimate of Fig. I I(top). Because ( I j through (4) are functions of 0, it is common for slight amounis of mechanical misalignment (or similar sourccs) to cause the characteristic fault frcquencics to deviate slightly from their exact predicted values. The application ofthc AM detector to this data is illustrated in Fig. 1 I(bottom). In this plot, there is a substantial amount of AM indicated. All significant peaks (above 500 Hz on the horizontal axis) are separated by approximately 212 Hz in the horizontal and vertical directions. This suggests the presence of a significant fault on the outer raceway. Squared Bicoherence I5O0\u2019 - 2 1000 I 6 c m 3 500 LL I I - I \u20180 500 1000 1500 2000 2500 Frequency (MI Fig. IO. Squared bicoherence applied to the same bearing vibration data. While B minor amount of QPC is indicated, the arbihary spacing makes it impossible to determine if this is the result of a bearing fault or some other non-fault some" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002893_dscc2008-2112-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002893_dscc2008-2112-Figure7-1.png", "caption": "Figure 7: Fish\u2013shaped part schematic showing modified powder flow rate reference parameters.", "texts": [ " However, the height variation at the sharp corners is not improved due to the bandwidth limitations of the powder feeder system and the fact that the process dynamics significantly change when the motion system slows down and the laser delivers more energy. Also, the excessive height where the deposition paths intersect is not affected since it is not considered in the powder flow rate reference design. The VPFRC is utilized with a modified powder flow rate reference for the third fish\u2013shaped part. The powder flow rate reference is adjusted manually, as shown in Figure 7, at the locations where excessive deposition exists. The parameters for these adjustments are listed in Table 1. The strategy is to deposit less powder where there is excessive deposition, to deposit powder only once per layer at the location where the deposition paths intersect, and to make the transitions smoother (i.e., decrease the powder flow rate reference derivatives). The modified powder flow rate reference and its first and second derivative are shown for one layer in Figure 8. As compared to Figure 4 the peak powder flow rate reference first and second derivates have been substantially reduced" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000983_icit.2003.1290263-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000983_icit.2003.1290263-Figure8-1.png", "caption": "Fig. 8: The Robot Model in Aerial Phase", "texts": [ " % = JF'IGwG (19) In HVC, the hopping velocity wV&h which is to realizing the desired hopping height \"'hZmd is acquired. Therefore, the height z, ( t ) in aerial phase is derived as follows. (20) Q 2 2 zo ( t ) = --t + WVGzht W h C h Here, if the height at the instant of the next landing is set to himd, the period t , in aerial phase is derived as follows. 9 9 0 = --t2 + WVGzht + (whGh - h r d ) J w chg - uGxh - (15) The velocity command of joint 3 can be derived from Eq. (lo), Eq. (13) and Eq. (15). Fig. 8 shows the posture change in aerial phase. Here, the attitude angular change y of the robot is as follows. iS\"d-\" C h 9 Sin e, I , cos - t voZ,,sine, I J cos y y = arctan J wVGzh2 w + '\"VCyh' (22) VGzh EerT < 0 : 8;m-d = - E\"' 2 0 : 0;md - 0 (16) + arctan d w U G z h 2 + \"VGyh2 p d yl =h? ~ V G . , . I , ~ + 29 ( w h c h - wh;md) The angular velocity WG of COG is derived from Eq. (21), Eq. (22). This is the posture command which is defined for the next landing state. It is substituted for Eq. (19) and changes into an angle command" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002491_1.3159892-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002491_1.3159892-Figure2-1.png", "caption": "FIG. 2. Three dimensional mesh used in plasma simulation 154 488 cells .", "texts": [ " In turbulent [This article is copyrighted as indicated in the article. Reuse of AIP content is subject to the terms at: http://scitation.aip.org/termsconditions. Downloaded to ] IP: 150.135.239.97 On: Wed, 17 Dec 2014 21:30:58 flows, FLUENT computes the mass diffusion in the following form: J i = \u2212 Di,m + t Sct Yi , 7 where Sct is the turbulent Schmidt number which is taken to be 0.7 by default, t is the turbulent viscosity and Di,m is the diffusion coefficient for metallic vapor in the mixture. Shown in Fig. 2 is the mesh used in the simulation of laser induced plasma. The intersection of the laser beam axis and workpiece top surface is the origin of the Cartesian coordinate system. Boundary conditions used in this study are shown in Fig. 3. Details of boundary conditions employed in simulation are as follows. Metal vapor with fixed temperature enters into the computational domain at constant mass generation rate at the bottom of the keyhole. The mass generation rate T\u0304 is determined by conservation of energy" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002632_s11071-009-9565-1-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002632_s11071-009-9565-1-Figure1-1.png", "caption": "Fig. 1 A two-link robot", "texts": [ " Choose the proper values \u03c30, \u03c31 in Assumption 1 and FU in Assumption 2. Choose the value r to determine R in (15). Select the values \u03b31, \u03b32, \u03b33 in (43)\u2013(45). Step 6: Obtain the fuzzy controller (14) and the adaptation laws (43)\u2013(45). In the following section, a simulation example is given to show the effectiveness of the proposed approach. 4 Simulation example In this section, we test the proposed observer-based adaptive fuzzy control design on the tracking control of a two-link robot. Consider a two-link manipulator described in Fig. 1. The dynamic equation of a two-link robot arm can be expressed as [17] M(q)q\u0308 + c(q, q\u0307) + g(q) + \u03c4c(q, q\u0307) + \u03c4d(q, q\u0307) = u (52) where M(q) \u2208 R2\u00d72 is the bounded positive-definite inertia matrix, c(q, q\u0307) \u2208 R2\u00d72 the vector representing centrifugal and Coriolis effects, g(q) \u2208 R2 the vector representing gravitational torques, and \u03c4c(q, q\u0307) \u2208 R2, \u03c4d(q, q\u0307) \u2208 R2 are the vectors representing the dynamic effects as nonlinear frictions, small joint and link elasticities, backlash and bounded torque disturbances" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002441_detc2009-86358-Figure12-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002441_detc2009-86358-Figure12-1.png", "caption": "Figure 12. Derivation of Crowning Radius", "texts": [ " Some gears are tilted by deflection, so the center of the crown is not at the center of the tooth face (called bias crown). Some gears have more misalignment than deflection, as in the case of crowned splines, and contact ellipse is small compared to the face width. Here, the center of the crown is at the center of the tooth face (called full crown). A crowned spline has one more limitation: The tooth thickness has to be modified from the standard because the minimum effective clearance is zero. The tooth thickness Tmod is dependent on R1 and \u03b8. From Figure 12, the following equations can be derived. \u03b8cos/)2/2/( mod dTTX += (14) \u03b8tan/Xy = (15) L1 R1 Drop Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 5 Copyright \u00a9 2009 by ASME )cos1(2/ \u03b8+\u22c5= yG (16) \u03b8tan)2/( \u22c5= GC (17) 22 1 2 1 )2/()( GCRR +\u2212= (18) Where, C is drop in the normal plane, G is total gage length and gage from center of crown is G/2 and dT is the clearance between space width of internal teeth and tooth thickness of external teeth" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000746_12.619143-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000746_12.619143-Figure1-1.png", "caption": "Fig. 1 Schematic of a biconvex singlet lens mounted in a simple cell", "texts": [ " To minimize this threat, it is advisable to apply sufficient preload at assembly PA so the residual preload existing at T will hold the lens against the mechanical interface under the maximum expected axial acceleration. The following equation defmes the minimum required PA: PA Wa -K3(T - TA) (3) where W is the lens weight and aG is the acceleration level The factor K3 that makes these types of estimations possible depends upon the optomechanical design of the subassembly and the pertinent material characteristics. It is difficult to quantify completely and accurately, even for a simple lens/mount configuration. For example, consider the design shown schematically in Fig. 1 . Here, a biconvex lens is clamped axially with some nominal PA in a cell between a shoulder and a threaded retainer. The glass-to-metal interfaces appear in the figure to be sharp corners, but tangential or toroidal interfaces would be more appropriate in an actual design. Key mechanical changes that can occur in this design and that contribute to the magnitude of its K3 factor are as indicated in Table 1. Designs, such as those for cemented doublet lenses (see Fig. 2) or lenses separated by spacers (see Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003904_2013-01-2373-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003904_2013-01-2373-Figure1-1.png", "caption": "Figure 1. Two DOF vehicle model", "texts": [ " Generally, the most important vehicle state variables which illustrate vehicle lateral stability are lateral velocity v and yaw rate r. In this paper, it is assumed that the vehicle longitudinal velocity u is a constant. On this basis, considering the nonlinear characteristic of the tire, it is possible to obtain the vehicle lateral stability region based on lateral velocity and yaw rate using lyapunov's second method under particular surface condition, longitudinal velocity and steering angle. The 2 DOF vehicle model is shown in Figure 1. Vehicle equations of motion in lateral and yaw directions can be expressed as follows: (1) (2) Where m is the mass of the vehicle, Iz is the yaw inertia; Fyf and Fyr are the total lateral forces exerted on the front wheels and rear wheels, respectively; lf and lr are the distance of c.g. from the front and rear axles respectively; \u03b4 is the road wheel angel of the front wheels; u, v and r are the longitudinal speed, lateral speed and yaw rate, respectively. Nonlinearity of the vehicle is mainly embodied in the nonlinear characteristic of the tire" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002072_s10589-007-9090-4-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002072_s10589-007-9090-4-Figure1-1.png", "caption": "Fig. 1 Numerical convergence example", "texts": [ " If the algorithm eventually stays computing iterations at Step 2, the sufficient-decrease criterion for the KKT norm guarantees that KKT points are computed with an arbitrary precision. Otherwise, the algorithm shares the global convergence properties of the Augmented Lagrangian Method [1, 2]. In what follows, we present a very simple numerical example which illustrates the acceleration property of the proposed algorithm. Consider the problem of solving min ((x \u2212 2)2 + (y \u2212 1)2) sin((x \u2212 2)2 + (y \u2212 1)2) s.t. x \u2265 0, y \u2265 0, x \u2212 2y + 2 \u2265 0, x \u2212 3y \u2212 4 \u2264 0, 2x + 5y \u2212 12.73 \u2264 0. In Fig. 1, we can see: \u2022 the feasibility region (the shaded pentagon); \u2022 the level curves; \u2022 the initial vector: v0 = (5;1)T (black square); \u2022 a local maximizer (black triangle); \u2022 the global minimizer (black circle). These results correspond to the Augmented Lagrangian working alone (AL), the SR1 working alone and the combined Augmented Lagrangian-SR1 algorithm called the Accelerated Augmented Lagrangian (AAL). We observe that we started all the algorithms with the vector v0. The SR1 algorithm alone converged to the maximizer shown in the figure (black triangle)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000610_bf02138004-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000610_bf02138004-Figure5-1.png", "caption": "Fig. 5", "texts": [ " - - 2 \" - - : . 080 (185 G80 G95 1 105 1.10 1.15 1 l j / \u00b0 \u00b0 , \" ~ , , _ 00eo DO 0 e _ \u2022 ~ . . . . . ----~. 1.20 125 130 1.38 1AO 1.45 ItS, !.50 FREOUENCY Fig. 7 t=el mode z2 have been drawn. The points referring to the secondary flexural strains, have ~ been joined by broken lines to sho~\" up bet.ter the resonances due to lateral vibrations. To distinguish the strains in sections 3 and 5 from the ones in sections 4 and 6, we'll indicate them with I and II respectively. Indexes I and II will indicate (fig. 5) the axial planes passing through points 3, 4 and 5, 6 respectively, and the direction constituted by their traces. The examination of the results related in fig. 7 will be developed considering the different resonances first; in a second moment we'll try to make clear the probable causes of discordances between theoretic and experimental results. Fig. 7 shows clearly the presence of 5 resonances and one less maxkes. A ) Axial or torsional resonance: It arises for a value of excitation frequency f z equal or very near to the or", " C) Zmteralflexural resonance : Given the above mentioned frequency limit, we have been able to show only the two resonances corresponding to the spring lateral displacements in two different vibrations planes, is The relative frequencies indicated by f a and f 4 are : f a = l . 1 2 f l ; f 4 = l . 1 9 f l In correspondence with the first one, the spring lateral displacements direction is relatively close to the direction II, while in correspondence with f 4 such a direction almost coincides with the I one. Comparing these directions with the ones singled out for the spring 1, and considering the disposition of the strain gauges as fig. 5a and 5b show, we may deduce that, in both cases, in correspondence with the lower frequency (f3) the spring deformates in the axial plane passing through the ends of the two end coils; for the higher frequency (3~) the lateral deformation plane is approximatively the one normal to the first. The theoretic value of the lateral vibrations frequencyf L is given by: f L = 1 .07 f . The percentual differences of the experimental values are 4.6% and 11.2% respectively; they are almost equal to the ones found in the spring no" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001611_icsens.2007.355544-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001611_icsens.2007.355544-Figure1-1.png", "caption": "Figure 1. Structure of the flexible and biocompatible glucose sensor for tear glucose measurement. The sensor has a laminar structure of functional polymer and film electrodes. The sensor electrochemically measures hydrogen peroxide generated by the enzyme reaction at GOD immobilized PMD membrane.", "texts": [ " In particular, phospholipid polymer, which is so-called MPC polymer, was used for the sensing region. MPC polymer has molecular configuration which is similar to a cell membrane, carried out by the techniques of polymer chemistry [15-16]. In this paper, details of the structure, fabrication, basic characteristics of the glucose sensor using phosholipid polymer is presented. Also, the result of hydrogen peroxide measurement through PMD membrane is reported. II. EXPERIMENTAL SECTION A. Construction ofaflexible glucose sensor The structure of the flexible glucose sensor is illustrated in Fig. 1. The sensor was constructed by immobilizing glucose oxidase (GOD: EC1.1.3.4, Wako Pure Chemical Industries Ltd., Japan) onto a flexible hydrogen peroxide electrode. The hydrogen peroxide electrode has a Pt working electrode and an Ag/AgCl reference/counter electrode formed on a 3mm x 50 mm x 350 pm polydimethyl siloxane (PDMS) membrane. The scale of the sensor was optimized for the purpose of attachment and handling in bioinstrumentation. The gas-permeable PMD membrane was placed on the other surface of the electrodes" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003437_9781118516072.ch4-Figure4.8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003437_9781118516072.ch4-Figure4.8-1.png", "caption": "Figure 4.8. The basic configuration of a SCIG wind turbine [7].", "texts": [ " The much simpler construction of the machine and the gain in dynamic characteristics must, however, be set against the severe disadvantage of feeding electricity through brushes and slip rings. Higher frictional losses, brush and slip-ring erosion, and higher maintenance costs are the consequences. 4.3.1.2 Squirrel Cage Induction Generator. The simplest electrical topology of a wind turbine system incorporates a fixed-speed wind turbine with a squirrel cage induction generator. The basic configuration of the wind turbine system is shown in Figure 4.8 [7]. The main components of a SCIG wind turbine system are turbine aerodynamics, blade control system, mechanical drive train, induction generator, reactive power compensation device, coupling transformer, protection (especially under voltage protection). This type of wind turbine is directly connected to the electrical network. For this reason it is simple and cheap. Furthermore, no synchronization device is required. However, the wind turbine has to operate at constant speed as the frequency of the grid determines the speed of the generator rotor", " As the diameter of the turbine rotor increases, the permanent magnet generators seem to be the first choice for the wind turbines manufacturers. 4.3.2.1 Fixed-Speed Wind Turbines. The early large-sized wind turbine technologies were based on generators operating at fixed speed, that is, the rotor speed is constant regardless of the wind speed, and is determined by the power grid frequency, the generator characteristics, and the gear ratio. The induction generator (squirrel cage or wound rotor) is more suited to operate at fixed speeds and its stator can be connected directly to the power grid through a soft-starter (see Figure 4.8). In case of a squirrel cage generator, the slip and hence the rotor speed can be varied, but the variation is very small so that the wind turbine is referred to as a fixed-speed system. Induction generators operate at higher speeds, and therefore a gearbox is used to transfer the mechanical energy from the low-speed shaft of the aerodynamic rotor to the high-speed shaft driving the generator. For size and cost reasons, the induction generator operates at a standard nominal speed of 1500 rpm. In terms of the rated power, the gearbox must provide a gear ratio of about 30\u2013100" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003125_fie.2009.5350746-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003125_fie.2009.5350746-Figure4-1.png", "caption": "FIGURE 4 THE LOAD SPECTRUM ANALYZER\u2019S GRAPHICAL USER INTERFACE FOR THE SELECTION OF THE FRONT SUSPENSION\u2019S COMPONENTS", "texts": [ " The suspension sub-project team proposed and initiated a second semester student project for the development of a data evaluation computer program. This program should facilitate the design layout by delivering relevant information about the load spectrum of every individual component of the suspension. The acquired data together with a preceding finite element analysis enabled the minimization of suspension weight at equal strength. The project team developed a highly comfortable graphical user interface that allows the selection of the front and rear suspension components by mouse-click (see Figure 4). The data provided by the sensors are appropriately visualized and processed by the program. The load spectra for each individual component are determined from the recorded data and visualized as histograms. The rain-flow counting method is used in order to reduce the spectra of varying stress into sets of simple stress reversals. 978-1-4244-4714-5/09/$25.00 \u00a92009 IEEE October 18 - 21, 2009, San Antonio, TX 39th ASEE/IEEE Frontiers in Education Conference M4D-4 This is necessary for the application of Miner\u2019s rule for the estimation of fatigue life of a component subjected to complex loading conditions (see e" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001139_iros.1992.587374-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001139_iros.1992.587374-Figure6-1.png", "caption": "Fig. 6: Change of 6PM for a moving object (top view).", "texts": [ " reduction of the 6PM on one or more of the six faces ( T I , rz , p1, pz, z1, z2) with respect t o the previous sampling instant. Transformation After an object has appeared in the work-space it has to be mapped into e-space in a first step completely. Therefore it is descrihed by the 6PM arid afterwards every elementary cell is transformed by using the OCMEMalgorithm introduced in section 2.3. At every sampling t ime a new 6PM has to be calculated for all moving objects and compared with the 6PM of t,he previous step. Fig. 6 shows the top view of the 6PM for such an object in two succeeding sampling periods. For this special case an enlargement on the p - f a c e and the r.a-face as well as an reduction on pl-face takes place. The algorithm for mapping this change into c-space depends 011 the change: Enlargement/reduction at p-face: Because of the rotational symmetry of the robot system relating to the z-axis of the work-space an enlargement of the 6PM in work-space results in a similar enlargement of the forbidden region in c-space, as shown in fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002222_978-0-387-46283-7-Figure2.1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002222_978-0-387-46283-7-Figure2.1-1.png", "caption": "Fig. 2.1 Schematic experimental setup for the growth of one-dimensional oxide nanostructures via an evaporation-based synthetic method", "texts": [ " C\u00a9 Springer 2008 nanoparticles or thin films in an effort to enhance their surface sensitivity, and they have recently been successfully synthesized into nanowire-like structures. Utilizing the high surface area of nanowire-like structures, it may be possible to fabricate nano-scale devices with superior performance and sensitivity. This chapter reviews the general techniques used for growing one-dimensional oxide nanostructures. The vapor phase evaporation represents the simplest method for the synthesis of one-dimensional oxide nanostructures. The syntheses were usually conducted in a tube furnace as that schematically shown in Fig. 2.1 [2]. The desired source oxide materials (usually in the form of powders) were placed at the center of an alumina or quartz tube that was inserted in a horizontal tube furnace, where the temperatures, pressure, and evaporation time were controlled. Before evaporation, the reaction chamber was evacuated to \u223c1\u20133\u00d710\u20133 Torr by a mechanical rotary pump. At the reaction temperature, the source materials were heated and evaporated, and the vapor was transported by the carrier gas (such as Ar) to the downstream end of the tube, and finally deposited onto either a growth substrate or the inner wall of the alumina or quartz tube" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003913_2011-01-1400-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003913_2011-01-1400-Figure2-1.png", "caption": "Figure 2. Piston Ring Adapter and Holder", "texts": [ " A Plint TE-77 high frequency reciprocating tribometer combined with an ultrasonic pulsing system, shown in Figure 1, was used in this work. The TE77 is flexible tribometer that has been used by several researches [6, 7], for piston ringliner wear studies. A special piston ring adapter was used to retain a section of piston ring and load it against a section of liner. The conformability of the ring-liner contact was adjusted by two locating screws located at the both edges of the ring holder (see Figure 2). The normal force was transmitted directly onto the moving specimen by means of a needle roller cam follower on the adapter and a running plate on the loading stirrup. The oscillations were produced by a motor with an eccentric cam, scotch yoke and guide block arrangement. The assembly was mounted on flexible supports, which allowed free movement in horizontal directions. A stiff piezo-electric force transducer connected to the assembly measured the friction force in the reciprocating directions" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003756_icelmach.2010.5607764-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003756_icelmach.2010.5607764-Figure2-1.png", "caption": "Fig. 2. Schematic example of rotor excentricity (stator slotting and permanent magnets not shown)", "texts": [ " If the two waves of a couple have unequal circumferential travel-directions, additionally to 1., the resulting force vector circulates with an angular frequency abf ,, \u03bd\u03c9\u03bd\u03c9 +=\u2126 . In this case, the direction of travel follows the highest absolute ordinal number. These rules are useful in the design process of a machine (e.g. elimination of noise generating wavelengths in advance) as well as explaining the fundamental principle of a bearingless drive. Concluding this chapter, a concrete example of rotor excentricity is presented. Thereto the schematic assembly according to Fig. 2 will be used. Theoretically, a rotor displacement e causes excentricity field waves that force the rotor in the direction of displacement (see Fig. 2). The example refers to a 4-pole, permanent magnet excited rotor without stator excitation. To solve this field problem a static FE-calculation, with excentricity 2/1/ =\u03b4e , was carried out. The flux density components in the air-gap were evaluated and treated by FFT. Essential results of this calculation are presented in Fig. 3. Apparently new ordinal numbers due to the excentrical rotor appear next to the former ones. With respect to the deduced force conditions it is obvious that, concerning this new field waves, an excentricity force exF exists" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002149_iccas.2007.4406709-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002149_iccas.2007.4406709-Figure2-1.png", "caption": "Fig. 2. This Figure depicts the division of the space in four quadrants.", "texts": [ " (To decide whether to use the goal for motion or to select a sub-goal of the list). The algorithm returns: \u2022 OPENED: There exists a set of homo topic paths joining both locations in the tunnel. \u2022 CLOSED: There is no local path joining both locations within the tunnel. The result of this process is a sub-goal that can be reached from the current robot location. In order to select a sub-goal we first use the algorithm with the goal. If the result is closed, we choose the closest sub-goal to the goal that has a path that reaches it. An example is showed in the Fig.2. Let the frame of reference be the robot frame with target tF , with origin 0F = (0, 0, 0) and unitary vectors ),,( zyx uuu , where xu is aligned with the current main direction of motion of the vehicle. Let S be a security distance around the robot bounds and R is the radius of the robot. We divide the space in four quadrants depending on the relation between the robot configuration and the target direction as follows. Let e be a given vector, the quadrant is <\u22c5<\u22c5<\u22c5 \u2265\u22c5<\u22c5<\u22c5 <\u22c5\u2265\u22c5\u2265\u22c5 \u2265\u22c5\u2265\u22c5\u2265\u22c5 =\u2126 0(&)0(&)0( 0(&)0(&)0( 0(&)0(&)0( 0(&)0(&)0( )( 324 323 322 321 1 1 1 1 PPP PPP PPP PPP eeeeeeifC eeeeeeifC eeeeeeifC eeeeeeifC e (1) Where: 1P , 2P and 3P are the planes defined by ],,[ 0 zx uuF , ],,[ 0 tz FuF and ],,[ 0 ty FuF ; and yP ue = 1 , tzP Fue \u2297= 2 and ytP uFe \u2297= 3 be the normal to these planes respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002995_icicta.2009.604-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002995_icicta.2009.604-Figure1-1.png", "caption": "Figure 1. AUV frames", "texts": [ " Aiming at these reasons, the main contribution of this paper is to build the maneuverability model of the long- distance torpedo-like AUV, and in the model the effect of the ocean current disturb is considered, and the maneuverability of the torpedo-like AUV effected by the ocean current is simulated when the AUV hovering, and some conclusions are educed. II. DIMENSIONAL MANEUVERABILITY MODEL A earth-fixed frame 0 0 0 ( , , , ) E S o x y z and a body-fixed frame ( , , , )BS B x y z , as shown in Figure 1, are defined, and their transform matrixes are B EC and E BC . The AUV\u2019s velocity vector is considered to be T T T[ , ]=V v T[ , , ]x y zv v v=v T[ , , ]x y z\u03c9 \u03c9 \u03c9= , and the earth-fixed position/orientation vector is T T T[ , ]=R r T 0 0 0[ , , ]x y z=r T[ , , ]\u03b8 \u03c8 \u03d5= . A detailed explanation of this can be found in the literature [3]. The velocity vectors in the earth-fixed frame and in the body-fixed frame have the following relationship:[4][5] ( )R J R V= (1) where ( )J R is the rotation matrix, which is shown in equation (2)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003958_09205071.2013.753662-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003958_09205071.2013.753662-Figure4-1.png", "caption": "Figure 4. Bearing model velocities.", "texts": [ " The deformation of the balls is a combination of the waviness of the inner race (r), outer race (R), and ball (d), in addition to the shift in location of the shaft center position in the x and y directions that needed to maintain force equilibrium, and any bearing preload (e). The equation for the deformation of the ball is given by (9) [16]. di \u00bc r\u00f0hi\u00de \u00fe R\u00f0hi\u00de \u00fe di\u00f0wbt\u00de \u00fe x cos\u00f0hi\u00de \u00fe y sin\u00f0hi\u00de \u00fe e \u00f09\u00de where h \u2013 angular position of each ball about the axis of the bearing, x \u2013 angular velocity of the shaft, and wbt \u2013 angular position of each ball about its own axis (see Figure 4). If the inertia force of the spindle is ignored, the resulting contact force equilibrium equations are given by Equations (10) and (11) for the x and y directions, respectively. Xn i\u00bc1 kd1:5i cos\u00f0hi\u00de \u00bc 0 \u00f010\u00de Xn i\u00bc1 kd1:5i sin\u00f0hi\u00de \u00bc 0 \u00f011\u00de Because of the nonlinearity of the Hertzian contact force, the Newton-Raphson method was used to find a numerical solution to these equations. These allow us to find the vibration of the shaft. Figure 5 shows an artificial defect introduced by our model on the inner and outer bearing races" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000063_robot.1999.772568-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000063_robot.1999.772568-Figure1-1.png", "caption": "Figure 1 : Robot finger with soft tip in contact with a rigid surface", "texts": [ " Based on this last work, this paper examines the physical interaction between a rigid object and a soft tip robot finger with unknown non-linearities of the reproducing forces and under general kinematic (and dynamic) uncertainties arising ftom both the robot finger and the rigid object. The asymptotic stability of the force regulation problem under kinematic uncertainties is addressed for a planar three degree of fi-eedom (dof) robot finger with a soft tip which is required to exert a desired force on a rigid object by pressing on it. 2 Task description Consider a three degree of fteedom robot finger with soft tip which is required to exert a desired force fmi on a rigid object by pressing on it and which has initially established contact with the rigid object (figure 1). Let {T} be a fixed task frame placed on the rigid surface so that the x and y task coordinate align respectively with the surface normal and the surface tangent at the point of contact c. Let the 0 task coordinate denote the angle formed between the third link and the surface as shown in the figure. Let X = [x y e] f = [f, f,, fe] denote the position and the generalized force in fixed task coordinates. T T 0-7803-51 80-04/99 $10.00 0 1999 IEEE 1475 We assume that the reproducing force fx is dependent on the displacement Ax through a generally unknown or uncertain function f, = fx(Ax) which is however a strictly increasing function of Ax (figure 2)", " using (8) to express j in (22) with respect to J, we get: A-'X+e,(aAx+PAF) = cosGx+ sinGy+ (d cos8 - d, sin3)b + ( a h + PAF) jq + e, (aAx + PAF) = 0 (22) - sin Gx + cosgy + (-d sin - d, COS^)^ e COSGX + sinGy + (aAx + P A F ) = 0 - sin Fix + cosejr = 0 9 = 0 (23) It follows that: A% + cosG(aAx + PAF) = 0 (24) i s = o a We can solve the second equation of (23) foryand substitute in the first: Thus, This last equation implies that x + 0 , Ax -+ A% and fx+fh, as t+x [3]. But x + 0 implies y + 0 which in turn implies f,+O and 6 + 0 implies fe-+O. Thus, by virtue of Lasalle's invariance theorem the asymptotic convergence of the force error and of the velocities to zero has been proved. 5 Simulation experiments and results Simulation experiments for a 3 dof planar robotic finger have been performed under the control law (15) with parameter values given in [ 5 ] . The robotic finger is assumed to be initially in contact with a rigid horizontal surface as in figure 1 with its third link forming an angle of 0=-30\" with the surface and it is desired to exert a force of 5 N on it. The reproducing force is simulated by the non-linear h c t i o n fx(AX) = k,(Ax)X. Dynamic parameters are assumed to be known and damping factors negligible but kinematic uncertainties arise (i) from uncertain link lengths and (ii) fiom uncertain rigid surface orientation. In (i) link length estimates of 7, 6 and 5 cm were used as opposed to the actual link lengths of 5 , 4 and 3 cm. In (ii) the estimate of the surface orientation assumes that the rigid surface forms a 30\" angle with its actual horizontal position" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002542_00029890.2009.11920919-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002542_00029890.2009.11920919-Figure2-1.png", "caption": "Figure 2. (a) Midpoint M traces a circle of radius OM. (b) Triangle AOB divided into four congruent right triangles. (c) Diagonals of a rectangle are equal and bisect each other.", "texts": [ " The ellipsograph feature alone makes the trammel an important object of study, but it has other interesting properties discovered with methods of calculus that deserve to be better known. This article describes known properties of the standard trammel as well as new ones that can be studied by simple geometric methods that do not require calculus. The generalization to flexible trammels increases our understanding of these properties and also leads to engaging classroom activities. 2. ELLIPSE TRACED BY A POINT ON A STANDARD TRAMMEL. To see why a point on a trammel AB of fixed length L traces an ellipse, we show first that the midpoint M of AB traces a circle, as suggested by Figure 2a. Figure 2b shows segment OM, together with perpendiculars from M to each axis, dividing the large right triangle AOB into four congruent smaller right triangles, each having hypotenuse of common length OM = AM = MB = L/2. Therefore, M always lies on a circle with center at O and radius L/2. This also follows by completing the rectangle OACB in Figure 2c. The circle of radius L/2 has Cartesian equation ( x L/2 )2 + ( y L/2 )2 = 1. (1) February 2009] THE TRAMMEL OF ARCHIMEDES 115 This is a special case of the general equation of an ellipse, ( x a )2 + ( y b )2 = 1, (2) traced by a point E = (x, y) on the trammel that divides it into segments of lengths a = AE and b = EB, as indicated in Figure 3a. We deduce (2) from (1) by relating the coordinates (x, y) of E with the coordinates (X, Y ) of the midpoint M which satisfy (1). Similar triangles in Figure 3a reveal that x/a = X/(L/2) and y/b = Y/(L/2); hence (2) follows from (1)", " Instead of m2 + n2 being constant, we assume that the intercepts (m, n) satisfy a more general relation, say G(x, y) = 0, (14) where G is a prescribed function we call the governor of the trammel. If the intercepts m and n satisfy G(m, n) = 0, we refer to the portion of line (13) in a given quadrant as a flexible trammel governed by G. The length of the trammel is equal to that of the radial segment from the origin to (m, n), where G(m, n) = 0. This segment and the trammel are the diagonals of a rectangle with opposite vertices at the origin and at (m, n) as shown in Figure 15b. This extends Figure 2c for a standard trammel. When G(m, n) = m2 + n2 \u2212 1, the trammel has constant length 1 (it is a standard trammel) and any point P on it traces an ellipse. Now we seek the path traced by any point P on a flexible trammel as the trammel slides along the axes. We treat only the simple case when P divides the trammel into two pieces whose lengths are in constant ratio. It can be realized physically by an elastic string with a knot on it. Theorem 5. Consider a flexible trammel governed by a function G" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000173_gt2003-38718-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000173_gt2003-38718-Figure1-1.png", "caption": "Fig. 1 Brush seal geometry and boundary conditions.", "texts": [ " In order to explore brush seal bristle stresses with frictional effects, this paper presents a study using a 3-D finite element analysis. A maximum bristle stress relation is derived based on statistically designed experiments. Model accuracy is determined through verification simulations. A discussion on the effects of design and loading parameters on maximum stress is also included. A brush seal is a set of fine diameter metallic or ceramic fibers densely packed between retaining and backing plates. As illustrated in Fig. 1, L denotes free bristle length, \u03b8 denotes bristle lay angle, and BH shows free bristle height. Fence height (backing plate clearance) is denoted by FH, and R is used for rotor radius. The backing plate is positioned downstream of the bristles to provide mechanical support for the differential pressure loads. The circular seal is installed in a static member with bristles touching the rotor with an angle in the direction of the rotor rotation. In the case of rotor excursions, this cant angle helps reduce the contact loads allowing bristles to bend rather than buckle", " The pinch point at the inner edge of the retaining plate is the other location where bristles bend in radial plane due to the rotor excursion. THE 3-D FINITE ELEMENT MODEL The model consists of a representative bristle bundle with a backing plate and a rotor surface (Fig. 3). Every bristle is defined by a number of 3-D quadratic beam elements. The rotor and the backing plate are defined as rigid surfaces. A nloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?u representative backing plate is placed behind the first bristle row of bristles (see Fig. 4). Additional to the main design parameters defined in Fig. 1, analysis requires information on material properties, bristle diameter and friction coefficients for various contact locations. Modeling of bristle spacing involves layout and proximity. Within the brush pack two types of bristle layouts can be considered in the circumferential plane, namely, in-line or staggered (Fig. 4). The actual spacing will be a mixture of these two. Microscopic inspections reveal that in a typical seal most of the bristles tend to stay in a staggered configuration. Therefore, the presented analysis incorporates the staggered configuration to achieve a better simulation of the real case", " Although there is a reported value of 0.10 [16], studies by Crudgington [14] and Fellenstein et al. [17] set range of friction coefficients from 0.18 to 0.47. Therefore, values ranging from 0.25 to 0.35 are used in the current model. In order to accurately simulate clamping of the bristles at the outer radius, top nodes of bristles are fixed in space (all 6 degrees of freedom), while bristle tips are free to slide on the rotor surface. Bristles are modeled as staggered 3-D cantilever beams clamped at top nodes as shown in Fig. 1. Frictional contact defined at backing plate limits axial motion of bristles, while they are free to slide on the backing plate or bend below the fence height. Seal-rotor interference is simulated by applying displacements to rotor surface. Frictional contact defined at bristle tips allow bristles to slide or bend when rotor surface is moved towards the brush pack. To provide circumferential periodicity, first and last bristles in each row are coupled in a master-slave relationship. The last bristle at every row experiences a pull from the first, instead of the resistance from the rest of the bristles behind it" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000572_1.1637648-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000572_1.1637648-Figure6-1.png", "caption": "Fig. 6 Example", "texts": [ " satisfy a constraint equation below: lxlxp1lylyp50 (31) Therefore, under Parkin\u2019s definition of quatch, the set of screws associated with a finite line displacement can be represented by a subset of the 4-system. Note that under Dimentberg\u2019s definition of pitch @12#, the screws associated with a finite line displacement can be expressed as a 3-system @10,11#. The purpose of describing the screws related to a finite line displacement in this paper is to demonstrate the capability of the method to characterize the screws of both point-line displacements and line displacements in a unified way. Consider the initial and final positions A1B1 and A3B3 of a point-line ~Fig. 6!. a1 and a3 are the direction cosines of the pointline at the two positions and A1 and A3 are the position vectors of the corresponding end point positions, respectively. Let a15~0.6287, 0.7544, 0.1886!, A15~1.0, 22.0, 0.7!, a35~20.2403, 0.9013, 0.3605!, A35~22.0, 0.6, 1.5!. Note that the above initial and final point-line positions are defined in a global reference frame O-XYZ . To describe the screw system following the method introduced in this paper, a canonical coordinate system o-xyz must be determined at first" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000409_bfb0082609-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000409_bfb0082609-Figure1-1.png", "caption": "Figure 1.", "texts": [ " Z = 0, X is tangent to the foldl ine means that We see easily that the projections of the solutions of (V,X) on ~ 2 M(the (x 1 ,x 2 ) -plane) are integral curves of X = i~1 X{x 1 ,x 2 ,Jx1) %x i Next we perform a coordinate change :s = Jx 1 ; 0/ox 1 = (1/2s)0/os ; - 2 2 and get X = X 1 (s ,x 2 ,S) (1/2s)0/os + X 2 (s ,x 2 ,s)0/ox 2 ' Finally, to make this continuous, we multiply by s and obtain X has a singularity for (s ,x 2 ) = (0,0). To get a funnel, we want this to be a hyperbol ic attractor with real eigenvalues; this occurs if and only if oX 1\u00b0 > -(0)oZ oX1 2\u00b0 < ~(-(O))oZ This is clearly an open condition on X at 0 and hence the situation with X having a hyperbolic attractor with real eigenvalues will in general occur. In that case, the phase portrait of X is as in Figure 1 and the projection of the solutions on the (x 1 ,x 2 )-Plane looks as in Figure 2. we get a funnel. This shows clear1.y that Many more details are in a forthcoming paper. Reference. 1 \u2022 E \u2022C. Zeeman, Differential equations for the heartbeat and nerve impulse, Proc. Symp. Dynamical Systems Salvador, Academic Press, 1973, pp. 663-741 Research supported by the National Science Foundation under grant GP 29321. Address. F. Takens, Mathematisch Instituut, Postbus 800, Groningen, The Netherlands. 38. What is the Unit of Selection" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001410_app.23958-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001410_app.23958-Figure4-1.png", "caption": "Figure 4 CVs of (a) Ni-deposited PANI and (b) bare PANI matrices (matrices were prepared as in Fig. 3) in the 0.1M Ni-complex solution at 200 mV/s.", "texts": [ " The sharp rise in cathodic current at about 1.6 V and in the anodic current at about 0.3 V is shown in the voltammogram. The results seem to be identical with those depicted in Figure 1(a) suggesting that Ni\u00fe2 from the electrolytic solution was reduced to metallic Ni at the PANI substrate (PANI/Ni) at about 1.6 V, and it oxidized back in the reverse scan to the electrolytic solution around 0.3 V. The deposition of Ni onto the PANI substrate was evidenced further from its dissolution process. The result is described in Figure 4. The oxidation and, hence, dissolution of the Ni deposit from the polymer was performed by the sweeping of the potential as restricted between 0.5 and 0.3 V in 0.1M Ni-complex solution at 200 mV/s, as shown in Figure 4(a). The anodic current at about 0.3 V and higher decreased as the sweeping was repeated within the restricted potential range. The decrease in anodic current may have resulted from the dissolution of Ni that incorporated previously into the PANI substrate. This observation was absent, as shown in Figure 4(b), if the PANI substrate without Ni was examined under similar electrolytic and experimental conditions, as shown in Figure 4(a). The results depicted in Figure 4 thus suggest that Ni was successfully incorporated into the PANI matrix under the experimental conditions we used. The charge versus time (Q\u2013t) plot for the Ni dissolution process was used to get insight into the real charge that indeed was consumed for Ni deposition onto the polymer matrix. For this purpose, Ni was deposited onto the PANI film by the passing of a 50-mC charge at 1.6 V. The film was then held at a constant potential of 0.1 V, and the charge released was recorded with time. The result is shown in Figure 5" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002169_rnc.1220-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002169_rnc.1220-Figure2-1.png", "caption": "Figure 2. Illustration for collision avoidance.", "texts": [ "1002/rnc attractiveness of a future target candidate xFj;l\u00f0t\u00de to MSAj at t: Aj\u00f0x F j;l\u00f0t\u00de\u00de \u00bc g0\u00f0l\u00de PNF j \u00f0t\u00de l0\u00bc1 g1\u00f0l; l0\u00de if XC j \u00f0t\u00de=| \u00f0i:e: NC j \u00f0t\u00de > 0\u00de g0\u00f0l\u00de P s2SU j \u00f0t\u00de g2\u00f0y F j;l ; s\u00de otherwise 8< : \u00f011\u00de where g0\u00f0l\u00de \u00bc 0 if d\u00f0s; sj\u00f0t\u00de\u00de5dw and /ssj\u00f0t\u00dex F j;l\u00f0t\u00de4arcsin dmin d\u00f0s;sj\u00f0t\u00de\u00de 8s 2 Sj\u00f0t\u00de 1 otherwise 8< : \u00f012\u00de g1\u00f0l; l 0\u00de \u00bc 1 sin yFj;l0 \u00f0t\u00de yFj;l\u00f0t\u00de 2 ! \u00f013\u00de and g2\u00f0y F j;l ; s\u00de \u00bc sin yFj;l\u00f0t\u00de c\u00f0s; sj\u00f0t\u00de\u00de 2 \u00f014\u00de where c\u00f0s; sj\u00f0t\u00de\u00de is the heading of s with respect to sj\u00f0t\u00de; and dw5dmin is a user-defined distance threshold for collision warning. The attractiveness function above is designed based on the following heuristics. g0\u00f0 \u00de: A future target candidate should have zero attractiveness to an MSA if another MSA that potentially leads to a collision is in the way (Figure 2). The second half of the first row of (12) basically says that there is an MSA blocking MSAj \u2019s path to xFj;l\u00f0t\u00de: Since the position of the MSAs may change very quickly over time, an MSA that intercepts in MSAj\u2019s path to a future target candidate at current time does not necessary mean it will actually be in MSAj \u2019s way later on. Thus, a threshold dw is put in (12) to prevent the collision assessment being activated until two MSAs are close enough. g1\u00f0 \u00de: A future target candidate should be more attractive to an MSA if there are more future target candidates in a similar direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003392_9780470876541.ch5-Figure5.4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003392_9780470876541.ch5-Figure5.4-1.png", "caption": "Figure 5.4 Four-pole interior permanent magnet synchronous machine (IPMSM) and its rotor reference d\u2013q frame.", "texts": [ " The current, which interacts with the field flux, should be regulated instan- taneously regardless of the variation of back EMF, leakage inductance, and winding resistance. 2. The field flux should be regulated regardless of the variation of the current, which interacts with the field flux. 3. The flux and the current should be kept as perpendicular instantaneously either by mechanical means or by angle measurement and control. If above three conditions are fulfilled simultaneously, then the torque of the electric machine can be regulated instantaneously. The torque of the IPMSM in Fig. 5.4 can be represented as Te \u00bc 3 2 P 2 f\u00f0Ld Lq\u00deirdsirqs \u00fe lf irqsg at rotor reference d\u2013q frame derived at Section 3.3.3.2. It is the sum of the reluctance torque and the field torque. And, even if the field flux, lf , by the magnet is kept constant, the pair of irds and irqs is enormous to generate the given torque reference, T* e . However, if the total losses are minimized at the given torque, the pair is uniquely decided. If the iron loss can be neglected, theminimization of themagnitude of the stator current vector, irdqs, is the minimization of the copper loss, which is the only loss now considered" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002175_robot.2007.363146-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002175_robot.2007.363146-Figure1-1.png", "caption": "Fig. 1. Sketch of 5-RRR(RR)", "texts": [ " Kinematic screw for a prismatic pair is a couple vector whose h=\u221e, $prismatic=[0T; ST]T (3) where $prismatic denotes a kinematic screw for a prismatic pair; 0 is a 3\u00d71 zero-vector. Two screws ($A and $B) are reciprocal to each other if they satisfy[15;16] SA \u00b7 S0B + S0A\u00b7SB =(lA\u00b7pB+mA\u00b7qB+nA\u00b7rB)+(pA\u00b7lB+qA\u00b7mB+rA\u00b7nB) = 0 (4) where $A=[SA T;S0A T]=[lA mA nA ; pA qA rA]T, $B=[SB T;S0B T]T = [lB mB nB ; pB qB rB ]T. According to Eq. (4), two coplanar linear vectors $A and $B are reciprocal to each other. As shown in Fig. 1, a movable platform (end-effector) and base are connected by five identical limbs each with five joints Ri, i=1,2,3,4,5. Axes of R1, R2, R3 are perpendicular to the base plane and the parallel structure is denoted with underline RRR. The other two joints, R4 and R5, intersect at a common point called rotation center. The intersection structure is denoted with parentheses (RR). Five R1 are chosen as actuators. The kinematical screw system[15;16] for one RRR(RR) limb consists of five joint screws ($i, i=1,2,3,4,5) corresponding to five joints, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001872_978-1-4020-6366-4_15-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001872_978-1-4020-6366-4_15-Figure4-1.png", "caption": "Fig. 4. The pantograph mechanism.", "texts": [ " As brilliant as the conception of the parallel motion linkage was, it was followed up by a synthesis that is very little short of incredible. In order to make the linkage attached to the beam of his engines more compact, Watt plumbed the depths of his experience for ideas. This experience yielded up the work that was completed much earlier on a drafting machine that made use of a pantograph. Watt combined his straight-line linkage with a pantograph, one link becoming a member of the pantograph. This pantograph mechanism [16], denoted as ABEG, is shown in Figure 4. With this design, the length of each oscillating link of the straight-line linkage was reduced to one-fourth instead of one-half the beam length. The entire mechanism could then be constructed so that it would not extend beyond the end of the working beam. This arrangement soon came to be known as Watt\u2019s parallel motion linkage, denoted as O2ABO4 in Figure 5. Through insight we can detect in this straight-line linkage the birth of a very ordered and advanced synthetic process. The kinematic analysis of the Watt four-bar linkage, see Figure 6a, and the geometry of the path of point M fixed in the coupler link AB (link 3) can be investigated using the method of kinematic coefficients [17]", " (22a) Substituting Equation (17b) into Equations (22), the Cartesian coordinates of the center of the curvature of the path traced by coupler point M can be written as Xcc = XM + \u03c1M [\u2212Y \u2032 M R\u2032 M ] (23a) and Ycc = YM + \u03c1M [ X\u2032 M R\u2032 M ] . (23b) In general, an arbitrary coupler point of a general four-bar linkage will trace a curve which is described as a tricircular sextic [18\u201320]. However, coupler point M of the Watt four-bar linkage traces a special curve which is best described as a figure-eight-shaped curve, as shown in Figure 8. This curve is commonly referred to as a lemniscate and has two segments that approximate straight lines [21]. By means of the pantograph mechanism (see Figure 4), the path traced by point M \u2032 (see Figure 5) is similar to the path traced by coupler point M. The Watt four-bar linkage was employed some 75 years after its inception by Richards when, in 1861, he designed his first high-speed engine indicator. The Richards indicator, which was introduced into England the following year, was an immediate success, and many thousands were sold over the next several decades. In considering the order of synthetic ability required to design the straight-line linkage and to combine it with a pantograph, it should be kept in mind that this was the first one of a long line of such mechanisms" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002542_00029890.2009.11920919-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002542_00029890.2009.11920919-Figure5-1.png", "caption": "Figure 5. Angle of inclination of a trammel is also the eccentric angle of traced ellipse.", "texts": [ " Therefore AM = MB = MO = L/2, which shows that AB has fixed length L , so it is a standard trammel, and we have already seen that it is always tangent to the astroid. This property is illustrated in Figure 7a, which shows various positions of the trammel as it slides along the coordinate axes in all four quadrants. It is described by saying that: The envelope of a moving trammel of fixed length is an astroid. For calculus derivations of this envelope see [5] and [6]. 4. EQUATIONS FOR ELLIPSE, TRAMMEL, AND ASTROID. Figure 5 shows an ellipse with semiaxes a and b and two concentric circles with radii a and b. A line February 2009] THE TRAMMEL OF ARCHIMEDES 117 from the origin making angle \u03b1 with the x axis is related to the coordinates of a point E = (x, y) on the ellipse by the parametric equations x = a cos \u03b1, y = b sin \u03b1. (3) Angle \u03b1 in (3) is called the eccentric angle of the ellipse (see [1, p. 522]). Figure 5 also shows a line segment AB through E making the same angle \u03b1 with the x axis. It is easy to show that AB has length a + b. Therefore it is a trammel whose angle of inclination is the eccentric angle. We simply note that AE cos \u03b1 = x , so according to (3), AE = a. Similarly, BE sin \u03b1 = y, giving BE = b. In Figure 6a, a line AB of length L makes an angle \u03b1 with the x axis and is tangent to the astroid at point P with coordinates (x, y). First we note that every point on the trammel satisfies the linear Cartesian equation x L cos \u03b1 + y L sin \u03b1 = 1, (4) because L cos \u03b1 and L sin \u03b1 are the x and y intercepts of the trammel", " Similarly we find PB = L sin2 \u03b1. Using these in (5) we see that each point (x, y) on the astroid satisfies the parametric equations x = L cos3 \u03b1, y = L sin3 \u03b1. (6) Using (6) together with cos2 \u03b1 + sin2 \u03b1 = 1 we obtain the following Cartesian equation of the astroid: ( x L )2/3 + ( y L )2/3 = 1. (7) Figure 6a point P with coordinates (5) is on the trammel (4) and on the astroid (7). The trammel is always tangent to the astroid but in general it intersects the ellipse described by (3) at two points as in Figure 5. As the trammel moves, these two points will coincide for some critical angle \u03b1, say \u03b1 = \u03b8 as shown in Figure 6b, and in this position the trammel is simultaneously tangent to the ellipse and the astroid at some point E with coordinates (a cos \u03b8, b sin \u03b8). The following theorem determines \u03b8 in terms of the semiaxes a and b of the ellipse. Theorem 1. The trammel is simultaneously tangent to the ellipse and the astroid when its angle of inclination \u03b8 with the x axis satisfies tan \u03b8 = \u221a b a . (8) Proof" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002100_jmes_jour_1969_011_071_02-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002100_jmes_jour_1969_011_071_02-Figure8-1.png", "caption": "Fig. 8. Conjugate centre distances for pu2 > p a ,", "texts": [ " (20) the discrepancy in two domains for centre distances vanishes. Afterwards it is easy to conclude that the method of Fig. 7 is also operative for a > ( r h 2 - r h , ) . ry2 sin (vh+4h) = ral sin #1 a, = rU2 cos (&2+A$2)-r( l , cos $l . (22) A+2 = - 1nv %I1 1 . (21) . t j . (23) CONJUGATION O F CENTRE DISTANCES z2 - 21. '2 i REPRESENTED FOR p.2 > p a l (ra2-rul) < 01 < (ru2-rbl) J ' (24) The representation of the conjugate centre distance in Without detriment to general validity a simplification Fig. 8 confirms statement (12). The point a, = uI I = , / [ ( f a 2 - P u 1 ) 2 + ( t b 2 - T b 1 ) 2 ] relates to the case of Fig. 5 rv2 = ra2, au2 = c i a 2 . . . (25) where the tooth tips touch on the contact line. In all other cases a,, < a, indicates non-interference. may be introduced BY 4, = $1 7 . . (26) 2 2 - 2 1 , > 4 2 = h+- inv u~~~~ I . . (27) z2 CONJUGATION OF CENTRE DISTANCES REPRESENTED FOR poz < pol the equations (20)-(23),lead to equation (15), by which The representation of the conjugate centre distance in J O U R N A L M E C H A N I C A L E N G I N E E R I N G S C I E N C E Vol I1 No 6 I969 at NANYANG TECH UNIV LIBRARY on June 5, 2016jms" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003437_9781118516072.ch4-Figure4.15-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003437_9781118516072.ch4-Figure4.15-1.png", "caption": "Figure 4.15. Drive trains of (a) conventional wind turbine; (b) direct-drive generator.", "texts": [ " Large wind turbines were classically designed with generators with small number of poles and high speeds, up to 1500 rpm at 50Hz and 1800 rpm at 60Hz. As the turbine speed is much lower than the generator speed, typically between 20 and 60 rpm, a gearbox was required between the turbine and generator to adapt the speed. In order to eliminate the problems introduced by the gearbox (e.g., high losses and high noise), nowadays there is a trend to use low-speed generators, which can be directly connected to the turbine shaft. Figure 4.15 shows comparatively the drive trains of a conventional wind turbine and one with a direct-drive generator [11]. Direct connection of the generator to the turbine involves a very high torque developed at the rotor shaft. For instance, a 500 kW direct-drive generator, with 30 rpm, has the same rated torque as a 50MW steam turbine generator, but with 3000 rpm [5]. It is important to know that the size and the losses of a low-speed generator depend on the rated torque rather than on the rated power" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001604_robot.1987.1087799-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001604_robot.1987.1087799-Figure8-1.png", "caption": "Figure 8", "texts": [ " c Remark 1 The collision-free joint space with respect to any obstacle can never be of the form shown in Figure 7. To see why this is so, consider the case (a) in Figure 7. Let (811: 821) and ( 8 1 2 , B z z ) be two points on the boundary of the J-obstacle as indicated in the figure, and let w1 = {w(r,q1,qz) : 91 = h q 2 E (821,82),0 5 r 5121, wz = { w ( w , ~ ~ ) : q1 = o 1 2 , q 2 E (e22,82) ,o 5 5 12), and where I , is the length of the ith link. Clearly any posture in the set W1 or the set Wz is achievable. But it can be seen from Figure 8 tha t some postures in the set W1 (i.e. those in the shaded area) are unachievable when q1 = 812 - a contradiction. Hence, the collision-free joint space cannot be of the given form. Similarly, it can be shown that the remaining cases are not possible. Remark 2 It is interesting to note that the configurations at which cases (c)-(f) occur are the local maxima or local minima of the curve ~ A J plotted on the 91- q z plane (see Figure 9). The rest of the curve corresponds to the cases (a) and (b)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000985_ichr.2004.1442686-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000985_ichr.2004.1442686-Figure7-1.png", "caption": "Fig. 7. Reliability based on change in range information taken from an ultrasonic range sensor", "texts": [ " Ultrasonic range sensors The robot can obtain omni-directional range information from the ring-type ultrasonic range sensors. We calculate its reliability distribution from changes in the range information, which describe the existence of moving objects. The distribution also conforms to a normal distribution. Its mean and the variance are 0.2 x &in D ' a, = P7k U07 and 0,' = D, where D, D,,,, and U, denote the distance between the object and the robot, the minimum distance that can be measured by the sensor, and its direction. The variables and the distribution are shown in Fig. 7. While the robot is moving, au is 0. Pyroelectn'c infrared Sensors The pyroelectric infra-red sensors are used to detect the moving objects by measuring changes in inka-red readings. The sensors' output is either l (detected) or 0 (not detected). We define these sensors\u2019 reliability distribution, fp(z), as 0.3 x (-30 5 xP 5 30), f p ( z ) { 0 (otherwise), (15) where ap is the weight value 0 whiIe the robot is moving and 1 while it is not moving. The distribution is shown in Fig. 8. After calculating all reliability distributions for all sensors, we combine the distributions into a reliability distribution of human existence around the robot by the following equation" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000523_1.1850942-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000523_1.1850942-Figure2-1.png", "caption": "Fig. 2 Location of the bearing centers with respect to the deformed rotor centreline", "texts": [ " Equation 4 then simplifies to Kr xro f xro c T U V W xro f xro c 0 fbo c fe f fe c (23) and on eliminating the nonconnection rotor displacements W VT 1U xro c Kr cxro c fbo c fe c VT 1fe f (24) Similarly, Eq. 11 simplifies to KP xpo f xpo c P Q R S xpo f xpo c 0 fpo c (25) and on eliminating the nonconnection pedestal displacements S RP 1Q xpo c Kp c xpo c Kp c b fpo c (26) Also in conformity with Eq. 13 fpo c fbo c (27) The configuration state identification problem is illustrated for a four-bearing system in Fig. 2 where the elements of xro corresponding to the lateral displacements of the deformed rotor centerline ri can be measured with respect to any arbitrary fixed datum line, here selected to be the line through the centers of the first and last unloaded bearing housings. The locations of the centers of the other unloaded bearing housings relative to this datum line define the configuration state of the system (c2 and c3 in Fig. 2 . If the mean eccentricities ei of the rotor journals define the locations of the centers of the bearing housings relative to the deformed rotor centreline and the mean displacements bi define the displacements of the centers of the bearing housings from their original unloaded positions, then one can see from Fig. 2 that at the bearing stations, ri ci bi ei i 1,2,3,4 (28) Hence, if we restrict the connection DOFs to the two lateral directions at each bearing, then Eq. 24 involves eight simultaneous equations in the eight unknown bearing force components in fbo and the four unknown configuration components in c2 and c3 since the ei are measured and the bi can be expressed in terms of the fbo by virtue of Eqs. 26 and 27 . Thus, one has an excess of unknowns. Taking measurements at another speed would yield eight additional equations, but there would also result eight additional unknown values of fbo ", " 14 b vector of the mean displacement of the bearing housings bi mean displacement of ith bearing housing C, C damping and gyroscopic matrix, element therein ci location of ith unloaded bearing housing ei mean journal eccentricity at ith bearing F force vector f, f force frequency component vector, element therein H dynamic stiffness matrix I, J, L, N partition matrices defined by Eq. 16 K, K stiffness matrix, element therein k order of harmonic M, M mass matrix, element therein n highest order harmonic P, Q, R, S partition matrices defined by Eq. 25 q number of rotor connection degrees of freedom ri location of rotor centreline at ith bearing sta- tion as measured from datum line in Fig. 2 s harmonic order of synchronous component T, U, V, W partition matrices defined by Eq. 23 t time X, X displacement vector, element therein x, x displacement frequency component vector, ele- ment therein rotor speed fundamental frequency of steady state solution Transactions of the ASME 016 Terms of Use: http://www.asme.org/about-asme/terms-of-use Download \u2022 differentiation with respect to time t c connection degree of freedom f free nonconnection degree of freedom Subscripts \u201eif not otherwise defined\u2026 b bearing e external o zero-order harmonic or mean value p foundation or pedestal r rotor u unbalance y, z horizontal and vertical directions, respectively 1 Lund, J" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000659_j.jmatprotec.2004.09.006-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000659_j.jmatprotec.2004.09.006-Figure2-1.png", "caption": "Fig. 2. Rigid body small displacement and screw parameter.", "texts": [ " Referring to Table 1, the MGRE of the prismatic class is ormed by a hierarchical set of a plane and a straight line. The absorption law states that the plane is identified by he cylinders axes (primary datum, PL in Fig. 1) and by the xis of cylinder S1 (secondary datum, MGRE 1 in Fig. 1). n this reference frame (the reference frame of TTRS1,2,) the olerance zone of the MGRE 2 of the toleranced cylinder S2 ill be positioned. mal translation T along this axis and an infinitesimal rotation around this axis. Let (Fig. 2) O a point in the Euclidean space. The set \u03c4, amed \u201cscrew parameter\u201d, is composed by the kinetic charcteristics \u2126 (infinitesimal rotation) and DO (infinitesimal isplacement) in O. O = ( \u2126 T+ \u2192 OC \u00d7 \u2126 ) O = ( \u2126 DO ) (4) By virtue of the small rigid body displacement hypothesis, is invariant with respect to the point considered, whereas O depends on it. Let M a further point in the Euclidean Table 3 TTRS reclassifications Complex Prismatic Revolution Helicoidal Cylindrical Plane Spherical Complex Complex Complex Complex Complex Complex Complex Complex Prismatic Complex prismatic Complex Complex Complex prismatic Complex prismatic Complex Revolution Complex revolution Complex Complex revolution Complex revolution Complex revolution" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001836_978-3-540-71967-0_2-Figure2.4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001836_978-3-540-71967-0_2-Figure2.4-1.png", "caption": "Fig. 2.4. MD 900 helicopter hingeless blade displaying the planned trim tab for in-flight tracking and active control flap for noise and vibration reduction [10]", "texts": [ " The flaps showed excellent authority with oscillatory thrust greater than 10% of the steady baseline thrust. Various flap actuation frequency sweeps were run to investigate the dynamics of the rotor and the flap system. Limited closed loop tests used hub accelerations and hub loads for feedback. Proving the integration, robust operation, and authority of the flap system were the key objectives met by the whirl tower test. This success depended on tailoring the piezoelectric materials and actuator to the application and meeting actuator/blade integration requirements (Fig. 2.4). Induced-strain Actuation of Fixed-Wing Aircraft. The feasibility of using active piezoelectric control to alleviate vertical tail buffeting was inves- tigated under the actively controlled response of buffet affected tails (ACROBAT) program [11]. Tail buffeting is a significant concern from fatigue and maintenance standpoints. During the ACROBAT program, active materials solutions to buffet problems were studied on 1/6-scale rigid full-span model of the F/A-18 aircraft tested in the Langley transonic dynamics tunnel (TDT)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002685_rnc.1378-Figure13-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002685_rnc.1378-Figure13-1.png", "caption": "Figure 13. The implicit function method.", "texts": [], "surrounding_texts": [ "Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 Published online 9 September 2008 in Wiley InterScience (www.interscience.wiley.com). DOI: 10.1002/rnc.1378 A homotopy method for exact output tracking of some non-minimum phase nonlinear control systems L. Consolini1,\u2217,\u2020 and M. Tosques2 1Dipartimento di Ingegneria dell\u2019informazione, Viale Usberti 181/A, 43100 Parma, Italy 2Dipartimento di Ingegneria Civile, Viale Usberti 181/A, 43100 Parma, Italy SUMMARY This paper presents a method for non-causal exact dynamic inversion for a class of non-minimum phase nonlinear systems, which seems to be an alternative to those existing in the literature. This method is based on a homotopy procedure that allows to find a \u2018small\u2019 periodic solution of a desired equation by a continuous deformation of a known periodic solution of a simpler auxiliary system. This method allows to face the exact output tracking problem for some non-minimum phase systems that are well known in the literature, such as the inverted pendulum, the motorcycle and the CTOL aircraft. Copyright q 2008 John Wiley & Sons, Ltd. Received 20 April 2007; Revised 19 June 2008; Accepted 23 July 2008 KEY WORDS: nonlinear non-minimum phase systems; output tracking; dynamic inversion 1. INTRODUCTION In this paper we consider the problem of exact output tracking a periodic reference trajectory for a suitable family of non-minimum phase nonlinear control systems, which includes some well-known benchmarks for nonlinear control theory such as the inverted pendulum [1, 2], the VTOL aircraft [3\u20136], the motorcycle [7] and the CTOL aircraft [8, 9]. Solving the exact output tracking problem for one of these systems, the resulting unstable internal dynamic has the form \u0308= f (t, ) (1) our goal is to find an initial condition for Equation (1), such that the resulting solution is periodic and sufficiently small. To this end we apply a homotopy method that relies essentially on the following steps. We find a family of differential problems dependent on s \u0308= f (t,s, ) such that f (t,1, )= f (t, ) and \u0308= f (t,0, ) admits a \u2018small\u2019 periodic solution \u0303. In Theorem 1, we \u2217Correspondence to: L. Consolini, Dipartimento di Ingegneria dell\u2019Informazione, Viale Usberti 181/A, Parma, Italy. \u2020E-mail: luca.consolini@polirone.mn.it Copyright q 2008 John Wiley & Sons, Ltd. present a sufficient condition that implies that there exists an s\u0304>1 with the property that \u2200s\u2208[0, s\u0304[ (therefore also for s=1) there exists a periodic solution s of \u0308= f (t,s, ) such that \u2016 s\u2212 \u0303\u2016\u221e and \u2016\u0307s\u2212 \u0307\u2016\u221e are small. This method represents a procedure for non-causal dynamic inversion, which is an alternative to the method presented in [10] and has the advantage of not requiring the linearized internal dynamics to be \u2018slowly varying\u2019. Homotopy (or continuation) methods have been used in the literature for solving a difficult nonlinear problem by means of a continuous deformation of it into an easier one (see [11] for a survey on the topic). These methods have received considerable attention in nonlinear control theory as well: for instance, in [12] for the feedback stabilization of linear systems, in [13] for the least-squares estimation, in [14] for boundary value problems in optimal control, in [15] for finding an input signal that drives a nonlinear system to a given state. This paper is organized as follows: from Section 2 to Section 4 the exact output tracking problem for the inverted pendulum on a cart, the motorcycle, the CTOL-aircraft are presented and discussed by means of the homotopy method presented in Theorem 1, which is finally stated and proved in Section 5. We remark that an analogous problem for the VTOL-aircraft has been considered in [16]. The following notations will be used: \u2200a,b\u2208R, a\u2227b=min{a, b}, a\u2228b=max{a, b} and [a,b]= {x \u2208R|a x b}, ]a,b[={x \u2208R|ask} then it is easy to see that there exists \u0304 0 and >1 : [0, ]\u00d7[0, \u0304]\u2282R therefore (13) of Theorem 1 is verified with f1 and f2 given above. Since s f (t,s, )= \u2329 \u0308s(t) , ( cos sin )\u232a if we set \u2016\u0308\u2016\u221e kg, we can set (s, )=kg which implies, by (14), that (s, )= (1+sk)k cos \u2212sk By numerical computation, we see that if 0 k 0.2672 the maximal interval of existence [0, s\u0304[ of the solution of system (16) is such that s\u0304>1 and we found the values (1) in terms of the values of k as shown in Figure 2. In particular if k=0.2672, it is (1)=1.4302. Therefore by Theorem 1, applied with s=1, the following result holds. Result 1 For every periodic curve of class C2, of period T , such that \u2016\u0308\u2016\u221e 0.2672g, there exists an initial condition ( 0, \u03070)T such that system (3) has a periodic solution verifying the properties: \u2200t \u2208R | (t)| (1)=1.4302 |\u0307(t)| ( (1),1) A simulation example. We apply this method to a periodic curve given by a fifth-order spline of period T =5s. This function has a continuous third-order derivative. The spline satisfies the properties (0) = (0,0), \u0307(0)=(0,\u22121), \u0308(0)=(\u22121,0), \u0308(0)=(0,0) (1) = (\u22122\u22122), (2)=(1,\u22121), (3)=(2,0) (4) = (\u22122,2) (5) and di (5)/dti =di (0)/dti , for i=0, . . . ,3 and is represented in Figure 3, it is \u2016\u0308\u2016\u221e =2.2m/s2. For numerical computation of use, for instance, Matlab Spline Toolbox. Assume d=1, g=9.8m/s2. Applying the method outlined above we can find a control force that drives the pendulum along the spline without overturning. In particular the initial condition for system (3) is given by y(1)=( 0, \u03070)T=(0.0283,0.0030)T where y(s) is the solution of Equation (19) and the couple ( , \u0307)T, associated with the corresponding solution of system (3), is shown in Figure 3. The numerical computation shows that | (t)| 0.085, |\u0307(t)| 0.24, in fact Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc the pendulum rod remains almost vertical (see Figure 1). Remark that the bounds (18) given by Theorem 1 are | (t)| 0.4861, |\u0307(t)| 1.1727. In this example, these bounds may appear very large, but this is justified by the fact that they must apply to any trajectory whose acceleration is bounded by 2.2m/s2 which is \u2016\u0308\u2016\u221e (Figures 4 and 5). 3. THE EXACT OUTPUT TRACKING PROBLEM FOR THE NON-HOLONOMIC MOTORCYCLE MODEL Consider the simple non-holonomic motorcycle model presented in [7]. The model kinematics are given by x\u0307 = v cos y\u0307 = v sin \u0307 = v (6) where (x, y) is the position of the motorcycle, is its leading angle with controls given by the speed v and the curvature (see Figure 6). Let \u2208C3(R,R2) be a periodic curve of period T such that \u2016\u0307(t)\u2016>0, \u2200t \u2208R; in order to track exactly this periodic trajectory the control must be v=\u2016\u0307\u2016 and =\u2016\u0307\u2016\u22121 d arg(\u0307) dt (7) Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc The problem is to show that there exists a value 0 of the roll angle such that starting with that value, the motorcycle does not overturn when it follows , with the controls given by (7). As shown in [7], the roll angle has to satisfy this equation: \u0308=h\u22121 [ g sin + ( h ( d arg(\u0307) dt )2 cos ) sin \u2212 ( \u2016\u0307\u2016d arg(\u0307) dt +b d2 arg(\u0307) dt2 ) cos ] (8) where g is the gravity acceleration, h is the height of the center of mass (when the vehicle is vertical), b is the distance of projection of the center of mass to the ground from the contact point of the real wheel. Equation (8) represents the system unstable internal dynamics. By the use of Theorem 1, we will show that there are suitable conditions for (see (9)), which guarantee the Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc existence of initial conditions ( 0, \u03070)T such that the solution of (8) is periodic of period-T and the angle remains small. In fact, set f (t,s, )=h\u22121 [ g sin + [ h ( d arg(\u0307) dt )2 cos ] sin \u2212s ( \u2016\u0307\u2016d arg(\u0307) dt +b d2 arg(\u0307) dt2 ) cos ] Remark that Equation (17) becomes (8) when s=1. Clearly it verifies hypotheses (1) and (2) of Theorem 1, taking as \u0303 the function identically zero. Since f (t,s, )=h\u22121 [ g cos +h ( d arg(\u0307) dt )2 cos(2 )+s ( \u2016\u0307\u2016d arg(\u0307) dt +b d2 arg(\u0307) dt2 ) sin ] we get that \u2200t \u2208R, \u2200 \u2208[0, /2[, \u2200 : | | , h\u22121 [ g cos +h \u2225\u2225\u2225\u2225d arg(\u0307)dt \u2225\u2225\u2225\u2225 2 \u221e (cos(2 )\u22270)\u2212s \u2225\u2225\u2225\u2225\u2016\u0307\u2016d arg(\u0307)dt +b d2 arg(\u0307) dt2 \u2225\u2225\u2225\u2225\u221e sin ] f (t, ,s) h\u22121 [ g+h \u2225\u2225\u2225\u2225d arg(\u0307)dt \u2225\u2225\u2225\u2225 2 \u221e +s \u2225\u2225\u2225\u2225\u2016\u0307\u2016d arg(\u0307)dt +b d2 arg(\u0307) dt2 \u2225\u2225\u2225\u2225\u221e sin ] Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc Therefore if \u2225\u2225\u2225\u2225\u2016\u0307\u2016d arg(\u0307)dt +b d2 arg(\u0307) dt2 \u2225\u2225\u2225\u2225\u221e k1g, h \u2225\u2225\u2225\u2225d arg(\u0307)dt \u2225\u2225\u2225\u2225 2 \u221e k2g (9) where k1 and k2 are two non-negative parameters, we have that f1(s, ) f (t,s, ) f2(s, ), \u2200t \u2208R, \u2200s\u2208R, \u2200 \u2208 [ 0, 2 [ , \u2200 : | | where f1(s, )=h\u22121g [cos +k2(cos(2 )\u22270)\u2212sk1 sin ] , f2(s, )=h\u22121g [1+k2+sk1 sin ] Set R={(s, ) \u2223\u2223cos +k2(cos(2 )\u22270)>sk1 sin }, clearly R is an open set of R2 such that \u2203 \u0304>0, >1 : [0, ]\u00d7[0, \u0304]\u2282R therefore condition (13) is verified. Since s f (t,s, )=h\u22121 ( \u2016\u0307\u2016d arg(\u0307) dt +b d2 arg(\u0307) dt2 ) cos we can take , in Theorem 1, in the following way: (s, )=k1g therefore (s, )=k1 (cos +k2(cos(2 )\u22270)\u2212sk1 sin )2 (1+k2+s sin k1)3 By numerical computation, we found the value k1,k2 for which the maximal interval of existence [0, s\u0304[ of the solution of system (16) is such that s\u0304>1. Figure 7 shows how (1) varies with respect to k1 for different values of k2. Remark that k2 is in general much smaller than k1 (in fact, setting k2=\u2016(\u2016\u0307\u20162/r)h/r\u2016\u221e, where r(t) is the trajectory radius of curvature, k1>\u2016\u2016\u0307\u20162/r\u2016\u221e, therefore k2/k1 h/\u2016r\u2016\u221e which is in general much smaller than 1). In particular if k1=0.33 and k2=0.1, it is s\u0304>1, therefore by Theorem 1, applied with s=1, we can deduce the following result Result 2 For every periodic curve \u2208C3 of periodic T such that \u2225\u2225\u2225\u2225\u2016\u0307\u2016d arg(\u0307)dt +b d2 arg(\u0307) dt2 \u2225\u2225\u2225\u2225\u221e (0.33)g h \u2225\u2225\u2225\u2225d arg(\u0307)dt \u2225\u2225\u2225\u2225 2 \u221e (0.1)g Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc there exists an initial condition ( 0, \u03070) such that Equation (8) has a periodic solution verifying the properties, \u2200t \u2208R: | (t)| (1)=0.45, |\u0307(t)| (1, (1)) A simulation example. Assume h=0.2, b=1m. Consider a periodic trajectory (see Figure 8), given by (t)= ( 0.5 sin( t), 0.5 sin ( 2 t ))T Applying the method outlined above we can find that the controls given by (7) will drive the motorcycle along the curve without overturning. In fact, the internal dynamic has the solution shown in Figure 9, associated with the initial condition given by y(1) where y(s) is the solution of (19). 4. THE EXACT OUTPUT TRACKING PROBLEM FOR THE CTOL AIRCRAFT Consider the simplified CTOL dynamic equation introduced in [8]:( x\u0308 y\u0308 ) = R( ) (\u2212D L ) +R( ) ( u1 \u2212 u2 ) + ( 0 1 ) \u0308 = u2 where (x, y) is the aircraft center of mass, is the pitch angle, u1 and u2 are the controls, R( )= ( cos \u2212sin sin cos ) Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc is the rotation matrix, is the flight path angle, D and L are the drag and lift forces given by L=aLv2(1+c ), D=aDv2(1+b(1+c )2) where v= \u2225\u2225\u2225\u2225\u2225 ( x\u0307 y\u0307 )\u2225\u2225\u2225\u2225\u2225 Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc is the aircraft speed, = \u2212 is the angle of attack and aL , aD , c, b and are given constants (see Figure 10). Given a curve \u2208C2(R,R2), if ( x(0) y(0) ) = (0), the controls that allow to exactly track are given by ( u1 \u2212 u2 ) = R(\u2212 ) [ \u0308+ ( 0 1 )] +R(arg(\u0307)\u2212 ) (\u2212D L ) where L = aL\u2016\u0307\u20162(1+c( \u2212arg(\u0307))) D = aD\u2016\u0307\u20162(1+b(1+c( \u2212arg(\u0307))2)) and is the solution of the following equation: \u0308=\u2212 \u22121 [\u2329(\u2212sin cos ) , \u0308+ ( 0 1 )\u232a + \u2329(\u2212sin( \u2212arg(\u0307)) cos( \u2212arg(\u0307)) ) , ( D \u2212L )\u232a] (10) which represents the unstable internal dynamics. Our problem is to show that there exists an initial condition ( 0, \u03070) such that the solution of the internal dynamics (10) remains small. To this end we apply the homotopy method given by Theorem 1. Remark that if v0 is such that 00, >0 and two locally Lipschitz functions, for i=1,2 : fi : [0, [\u00d7[0, \u0304[ \u2192 R (s, ) fi (s, ) such that 0< f1(s, ) f (t,s, ) f2(s, ) \u2200t \u2208R, \u2200 \u2208[0, \u0304[ \u2200 : | \u2212 \u0303(t)| , \u2200s\u2208[0, [ (13) Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc Let it be, \u2200(s, )\u2208[0, [\u00d7[0, \u0304[: (s, )= f 22 (s, ) f 31 (s, ) (s, ) (14) where is a locally lipschitz function such that \u2200(s, )\u2208[0, [\u00d7[0, \u0304[: (s, ) max{| s f (t,s, )| :0 t T, | \u2212 \u0303(t)| } (15) and let [0, s\u0304[ be the maximal interval of existence of the solution of the system \u0307(s) = (s, (s)) \u2200s\u2208[0, s\u0304[ (0) = 0 (16) Then there exists a family { s}0 s0, 0, 0 such that \u2200 \u2208]\u2212 0, T + 0[, \u2200y\u2208 B(y0( ), 0), \u2200s : |s|< 0, the maximal interval of existence of the solution of (23) contains properly [ ,T + ], in other words it well defines the map P( ,s, y)= x(T + , ,s, y) (26) on the open set ={( ,s, y)|\u2212 0< 0 and a C1 map y : [0,T ]\u00d7[0, ]\u2192R2, ( ,s) y( ,s) such that y( ,0)= y0( ) \u2200 \u2208 [0,T ], P( ,s, y( ,s))=0 \u2200( ,s)\u2208[0,T ]\u00d7[0, ] {( ,s, y( ,s))|( ,s) \u2208 [0,T ]\u00d7[0, ]}={( ,s, y)\u2208S |P( ,s, y)=0)} (30) where S ={( ,s, y)|0 T, s\u2208[0, ],\u2016y\u2212 y0( )\u2016 } is the cylinder of radius constructed aroundS0 (that is \u2200y( ,s)\u2208[0,T ]\u00d7[0, ], y( ,s) is the only solution of the equationP( ,s, y)=0, inside the cylinder S ), yP( ,s, y( ,s)) is invertible \u2200( ,s)\u2208[0,T ]\u00d7[0, ] (31) Let \u0304 be the supremum of \u2019s such that there exists a unique y\u2208C1([0,T ]\u00d7[0, ], R2) verifying the properties (30), (31); in other words we can say that there exists a C1 map y : [0,T ]\u00d7[0, \u0304[\u2192R2 such that (30), (31) hold with [0, ] substituted by [0, \u0304[. Let [0, s\u0304[ be the maximal interval of existence of the solution of system (16). Since the function (s, ) given by (14) is defined on [0, [\u00d7[0, \u0304[, it must be s\u0304 and \u0304=sup0 s0), to a function which has the properties (30) and (31), contradicting the definition of \u0304. Therefore \u0304 s\u0304 and if we set s(t)= y1(t,s) \u2200t \u2208R, \u2200s :0 s s\u0304 { s}0 s0, x2=\u221a f1x1} and F2={x \u2208R2|x1>0, x2=\u221a f2x2} Suppose that there exists a t\u0304>0 such that x(t\u0304) /\u2208 1 (49) Then (see Figure 16) there exists a t0 and >0 such that x(t0)\u2208 1 and x(t)\u2208R2\\ 1, \u2200t \u2208 ]t0, t0+ ]. Remark that if x(t0)=0, by the uniqueness of the solution x(t)=0, \u2200t \u2208R, which Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc implies that x(t\u0304)=0\u2208 1, which contradicts (49). Suppose now that x(t0)\u2208F1, then, unless of decreasing , we can suppose \u3008x(t) , v1\u3009 > 0 \u2200t \u2208]t0, t0+ ] (50) x1(t) > 0 \u2200t \u2208]t0, t0+ ] (51) Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc where v1= (\u221a f1\u22121 ) . Since \u3008x(t0) , v1\u3009=0, being x(t0)\u2208F1 it is \u3008x(t0+ ) , v1\u3009 = \u2329 x(t0)+ \u222b t0+ t0 x\u0307(t)dt , v1 \u232a = \u222b t0+ t0 \u3008x\u0307(t) , v1\u3009dt = \u222b t0+ t0 \u2329 A(t)x(t) , (\u221a f1 \u22121 )\u232a dt= \u222b t0+ t0 (x2(t) \u221a f1\u2212x1(t) f (t))dt \u222b t0+ t0 x1(t)( f1\u2212 f (t))dt since x2(t) x1(t) f2, \u2200t \u2208]t0, t0+ ] by (50). Therefore by (45) and (51) it is \u3008x(t0+ ) , v1\u3009 0 which is impossible since \u3008x(t0+ ) , v\u3009>0 by (50). If x(t0)\u2208F2, as before unless of decreasing >0, we can suppose that there exists >0 such that \u3008x(t) , v2\u3009 > 0, \u2200t \u2208]t0, t0+ ] (52) x1(t) > 0, \u2200t \u2208]t0, t0+ ] (53) where v2= (\u2212\u221a f2 1 ) . Since \u3008x(t0) , v2\u3009=0, being x(t0)\u2208F2 it is \u3008x(t0+ ) , v2\u3009 = \u2329 x(t0)+ \u222b t0+ t0 x\u0307(t)dt , v2 \u232a = \u222b t0+ t0 \u3008x\u0307(t) , v2\u3009dt = \u222b t0+ t0 \u2329 A(t)x(t) , (\u2212\u221a f2 1 )\u232a dt= \u222b t0+ t0 (x1(t) f (t)\u2212x2(t) \u221a f2)dt \u222b t0+ t0 x1(t)( f (t)\u2212 f2)dt since x1(t) \u221a f2 x2(t), \u2200t \u2208]t0, t0+ ] by (52). Therefore by (45) and (53) it is \u3008x(t0+ ) , v2\u3009 0, which is impossible since \u3008x(t0+ ) , v2\u3009>0 by (53). Then (49) cannot be verified and 1 is positively invariant with respect to A(t). Analogously, it is proved that 2 is positively invariant with respect to the vector field \u2212A(\u2212t)x . Moreover, \u2200i=1,2, the eigenvectors wi (t) of (t) satisfy wi (t)\u2208 i , \u2200t 0. In fact, set i = arcsin \u221a fi/ \u221a 1+ fi , i=1,2 and define the continuous map Gt : [ 1, 2]\u2192R by Gt ( )=arg( (t) ( ))\u2212 where \u2200 \u2208R, ( )=(cos , sin )T. Because of the positive invariance of 1, it is Gt ( 1) 0 and Gt ( 2) 0, therefore by Bolzano\u2019s theorem there exists \u0304\u2208[ 1, 2] such that Gt (\u0304)=0 that is there exists 1(t) such that (t) (\u0304(t))= 1(t) (\u0304(t)) therefore w1(t)= (\u0304(t)) is a normalized eigenvector. Analogously, using the positive invariance of 2 with respect to \u2212A(\u2212t) it is verified that the other eigenvector w2(t) of (t) belongs to 2. Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc Given any x\u0304 \u2208 1\\{0}, let it be x(t)= (t)x\u0304 , \u2200t 0. Since x(t)\u2208 1, by the positive invariance of 1, there exists function (t) and (t) such that x(t)= (t) ( 1 (t) ) , (t) 0, \u221a f1 (t) \u221a f2 \u2200t 0 Since \u2016x(t)\u2016= (t) \u221a 1+ 2(t), \u2200t with \u2016x(t)\u2016>0 it is d\u2016x(t)\u2016 dt = \u3008x\u0307(t) , x(t)\u3009 \u2016x(t)\u2016 = \u3008A(t)x(t) , x(t)\u3009 (t) \u221a 1+ 2(t) = (t) (t)\u221a 1+ 2(t) (1+ f (t))=\u2016x(t)\u2016 (t) 1+ (t)2 (1+ f (t)) Since \u221a f1 (t) \u221a f2(t), \u2200t 0 it is (t) 1+ 2(t) \u221a f1 1+ f1 \u2227 \u221a f2 1+ f2 and d\u2016x(t)\u2016 dt ( \u221a f1 1+ f1 \u2227 ( \u221a f2 1+ f2 (1+ f1) )) \u2016x(t)\u2016= (\u221a f1\u2227 (\u221a f2 1+ f1 1+ f2 )) \u2016x\u2016= 0( f1, f2)\u2016x\u2016 therefore \u2016x(t)\u2016 e 0t\u2016x(0)\u2016. Finally let x(t)= (t)w1(t), where w1(t) is the eigenvector of (t) belonging to 1. Being w1(t)\u2208 1, \u2200t 0, it is, by the previous reasoning, \u2016x(t)\u2016 e 0t\u2016w1(t)\u2016 which implies, since x(t)= 1(t)w(t), \u2200t 0 that 1(t) e 0t \u2200t 0 Therefore (48) holds since 2(t)= 1(t) \u22121 being the trace of A(t)=0. The following lemma gives an estimate on the fixed points of the solution for a non-homogeneous time-dependent linear system associated with a hyperbolic matrix A(t) having the form (44) with controls (45) on the coefficients. Lemma 2 Let A(t) be a continuous 2\u00d72 matrix defined on [0,+\u221e[, satisfying the hypotheses of Lemma 1 and let B(t)= ( 0 d(t) ) be a continuous vector on [0,+\u221e[. Let \u2208R2 be such that there exists 0 and T>0 for which = (T + , ) + \u222b T+ (T + , p)B(p)dp Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc where (t, ) is a solution of (46). Then | 1| c1 1+e\u2212 0T 0 \u2016B\u2016[ ,T+ ] | 2| c2 1+e\u2212 0T 0 \u2016B\u2016[ , T+ ] (54) where 0( f1, f2) = \u221a f1\u2227 (\u221a f2 1+ f1 1+ f2 ) c1( f1, f2) = \u221a f2 f1 1+ f2 1+ f1 c2( f1, f2) = f2 f1 \u221a 1+ f2\u221a 1+ f1 (55) and \u2016B\u2016[ ,T+ ] =max{\u2016B(t)\u2016| t T + }. Proof Without loss of generality, assume =0, therefore suppose that satisfies = (T,0) + \u222b T 0 (T, )B( )d (56) Set W (t, )=(w1(t, ),w2(t, )) where wi (t, ) are the normalized eigenvectors and i (t) the eigenvalues of (t, ). Remark that for any invertible 2\u00d72 matrix W =(w1,w2) it is \u2200z\u2208R2, z=(W\u22121z)1w1+(W\u22121z)2w2 (57) where (W\u22121z)i are the components of vector W\u22121z. Therefore (I \u2212 (T,0)) = (I \u2212 (T,0)) 2\u2211 i=1 (W\u22121(T,0) )iwi (T,0) = 2\u2211 i=1 (1\u2212 i (T,0))(W\u22121(T,0) )iwi (T,0) moreover\u222b T 0 (T, )B( )d = \u222b T 0 (T, ) 2\u2211 j=1 (W\u22121(T, )B( )) jw j (T, )d = 2\u2211 j=1 \u222b T 0 (W\u22121(T, )B( )) j j (T, )w j (T, )d = 2\u2211 i=1 ( 2\u2211 j=1 \u222b T 0 j (T, )(W\u22121(T, )B( )) j \u00b7(W\u22121(T,0)w j (T, ))id ) wi (T,0) Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc Therefore, by (56) it follows that (W\u22121(T,0) )1= 2\u2211 j=1 1 1\u2212 1(T,0) \u222b T 0 j (T, )(W\u22121(T, )B( )) j (W \u22121(T,0)w j (T, ))1 d (58) Moreover, Equation (56) can also be written in the form ( (T,0)\u22121\u2212 I ) = \u222b T 0 ( ,0)\u22121B( )d (59) and then we obtain ( (T,0)\u22121\u2212 I ) = 2\u2211 i=1 ( i (T,0)\u22121\u22121)(W\u22121(T,0) )iwi (T,0) \u222b T 0 ( ,0)\u22121B( )d = 2\u2211 i=1 ( 2\u2211 j=1 \u222b T 0 j ( ,0) \u22121(W\u22121( ,0)B( )) j (W \u22121(T,0)w j ( ,0))i d ) wi (T,0) moreover, by (59) (W\u22121(T,0) )2= 2\u2211 j=1 2(T,0) 1\u2212 2(T,0) \u222b T 0 j ( ,0) \u22121(W\u22121( ,0)B( )) j (W \u22121(T,0)w j ( ,0))2 d (60) Note that if w1\u2208 1 and w2\u2208 2, then \u2200x \u2208 1\u222a 2 it is |(W\u22121x)1|\u2228|(W\u22121x)2| \u221a f2(1+ f2) f1(1+ f1) \u2016x\u2016 (61) In fact, first of all, we can suppose that \u2016x\u2016=1, then W and x can be written in the following way: W = \u239b \u239c\u239c\u239c\u239d 1\u221a 1+a \u2212 1\u221a 1+b \u221a a\u221a 1+a \u221a b\u221a 1+b \u239e \u239f\u239f\u239f\u23a0 , x= \u239b \u239c\u239c\u239c\u239d 1\u221a 1+c \u00b1\u221a c\u221a 1+c \u239e \u239f\u239f\u239f\u23a0 where the sign is \u2018+\u2019 if x \u2208 1 and \u2018\u2212\u2019 if x \u2208 2 and a, b, c are suitable real numbers such that f1 a,b,c f2. Then W\u22121x= \u239b \u239c\u239c\u239c\u239c\u239c\u239d \u221a 1+a\u221a 1+b \u221a b\u00b1\u221a c\u221a a+\u221a b \u221a 1+b\u221a 1+c \u00b1\u221a c\u2212\u221a a\u221a a+\u221a b \u239e \u239f\u239f\u239f\u239f\u239f\u23a0 which implies (61), since \u221a 1+a\u221a 1+c \u2228 \u221a 1+b\u221a 1+c \u221a 1+ f2\u221a 1+ f1 , |\u221ab\u00b1\u221a c|\u221a a+\u221a b \u2228 |\u00b1\u221a c\u2212\u221a a|\u221a a+\u221a b \u221a f2\u221a f1 Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc Moreover, since B(t)= ( 0 d(t) ) , W\u22121B(t)= d(t)\u221a a+\u221a b (\u221a 1+a \u221a 1+b ) therefore |(W\u22121B(t))i | \u221a 1+ f2 2 \u221a f1 \u2016B(t)\u2016, i=1,2 (62) Furthermore, \u2200z\u2208R2, ( z1 z2 ) =W (W\u22121z)= \u239b \u239c\u239c\u239c\u239d (W\u22121z)1\u221a 1+a + (W\u22121z)2\u221a 1+b (W\u22121z)1 \u221a a\u221a 1+a + (W\u22121z)2 \u221a b\u221a 1+b \u239e \u239f\u239f\u239f\u23a0 which implies that, \u2200z\u2208R2, \u2200i=1,2 |z1| 1\u221a 1+ f1 (|(W\u22121z)1|+|(W\u22121z)2|) |z2| \u221a f2\u221a 1+ f2 (|(W\u22121z)1|+|(W\u22121z)2|) (63) Therefore, from (58), (60), (61), (62) and Lemma 1, setting c= 1 2 (1+ f2)/ f1 \u221a f2/(1+ f1) |(W\u22121(T,0) )1|\u2228(W\u22121(T,0) )2| c\u2016B\u2016[0,T ] { 1 1(T,0)\u22121 [ 1(T,0) \u222b T 0 1( ,0) \u22121d + \u222b T 0 2(T, )d ] \u2228 1 1\u2212 2(T,0) [ 2(T,0) \u222b T 0 1( ,0) \u22121 d + \u222b T 0 2(T, )d ]} c\u2016B\u2016[0,T ] { 1 1(T,0)\u22121 [ 1(T,0) \u222b T 0 e\u2212 0 d + \u222b T 0 e\u2212 0(T\u2212 ) d ] \u2228 1 1\u2212 2(T,0) [ 2(T,0) \u222b T 0 e\u2212 0 d + \u222b T 0 e\u2212 0(T\u2212 ) d ]} =c\u2016B\u2016[0,T ] ( 1(T,0)+1 1(T,0)\u22121 \u2228 2(T,0)+1 1\u2212 2(T,0) ) (1\u2212e\u2212 0T) c\u2016B\u2016[0,T ] 1+e\u2212 0T 1\u2212e\u2212 0T 1\u2212e\u2212 0T 0 which implies the thesis by (63). Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc Corollary 1 Assume that A(t) and B(t) satisfy the hypotheses of Lemma 2 and that A(t) and B(t) are periodic of period T . If \u2208R2 is such that there exists 0 and T 0 for which = (T + , ) + \u222b T+ (T + , p)B(p)dp (64) where (t, ) is a solution of (46), then | i | ci 0 \u2016B\u2016[ ,T+ ], i=1,2 (65) where 0 and ci are given by (55). Proof By Floquet\u2019s theorem, since A(t) is periodic there exists a T -periodic, invertible, continuous matrix P(t) and a constant matrix E such that (t,0)= P(t)eEt . Then (t, )= P(t)eE(t\u2212 )P\u22121( ), which implies by T -periodicity of P that (T + , )= ((n+1)T + ,nT + ) \u2200n\u2208N Then (nT + , )= (nT + , (n\u22121)T + ) ((n\u22121)T + , (n\u22122)T + ) \u00b7 \u00b7 \u00b7 (T + , )= (T + , )n Therefore from (64) it follows by induction that = (nT + , ) + \u222b nT+ (nT + , p)B(p)dp \u2200n\u2208N which implies (65), by (54). 7. CONCLUSIONS We have presented a homotopy method for non-causal dynamic inversion for non-minimum phase nonlinear systems, which seems to be an alternative to those existing in the literature and has been useful to deal in a rigorous way with the exact output tracking problem for some well-known systems with two-dimensional unstable internal dynamics. REFERENCES 1. Shiriaev A, Pogromsky A, Ludvigsen H, Egeland O. On global properties of passivity-based control of an inverted pendulum. International Journal of Robust and Nonlinear Control 2000; 10(4):283\u2013300. 2. Mazenc F, Bowong S. Tracking trajectories of the cart\u2013pendulum system. Automatica 2003; 39(4):677\u2013684. 3. Sastry H, Hauser J, Meyer G. Nonlinear control design for slightly non-minimum phase systems: application to v/stol aircraft. Automatica 1992; 28:665\u2013679. 4. Martin P, Devasia S, Paden B. A different look at output tracking: control of a vtol aircraft. Automatica 1996; 32:101\u2013107. 5. Al-Hiddabi SA, McClamroch NH. Tracking and maneuver regulation control for nonlinear non-minimum phase systems: application to flight control. IEEE Transactions on Control Systems Technology 2002; 10(6):780\u2013792. Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc 6. Do KD, Jiang ZP, Pan J. On global tracking control of a vtol aircraft without velocity measurements. IEEE Transactions on Automatic Control 2003; 48(12):2212\u20132217. 7. Hauser J, Saccon A, Frezza R. Achievable motorcycle trajectories. 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Dunyak J, Junkins JL, Watson LT. Robust nonlinear least squares estimation using the Chow\u2013Yorke homotopy method. Journal of Guidance, Control, and Dynamics 1984; 7:752\u2013755. 14. Dikoussar VV. Continuation methods in boundary value problems. Computational Optimal Control. Birkhauser Verlag: 1994; 65\u201370. 15. Reif K, Weinzierl K, Unbehauen AZR. A homotopy approach for nonlinear control synthesis. IEEE Transactions on Automatic Control 1998; 43(9):1311\u20131318. 16. Consolini L, Tosques M. On the vtol exact tracking with bounded internal dynamics via a Poincare\u0301 map approach. IEEE Transactions on Automatic Control 2007; 52(9):1757\u20131762. 17. Piazzi A, Visioli A. Optimal non-causal set-point regulation of scalar systems. Automatica 2001; 37(1):121\u2013127. Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc" ] }, { "image_filename": "designv11_61_0002412_icems.2009.5382984-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002412_icems.2009.5382984-Figure4-1.png", "caption": "Fig. 4. Analysis models", "texts": [ " A three-dimensional model is used for the vibration analysis. The obtained electromagnetic force is transformed as the vibration force in the vibration analysis. Although an outer case of the stator was not included in the model of the magnetic analysis, it should be included in a model of the structural analysis, because this much effects on the structural analysis. The triangle element is used for the magnetic analysis, and the hexahedron element is used for the structural analysis. The analysis models are shown in Fig. 4. The conditions for each analysis are explained. A PWM carrier frequency has little effect on the electromagnetic vibration, because the largest resonance frequency was found around 3000Hz, the PWM carrier frequency is higher than it. Therefore the sine wave current is added to the magnetic analysis. The front cover of the motor is tightly attached to the motor mount, the surface is selected as the constraint in the structural analysis. The surface is shown in Fig. 5. The electromagnetic force distribution is shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003757_siitme.2013.6743683-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003757_siitme.2013.6743683-Figure1-1.png", "caption": "Fig. 1. The kinematic scheme of the vibrating platform with", "texts": [ " According to the researches, these low Position Resonance frequency Body part Frequency value upright Eye 20 Hz Head 18 Hz Muscle 7-15 Hz Inner organs 8 Hz Spine 8 Hz Whole body 5 Hz The fact that Table I presents the resonance frequencies for certain body parts and some WBV-devices work exactly with this frequency might look contradictory, but the frequency for the different technologies are applied for a short time. 2013 IEEE 19 th International Symposium for Design and Technology in Electronic Packaging (SIITME) 978-1-4799-1555-2/13/$31.00 \u00a92013 IEEE 242 24\u201327 Oct 2013, Gala\u0163i, Romania As in [5], the frequency of 5 Hz used in case of stochastic resonance is on the lowest level of what training experts expect to be functional. II. THE VIBRATING PLATFORM MODEL As in Fig. 1 and according to [7, 8, 9, 10], a model of a vibrating platform was developed, based on an electromagnetic actuator. The vibrating platform consists of a rigid plate 1, articulated on one side to the frame 2, an elastic element 3, an electromagnetic actuator consisting of a moving core 4 and a fixed coil 5, an electronic control system 6 and a vibration transducer 7. electro-mechanic actuator The weight G of the person standing on the plate is compensated by the opposed elastic force F2. The electromagnetic force F1 induced by the movement of the mobile core inside the fixed coil of the electromagnetic actuator produces the plate oscillations in a vertical plane and the vibrations, with a required frequency and amplitude, are transmitted to the bones and to the muscles of the patient" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002678_msf.618-619.291-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002678_msf.618-619.291-Figure3-1.png", "caption": "Figure 3. Current solutions to tensile testing issues a) notched specimen [13] and b) simulated dental prosthesis [15].", "texts": [ " Beyond optical microscopy little work has been done on characterising these changes experienced during conduction welding. Reducing the detrimental effect of these changes through careful selection of process parameters is the key to improving the mechanical properties of the weld. This however is not an easy task and is impeded by a number of issues. Measuring the tensile properties of the weld is made difficult by the strengthening effect. In order to get around this most researchers use bend tests, notched specimens or attempt application specific geometries (Figure 3) [13-15]. Bend tests however give limited elongation data and little work has gone into comparing notched and un-notched tests. These problems make it hard to infer comprehensive mechanical property data. When evaluating GTAW titanium welds the weld is orientated along the gauge length rather than transverse to it. This might be achieved for laser welding by developing smaller test geometries. The number of process variables also makes it hard to reach general conclusions about laser welding. Process parameters such as surface finish can greatly affect the weld but a rarely specified in research making it hard to reproduce results even for application specific testing" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002930_icelmach.2008.4799943-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002930_icelmach.2008.4799943-Figure6-1.png", "caption": "Fig. 6. Field distribution inside PM-V machine under load", "texts": [], "surrounding_texts": [ "Referring to the figure 5 the variation of the effective air-gap length for the salient pole machines with non-uniform sinusoidal air-gap length along the circumferential direction can be described using the following function,\ng(O) =_--,,\"',..---__ 1 ___ _\na l + La2,J . cos(j. 2(;8 - 0) + IPJ) (5)\nj=1\nwhere the minimum air-gap length is (al + a 2,)-1 and the maximum is (al -a2,Jrl \u2022 With j are denoted the air-gap\npermeance harmonics, and ;8 is the angular position at the\nstator surface.\nUsing the correct expression for the effective air-gap length, the correct expressions for the self-, mutual- and phase inductance and also for the total current flux-linkage are derived in following.\nIII. I SELF AND MUTUAL INDUCTANCES\nFor the determination of self-inductance it is necessary to compute the flux linking a winding due to its own current. Let us consider the flux-linkage of a single tum of a stator winding which spans Ws and which is located at an angle ;8' In this case the flux can be determined by performing a surface integral over the open surface of the single tum, In order to obtain the flux-linkage of the entire winding, the flux linked by each tum must be summed, Therefore we have (/ represents the active length of the machine, the superscript (ole) denotes the parameters of the electric machine taking into account the air gap permeance harmonics),\n~\nII'\" =N f2 B('\" O),r,/-d'\" =J:..W,r,/.i . Tw s 'l'S' 'l's s u ~ It\n2\nSolving the above integral equation leads to\n(7)\nwhere,\n(8)\n(9)\nHaving the expression for the current flux-linkage, the self-inductance of the winding-u is obtained by dividing equ, (7) by iu ' Thus,\nL \" 4t'p N2 / L'\" ;,2{ It L'\" t;k,J ( , .\"., )} =- \u00b7r\u00b7\u00b7 a +_. -\u00b7a \u00b7cos j.=-m uu 2' k 1 2 1< 2.J Y\"J\nIt k=1 t'p J=I ':>k\n(10)\nIn analogy, the expression for the mutual inductance taking into account the rotor permeance effect is:\n(11)\nThe remammg self- and mutual inductances may be calculated using the same procedure as above. We can express these inductances compactly by defining,\n(12)\n2 '\" [} =-N2 \u00b7r\u00b7/\u00b7a \u00b7L~\u00b7'\u00b7 Bits 2,J k=1 k k,}\n(13)\nThus for the phase-u we have,\nL:\" =LA + i:L{. .cos(j\u00b720-IPJ) (14) J=I\n(15)\nThe above analysis shows that due to the air-gap permeance harmonics, high harmonic components appear in the self- and mutual inductances.\nIn the following analysis the inductances of a PM machine with V-magnets in the rotor (PM-V machine) are investigated using FEM. The following figures 6 and 7 shows flux density", "distribution and also the self- and mutual inductances of the PM-V machine under high load operation condition, / = 500A , 8=00.\nThe FE simulations show that due to the saturation effect and also due to the effect of the magnet pole shape (rotor permeance effect), high harmonic components appear in the inductance curves of the PM-V machine. Using FFT analysis the harmonic components of the self-inductance are obtained. The following figure 8 shows harmonic components of the LB\nparameter.\nFurther, the expressions for the self- and mutual inductances can be written as\nL*\",. =LA +L~ .cos(20)\n+ tLi. . cosU \u00b720 - tpj) (16)\n}=2\n\u2022 LA I (Ll 21f) L =--+L \u00b7cos 2u-uv 2 B 3\n\"\". ( 21f) + ~)Is'cos j\u00b720-tp}-}=2 3\n(17)\nSubstituting equs. (1) and (2) into (16) and (17), we have:\nL:u =Luu + tLi. .cosU\u00b720-tpj) (18) j=2\nL* =L + tv 'COS(j.20-tp _ 21f) uv uv }=2 B } 3 (19)\nThe above expressions show that the self- and mutual inductances of salient pole synchronous machines are influenced from the rotor permeance effect.\nIII.2 TOTAL PHASE INDUCTANCE\nThe phase inductance (synchronous inductance) is the effective inductance seen by one phase under the balanced 3- phase conditions of normal machine operation. Under balanced three phase condition the expression for the current flux linkage in the phase-u taking into account the air-gap rotor permeance harmonics is:\n,,< = L(8)./ . cos(wt + 8 -r)\n+ i\u00b7 tv. cos[(2j -1). wt - tp} -8] (20)\nj=2\nThe above equ. (20) shows that for the salient pole synchronous machines with non-uniform air-gap length the current flux-linkage contains some time-harmonics even if the electric machine is supplied with purely sinusoidal currents. However, the time-harmonics on the current flux-linkage appear as a result of the air-gap permeance harmonics. The following figures 9 and 10 show the current flux-linkage of the the PM-V machine obtained with FEM and corresponding harmonic components. Like in the previous section the FE simulations are performed under high load condition, i = 500A , 8=00.", "Comparing self-inductance harmonics presented in the figure 8 and the harmonic components of the total current flux linkage given in the above figure 10 it is shown that in the current flux-linkage appear high harmonic components of the order \"2j - I\"; the obtained results validate the expression for the current flux-linkage given in the equ. (20).\nThe above equ. (20) shows that for I! = 0 (electric machines\nwith sinusoidal air-gap length) the total current flux-linkage is purely sinusoidal. Thus,\nfor: I! =0 ~ V< =L(0).1 ,cos(IVt+o-y) (21)\nReferring to the equ. (20) and (21) it is shown that the time-harmonics appear in the current flux-linkage only as a result ofthe air-gap permeance harmonics;\n\u2022 winding harmonics don't contribute on the time-harmonics\n\u2022 time-harmonics of voltage supply aren't considered here\nFurther, the equ. (20) can be written as:\n(22)\nwhere, 'l/u represents the fundamental component of the\ncurrent flux-linkage,\n'l/u =L(o).i,cos(IVt+o-y) (23)\nand, i'l/u represents the harmonic components on the current\nflux-linkage curve,\ni'l/u = i\u00b7 I.Il . cos[(2i -1). IVt -fIJi - 0] (24) j~2\nAccording to the equs. (22) to (24) it is shown that by the salient pole machines with non-uniform air-gap length, the expression for the current flux-linkage '1/; consists of total\nphase inductance L( 0) , and harmonic inductances\nI! =(3/2).L1t.\nIV. CONCLUSION\nIt is well known that air-gap length has a remarkable effect on the performances of the electric machines. Since until now the theory of the salient pole synchronous machine is based only on the idealized machines (with sinusoidally variation of the air-gap length), the main motivation of this work was to investigate the influence of the air-gap length on the performances of this type of machines. Using the correct expression for the effective air-gap length the effect of the air gap rotor permeance harmonics on the self-, mutual and phase inductance and also on the current flux-linkage, is investigated and analysed. Based on the obtained results the following conclusions are made:\n\u2022 Air-gap rotor permeance harmonics lead to high harmonic components on the self-, mutual-, and phase inductances of the salient pole synchronous machines,\n\u2022 Time-harmonics appear in the current flux-linkage only as a result of the air-gap permeance harmonics;\no winding harmonics don't contribute on the time-harmonics,\no time-harmonics of voltage supply aren't considered here.\n\u2022 Since the phase voltage and electromagnetic torque depends on the flux-linkage of the machine, the rotor permeance harmonics cause high harmonics in the phase voltage and electromagnetic torque (torque ripple) of this type of machines. An accurate analysis of the air-gap permeance effect on the phase voltage and electromagnetic torque should be performed in a future work." ] }, { "image_filename": "designv11_61_0000010_s0167-8922(03)80068-8-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000010_s0167-8922(03)80068-8-Figure1-1.png", "caption": "Figure 1. The dynamic test rig: (1) vibration shaker, (2) force transducer, (3) moving cylinder, (4) accelerometer, (5) ball, (6) glass disc, (7) tri-axis force transducer, (8) rigid frame.", "texts": [ " Capabilities of the machine are given and frequency properties are reported. *Corresponding author. E-mail address: joel.perret-liaudet@ec-lyon.fr In a third part, we present the main characteristics of the EHL machine. The optical method uses interferences created in the EHL contact from a normal incident white light permits to measure the oil thickness and also to visualise the lubricated contact. Some preliminary results with or without rolling speed will be given to demonstrate the efficiency of the apparatus The experimental setup is shown in Fig. 1. It consists on a double sphere-plane Hertzian contact. The SAE 52100 steel ball is preloaded between two horizontal plane surfaces, which one is a glass disc fixed to a heavy rigid flame and the other is a SAE 52100 steel disc rigidly fixed to a cylinder moving like a rigid body. Planes were grotmd to obtain low roughness (Ra < 0.4 #m). Ball roughness is very small (Ra < 0.02 #m). With these conditions, the Hertzian theory can be applied. The contact is preloaded by the weight of the moving cylinder which is held by six thin stems in order to prevent lateral displacements and rotations" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000117_135065003322620246-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000117_135065003322620246-Figure4-1.png", "caption": "Fig. 4 Details of mounting arrangements for sensors used to extract signals from reciprocating machinery: (a) an inductive transducer mounted in the connecting rod bearing to measure oil \u00aelm thickness (and also incorporating a dummy transducer to offset transducer sensitivity to temperature) as used by Goodwin and Holmes [11]; (b) an inductive transducer mounted in the crank journal to measure oil \u00aelm thickness and extracting the signals via slip rings, as used by Goodwin et al. [53]", "texts": [ " The authors concluded that the discrepancies between theory and experiment were due to non-circularity of the bearing and to bearing elastic deformation. The following year, Bates and Evans [4] published details of a measurement procedure for recording bigend bearing oil \u00aelm thickness in a working engine. The method used a mechanical linkage similar to that described by Westbrook and Munro [40], but it had a lighter construction and the arrangement incorporated capacitance transducers within the bearing shell (see Fig. 4a) to record journal position, and thermistors for temperature measurement also in the bearing shell. The results obtained from the experimental work showed excellent repeatability from one engine run to the next, and showed oil \u00aelm thickness variation with crank angle to agree broadly with theoretical predictions obtained using the mobility method based on a short bearing model. A possible criticism of the technique used by the authors is the possibility of oil \u00aelm cavitation affecting the bearing clearance dielectric constant, and hence upsetting the oil \u00aelm thickness measurements", " His results also showed that the temperature range around the bearing circumference can be up to 8 K, and the mean bearing temperature can be over 30 K higher than the oil feed temperature at high engine speeds. Experimental measurements of big-end bearing journal orbits were published by Goodwin et al. [53] on the basis of measurements made on a single cylinder fourstroke diesel engine. Measurements were made using inductive transducers located in the crankshaft journal, and signals were extracted using slip rings (see Fig. 4b). The authors concluded that the procedure was a feasible method for obtaining experimental data but was dependent on accurate knowledge of the transducer temperature. The orbits published suggested considerable elastic deformation of the bearing during operation. A summary of the experimental work discussed above is given in Table 2. Consideration of the theoretical modelling work discussed above reveals that the complexity of modelling procedures has increased considerably. Whereas early work was based on rigid bearing assumptions and in some cases even on hand-worked solutions, modern computing power now enables sophisticated modelling practices to be used" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003233_pime_conf_1965_180_321_02-Figure24.9-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003233_pime_conf_1965_180_321_02-Figure24.9-1.png", "caption": "Fig. 24.9. Lubricant flow diagrams", "texts": [ " Cole and Hughes (12) Proc Iristii Mech Engrs 1965-66 found that when they had separated the supply pressure component from the total flow the remaining hydrodynamic side leakage flow (or zero supply pressure flow) was less than the theoretical side leakage flow. The discrepancy was attributed to reduction in film extent and circulation of oil to and from menisci at the sides of the bearing. It is now suggested that the concept of adding the flow Q\u201d, due to hydrodynamic pressures generated in the loadsupporting wedge, is an erroneous one. Fig. 24.9 illustrates diagrammatically the proposed models of lubricant flow in journal bearings with a single and a double inlet position. From continuity of flow the inlet flow Q must be equal to the sum of the outlet flows: QL due to side leakage flow from the loaded region and Qs a side leakage flow before the commencement of the hydrodynamic wedge at E. The flow rate Qs will depend on the inlet pressure p,, viscosity ?I, surface speed U, mean film thickness h, and the inlet groove to bearing proportions" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002250_978-3-540-92841-6_559-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002250_978-3-540-92841-6_559-Figure1-1.png", "caption": "Fig. 1. Passive Intelligent Walker", "texts": [ " Many elderly people would have some handicaps for locomotion, because of the decline of the physical strength, muscular strength, etc. In the daily life, the ability of walking is one of the most fundamental functions for humans to maintain the high quality of life. For assisting the walking of the human, several intelligent systems using robot technologies have been proposed [1]- [3]. We have also introduced two kinds of walking assist system. First one is the walker-type walking assist system called RT Walker as shown in Fig.1. Walker-type waling assist system could support the weight of the user and keep his/her balance. In addition, RT Walker achieves several functions such as collision avoidance and path following by controlling the brake torques of the rear wheels [4]. Second one is the wearable walking assist system called Wearable Walking Helper as shown in Fig.5, which consists of knee orthosis, prismatic actuator and sensors [5]. The knee joint of the orthosis has one degree of freedom rotating around the center of the knee joint of the user on sagittal plane" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003942_s13369-012-0287-1-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003942_s13369-012-0287-1-Figure2-1.png", "caption": "Fig. 2 Actual rotor and its solid model", "texts": [ " In general, the VPF vibration is regarded as one of the flow-induced vibration mechanisms in rotating machinery, which is primarily due to the vane-passing pressure pulsation. Normally, vibrations at VPF are kept at acceptable levels in the prototypal design stage of the pump, as a design acceptance criterion. Nevertheless, the diagnostic strategy of operating machines calls for investigating other possible excitation mechanisms before venturing into any procedure of design modifications. Accordingly, one of the preliminary steps of diagnostics is to examine the possibility of resonance at the recorded VPF. The BFP rotor-bearing system, shown in Fig. 2, is a stepped steel rotor with 31 stations, four impellers, two journal bearings and one tilting-pad thrust bearing. Rotor overall length is 2.38 m and total weight of 357 kg. The journal bearing is a two-lobe sleeve made of Carbon steel JIS S25CN and WJ2. The bearing lubricant is Mobil (Turbine Oil T32), with oil viscosity 32 cSt at 40 \u25e6C and 5.4 cSt at 100 \u25e6C. The bearing dimensions are Table 1 Journal bearing characteristics Speed (rpm) K b xx (N/m) K b xy (N/m) K b yx (N/m) K b yy (N/m) Cb xx (N s/m) Cb xy (N s/m) Cb yx (N s/m) Cb yy (N s/m) a" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003047_6.2009-6230-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003047_6.2009-6230-Figure5-1.png", "caption": "Figure 5. Schematic of a force-deflection curve where the hysteresis is due to damping.", "texts": [ " (9) sets a limit on the slip angle \u03b1 beyond which MK\u03b1 (p) is taken to be zero. I.A.5. Torsional and lateral damping Damping in general is a difficult quantity to measure in any given system. Especially, in the case of an aircraft tyre, whose natural frequency of oscillation changes with the inflation pressure, determining the damping coefficient is quite challenging. Typically, damping characteristics are measured using a force-deflection curve that forms a hysteresis loop due to damping, as shown in Fig. 5. The graph shows a dashed curve that represents a force-deflection curve in the case of a viscous damper. For an ideally elliptical force-deflection curve, the damping coefficient is given by cc = b bmax . (10) However, in the case of aircraft tyres, due to the complex nature of the inflated tyres the force-deflection curves deviate from the elliptical nature of the dashed curve to take the form of the solid curve in Fig. 5. In 7 of 12 American Institute of Aeronautics and Astronautics such cases using Eq. (10) may yield a reasonable estimate of the damping coefficient but is not necessarily very accurate. Despite its accuracy limitations Eq. (10) is used to compute damping coefficients in the case of aircraft tyres. Following is the experimental data representing torsional and lateral damping coefficients obtained using Eq. (10) for two different tyre pressures and several different vertical loads on the wheel. Figure 6(a) shows the lateral damping coefficient cl as a function of the normalized pressure p\u0303 for different vertical loads" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000052_cdc.1999.832740-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000052_cdc.1999.832740-Figure1-1.png", "caption": "Figure 1: The considered model of a car.", "texts": [ " In a previous work [21] we solved the problem of computing the distance to polygonal obstacles for a punctual\u2019 car-like robot. This paper presents a solution when both the robot and the obstacles have polygonal shape. The solution is based on the application of the Pontryagin Maximum Principle together with transversality conditions allowing to single out necessary conditions to be satisfied by feasible shortest paths to obstacles. Configuration space obstacles Consider the carlike robot sketched in Fig. 1, the coordinates x and y determine the position in the plane of the rear wheel We did not take into account the physical space occupied by the robot according to its shape. 0-7803-5250-5/99/$10.00 0 1999 IEEE 17 axle midpoint (reference point) of the car, while 8 is the angle that the unit direction vector v forms with the positive x axis. A configurntion of the car is given by a triple (x, y, 8) E C R2 x S\u2019, where S\u2019 is the unit circle in the plane. The point P(x, y) E R2 will be referred to as the position of the car in the plane" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003172_kikaic.74.1825-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003172_kikaic.74.1825-Figure2-1.png", "caption": "Fig. 2 A conceptual image", "texts": [], "surrounding_texts": [ "1825\n\u65e5\u672c\u6a5f\u68b0\u5b66\u4f1a\u8ad6 \u6587\u96c6(C\u7de8)\n74\u5dfb743\u53f7(2008-7)\n\u8ad6 \u6587No.08-0021\n\u30d4\u30b6\u8077 \u4eba \u306e\u30cf \u30f3 \u30c9\u30ea\u30f3\u30b0 \u30e1 \u30ab\u30cb\u30ba \u30e0\u306b\u7740 \u76ee \u3057\u305f\u52d5 \u7684\u64cd \u308a*\n\u6771 \u68ee \u5145*1, \u5185 \u6d77 \u572d \u7950*2 \u5927 \u672c \u5eb7 \u9686*3, \u91d1 \u5b50 \u771f*4\nDynamic Manipulation Inspired by Handling Mechanism of Pizza Master\nMitsuru HIGASHIMORI*5, Keisuke UTSUMI,\nYasutaka OMOTO and Makoto KANEKO\n*5 Graduate School of Engineering, Osaka University,\n2-1 Yamadaoka, Suita-shi, Osaka, 565-0871 Japan\nThis paper discusses a dynamic manipulation of an object on a plate attached at the tip of manipulator, where the basic concept is inspired by a handling mechanism of pizza master. Both the position and the orientation of object are dynamically controlled by the plate. The main driving force for manipulating the object is produced by the combination of inertial and frictional forces generated by one rotational and one translational motion of the plate. We first discuss how to make motions of object for three degrees of freedom on the plate. We then show that it is guaranteed to achieve an arbitrary set of desired position and orientation of object by the proposed manipulation scheme. Finally, we show a couple of experiments for confirming the basic idea by using a high-speed manipulator with an assistance of vision. Key Words: Robot, Skill, Manipulator, Dynamic Manipulation, High-Speed Robotics\n1. \u306f \u3058 \u3081 \u306b\n\u30ed\u30dc \u30c3 \u30c8\u30cf \u30f3 \u30c9\u306b \u3088\u308b\u5bfe\u8c61\u7269 \u306e\u628a\u63e1\u3084\u64cd \u308a\u306b\u95a2\u3059 \u308b\n\u69d8 \u3005\u306a\u7814 \u7a76\u304c\u884c\u308f\u308c \u3066\u3044\u308b\u304c,\u5f93 \u6765,\u5bfe \u8c61\u7269 \u304c\u9759\u6b62 \u3057 \u3066\u3044\u305f \u308a,\u3086 \u3063 \u304f\u308a\u3068\u904b\u52d5 \u3057\u3066\u3044\u305f \u308a\u3059 \u308b\u6e96\u9759 \u7684\u6761\u4ef6 \u3092\u4eee\u5b9a \u3057\u305f\u7814\u7a76 \u304c\u4e2d\u5fc3 \u3068\u306a\u3063\u3066\u3044\u305f(1)\uff5e(5).\u8fd1\u5e74\u3067\u306f, \u30bb \u30f3\u30b7 \u30f3\u30b0\u901f\u5ea6\u3084\u30a2 \u30af\u30c1\u30e5\u30a8\u30fc \u30b7 \u30e7\u30f3\u6280\u8853 \u306e\u5411\u4e0a\u306b \u3088\n\u308a,\u904b \u52d5\u7269\u4f53 \u306e\u30c0\u30a4\u30ca \u30df\u30c3\u30af\u306a\u6355\u7372\u3084\u64cd \u308a\u306b\u95a2\u3059 \u308b\u7814 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E-mail : higashi@mech.eng.osaka-u.ac.jp", "1826 \u30d4\u30b6\u8077 \u4eba\u306e\u30cf \u30f3 \u30c9\u30ea\u30f3\u30b0\u30e1\u30ab\u30cb\u30ba\u30e0\u306b\u7740 \u76ee\u3057\u305f\u52d5\u7684\u64cd \u308a\n\u3057,n,{IX,IY,IZ}\u306f,\u305d \u308c\u305e\u308c \u30d7 \u30ec\u30fc \u30c8\u306e\u8cea\u91cf\u304a \u3088\u3073\n\u91cd\u5fc3\u306b\u304a \u3051\u308b{X1,Y1,Z1}\u8ef8 \u56de \u308a\u306e\u6163 \u6027\u30e2\u30fc\u30e1\u30f3 \u30c8\u3067 \u3042 \u308a,\u7c21 \u5358 \u306e\u305f\u3081,\u91cd \u529b\u306e\u5f71\u97ff \u306f\u8003 \u3048\u306a\u3044,\u6b21 \u306b,\u56f32(b) \u306e \u3088 \u3046\u306b,\u30d7 \u30ec\u30fc \u30c8\u4e2d\u5fc3\u304b \u3089\u8ddd\u96e2l\u306e \u68d2\u4e0a\u306e\u70b9 \u3092\u64cd\u4f5c\n\u70b9 \u3068 \u3057\u3066\u8003 \u3048\u3066\u307f \u308b.\u68d2 \u306e\u8cea\u91cf \u3092\u7121\u8996\u3059 \u308b \u3068,\u5909 \u4f4d\u30d9 \u30af \u30c8\u30ebx2\u3013[X2,Y2,Z2,\u0398X2,\u0398Y2,\u0398Z2]T\u306e \u904b\u52d5 \u306b\u5bfe\u5fdc\u3059 \u308b\u6163 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\u3042\u308b.\u6700 \u5f8c\u306b,\u9ad8 \u901f\u30de\u30cb \u30d4\u30e5 \u30ec\u30fc \u30bf \u3068\u30d3\u30b8 \u30e7 \u30f3\u304b \u3089\u69cb\u6210 \u3055\u308c \u308b\u5b9f\u9a13\u30b7\u30b9\u30c6\u30e0 \u3092\u69cb\u7bc9 \u3057,\u63d0 \u6848\u3059 \u308b\u64cd\n\u308a\u624b\u6cd5 \u306e\u6709\u52b9\u6027 \u3092\u5b9f\u9a13 \u7684 \u306b\u78ba\u8a8d\u3059 \u308b.\n\u672c\u8ad6\u6587 \u306e\u69cb\u6210 \u306f\u4ee5 \u4e0b\u306e\u901a \u308a\u3067 \u3042\u308b.2\u7ae0 \u3067\u5f93\u6765\u7814\u7a76 \u306b\u3064 \u3044\u3066\u6574 \u7406 \u3057\u305f\u5f8c,3\u7ae0 \u3067\u306f,\u89e3 \u6790\u30e2 \u30c7\u30eb \u3092\u793a \u3057\u305f \u4e0a \u3067,\u672c \u7814 \u7a76\u3067\u53d6 \u308a\u6271 \u3046\u64cd \u308a\u554f\u984c \u306b\u3064\u3044\u3066 \u5b9a\u7fa9\u3059 \u308b. 4\u7ae0 \u3067\u306f,\u64cd \u308a\u6226\u7565 \u306e\u67a0\u7d44 \u307f\u306b\u3064\u3044\u3066\u793a\u3059.5\u7ae0 \u3067\u306f, \u5b9f\u6a5f \u5b9f\u9a13\u306e\u7d50\u679c \u3092\u793a\u3059.\n2. \u5f93 \u6765 \u7814 \u7a76\n\u30d7 \u30ec\u30fc \u30c8\u306b\u3088\u308b\u5bfe\u8c61\u7269 \u306e\u64cd \u308a\u554f\u984c \u3092\u53d6 \u308a\u6271\u3063\u305f\u7814\u7a76 \u3068\u3057\u3066,\u8352 \u4e95 \u3089(10)\u306f,\u30de \u30cb \u30d4\u30e5 \u30ec\u30fc \u30bf\u5148\u7aef \u306b\u53d6 \u308a\u4ed8 \u3051 \u3089\u308c\u305f\u30d7 \u30ec\u30fc \u30c8\u4e0a \u306b\u7f6e \u304b\u308c \u305f\u7acb\u65b9\u4f53\u5bfe\u8c61 \u7269\u306e\u64cd \u308a\u554f \u984c\u306b\u3064\u3044\u3066\u8b70\u8ad6 \u3057,\u5bfe \u8c61\u7269 \u3092 \u81ea\u8eab\u306e\u30a8 \u30c3\u30b8\u5468 \u308a\u306b\u56de\u8ee2 \u3055\u305b \u308b\u305f\u3081\u306e\u6226\u7565\u306b\u3064\u3044\u3066\u8003\u5bdf \u3057\u3066\u3044\u308b.Lynch\u3089(11) \u306f,\u624b \u306e\u5e73\u306e \u5185\u9762\u304a\u3088\u3073\u80cc \u9762\u306b\u6cbf\u3063\u3066\u30dc\u30fc\u30eb \u3092\u79fb\u52d5 \u3055 \u305b \u308b\u30b8\u30e3\u30b0\u30ea\u30f3\u30b0\u30b9\u30ad\u30eb\"Butterny\"\u3092 \u53d6 \u308a\u4e0a\u3052,\u30ed \u30dc \u30c3\n\u30c8\u30d1\u30fc \u30e0\u306e\u5f62 \u72b6 \u3068\u904b\u52d5\u8a08\u753b \u306b\u3064\u3044 \u3066\u8b70\u8ad6 \u3057\u3066\u3044 \u308b.\u96e8\n\u6d77 \u3089(12)\u306f,\u30d3 \u30b8 \u30e7\u30f3\u60c5\u5831 \u306b\u57fa\u3065 \u304d,6\u81ea \u7531\u5ea6 \u30de\u30cb \u30d4\u30e5\n\u30ec\u30fc \u30bf\u5148\u7aef \u306b\u53d6 \u308a\u4ed8 \u3051\u3089\u308c\u305f\u30d7 \u30ec\u30fc \u30c8\u4e0a \u3067\u5bfe\u8c61\u7269 \u3092\u64cd\n\u4f5c\u3059 \u308b\u5b9f\u9a13 \u3092\u884c\u3063\u3066\u3044 \u308b.Reznik\u3089(14)\u306f,3\u81ea \u7531\u5ea6\u306e \u6c34 \u5e73\u632f\u52d5 \u30d7 \u30ec\u30fc \u30c8\u306b\u3088\u308bUniversal Planar Manipulator (UPM)\u3092 \u958b\u767a \u3057,\u30d3 \u30b8 \u30e7\u30f3\u30b7\u30b9\u30c6 \u30e0\u3068\u7d44\u307f\u5408 \u308f\u305b,\u8907 \u6570 \u5bfe\u8c61\u7269\u3092\u540c\u6642 \u306b\u7570 \u306a\u308b \u76ee\u6a19\u65b9 \u5411\u3078 \u3068\u79fb\u52d5 \u3055\u305b \u308b\u5b9f\u9a13 \u3092\u884c\u3063\u3066\u3044 \u308b.Bohringer\u3089(15)\u304a \u3088\u3073Vose\u3089(16)\u306f, \u30d7 \u30ec\u30fc \u30c8\u306b\u4efb\u610f\u8ef8\u5468 \u308a\u306e\u632f \u52d5\u3092\u4e0e \u3048,\u30d7 \u30ec\u30fc \u30c8\u4e0a\u306e\u5bfe\n\u8c61\u7269\u306e\u6319\u52d5 \u3092\u30bb \u30f3\u30b5 \u30ec\u30b9\u3067\u7ba1\u7406\u3059 \u308b\u65b9\u6cd5\u306b\u3064\u3044\u3066\u8b70\u8ad6 \u3057,\u57fa \u672c \u539f\u7406 \u30921\u8ef8 \u632f\u52d5\u767a\u751f\u6a5f\u69cb \u3092\u7528\u3044\u3066\u5b9f\u9a13\u7684 \u306b\u78ba \u8a8d \u3057\u3066\u3044 \u308b.\u4ee5 \u4e0a \u306e\u3088 \u3046\u306b\u30d7 \u30ec\u30fc \u30c8\u306b \u3088\u308b\u5bfe\u8c61\u7269 \u306e\u64cd\n\u308a\u306b\u95a2\u3059 \u308b\u7814 \u7a76\u304c\u3044 \u304f\u3064\u304b\u884c\u308f\u308c\u3066\u3044 \u308b\u304c,\u7b46 \u8005 \u3089\u306e\n\u77e5 \u308b\u9650 \u308a,\u672c \u7814\u7a76 \u306e\u3088 \u3046\u306b\u30d7 \u30ec\u30fc \u30c8\u306e\u80fd\u52d5\u7684\u306a\u904b\u52d5 \u65b9 \u5411 \u30922\u81ea \u7531\u5ea6 \u306b\u9650\u5b9a \u3057\u305f\u4e0a\u3067,\u30c0 \u30a4\u30ca \u30df\u30c3\u30af\u306a\u52b9\u679c \u3092 \u7a4d\u6975\u7684 \u306b\u7528 \u3044\u3066\u5bfe\u8c61\u7269\u306e3\u81ea \u7531\u5ea6 \u3092\u81ea\u5728 \u306b\u64cd\u4f5c\u3059 \u308b \u3053\n\u3068\u3092\u8a66 \u307f\u305f\u7814\u7a76 \u306f\u898b \u5f53\u305f \u3089\u306a\u3044.", "\u30d4\u30b6\u8077\u4eba \u306e\u30cf \u30f3 \u30c9\u30ea\u30f3\u30b0\u30e1\u30ab\u30cb\u30ba\u30e0 \u306b\u7740 \u76ee \u3057\u305f\u52d5\u7684\u64cd \u308a 1827\n3. \u554f \u984c \u8a2d \u5b9a\n3\u30fb1 \u30e2\u30c7\u30eb\u5316 \u56f33\u306b \u793a\u3059 \u3088 \u3046\u306a \u30d7 \u30ec\u30fc \u30c8\u3068\u5bfe \u8c61 \u7269 \u3092\u8003 \u3048\u308b.\u7c21 \u5358\u5316 \u306e\u305f\u3081,\u4ee5 \u4e0b \u306e\u4eee \u5b9a\u3092\u8a2d \u3051\u308b.\n\u4eee \u5b9a1: \u5bfe\u8c61\u7269 \u3068\u30d7 \u30ec\u30fc \u30c8\u306f\u525b\u4f53 \u3068\u3059 \u308b. \u4eee \u5b9a2: \u5bfe\u8c61\u7269 \u306e\u8cea \u91cf\u5206\u5e03 \u306f\u4e00\u69d8 \u3068\u3057,\u539a \u307f\u306f\u7121\u8996 \u3067\n\u304d\u308b\u307b \u3069\u5c0f \u3055\u3044.\n\u4eee\u5b9a3: \u30d7 \u30ec\u30fc \u30c8\u306f,\u5bfe \u8c61\u7269 \u304c\u843d\u4e0b \u3057\u306a\u3044\u7a0b\u5ea6 \u306b\u5341\u5206\n\u5927 \u304d\u3044.\n\u4eee\u5b9a4: \u5bfe\u8c61\u7269 \u3068\u30d7 \u30ec\u30fc \u30c8\u3068\u306e\u63a5\u89e6\u306f,\u5e38 \u306b\u5747\u4e00 \u306a\u9762\n\u63a5\u89e6 \u304c\u7dad \u6301 \u3055\u308c \u308b.\n\u4eee\u5b9a5: \u5bfe\u8c61 \u7269 \u3068\u30d7 \u30ec\u30fc \u30c8\u9593\u306e\u6469\u64e6\u4fc2 \u6570 \u306f\u4e00\u69d8 \u306b \u03bc\n\u3068\u3057,\u9759 \u6469\u64e6 \u3068\u52d5\u6469\u64e6 \u306f\u533a\u5225 \u3057\u306a\u3044.\n\u4eee\u5b9a6: \u5bfe\u8c61\u7269\u304a \u3088\u3073\u30d7 \u30ec\u30fc \u30c8\u306e\u4f4d\u7f6e \u30fb\u59ff\u52e2 \u306f\u30bb \u30f3\u30b5\n\u7cfb\u304b \u3089\u5f97 \u3089\u308c \u308b.\n\u56f33\u5185 \u306b\u793a\u3059\u8a18 \u53f7\u306e\u610f\u5473\u306f,\u4ee5 \u4e0b\u306e\u901a \u308a\u3067 \u3042\u308b.\n\u03a3R: \u57fa \u6e96\u5ea7\u6a19 \u7cfb.\u305f \u3060 \u3057,xR-yR\u5e73 \u9762 \u306f\u6c34\u5e73 \u3068\n\u3059 \u308b.\n\u03a3m: \u30d7 \u30ec\u30fc \u30c8\u306b\u56fa\u5b9a \u3055\u308c \u305f\u5ea7\u6a19 \u7cfb. \u03a3B: \u5bfe\u8c61\u7269 \u306e\u91cd\u5fc3\u4f4d\u7f6e \u306b\u56fa\u5b9a \u3055\u308c\u305f\u5ea7\u6a19\u7cfb.\u305f\n\u3060 \u3057,ZB\u8ef8 \u306f\u30d7 \u30ec\u30fc \u30c8\u3068\u306e\u63a5\u89e6\u9762 \u306b\u76f4\u4ea4\u3059 \u308b.\nmxB ,myB: \u03a3m\u304b \u3089\u898b \u305f \u03a3B\u306e \u4f4d\u7f6e m\u03b8B: \u03a3 m\u304b \u3089\u898b \u305f \u03a3B\u306eZm\u8ef8 \u56de \u308a\u306e\u56de\u8ee2 \u89d2\u5ea6. Rx\nB,RyB: \u03a3R\u304b \u3089\u898b\u305f \u03a3B\u306e\u4f4d \u7f6e. R\u03b8B: \u03a3R\u304b \u3089\u898b\u305f \u03a3B\u306eZR\u8ef8 \u56de \u308a\u306e\u56de\u8ee2\u89d2\u5ea6 .\nmB: \u5bfe\u8c61\u7269 \u306e\u8cea \u91cf, AB: \u5bfe\u8c61\u7269\u306e\u63a5\u89e6 \u9762\u7a4d.\ng: \u91cd\u529b\u52a0\u901f\u5ea6.\n\u672c\u7814\u7a76 \u3067\u306f,\u30d4 \u30b6\u8077\u4eba \u306e\u64cd \u4f5c\u30a2\u30ca \u30ed\u30b8\u30fc \u3092\u8e0f\u307e \u3048\u3066, \u56f33\u306b \u793a\u3059 \u3088 \u3046\u306b,\u30d0 \u30fc\u5148\u7aef \u306b\u30d7 \u30ec\u30fc \u30c8\u304c\u88c5\u7740 \u3055\u308c\u3066 \u3044 \u308b\u3082\u306e \u3068\u3059 \u308b.\u4e21 \u8005 \u306e\u63a5\u7d9a\u90e8 \u306b \u03a3m\u304c \u56fa\u5b9a \u3055\u308c,xm \u8ef8\u304a \u3088\u3073Zm\u8ef8 \u306f,\u305d \u308c\u305e\u308c\u30d0 \u30fc\u306e\u9577\u624b \u65b9\u5411,\u30d7 \u30ec\u30fc\n\u30c8\u9762\u306e\u6cd5\u7dda\u65b9\u5411 \u306b\u4e00\u81f4 \u3057\u3066\u3044 \u308b\u3082\u306e \u3068\u3059 \u308b.X,\u0398 \u306f,\n\u305d\u308c \u305e\u308c \u30d7 \u30ec\u30fc \u30c8\u306e \u4e26\u9032\u904b \u52d5\u306e\u5909 \u4f4d,\u30d7 \u30ec\u30fc \u30c8\u306e\u50be \u304d\u904b\u52d5 \u306e\u5909\u4f4d \u306b\u5bfe\u5fdc\u3059 \u308b.\u305f \u3060 \u3057,\u30d7 \u30ec\u30fc \u30c8\u304cxR-yR \u5e73\u9762 \u306b\u5bfe \u3057\u3066\u5e73\u884c(\u6c34 \u5e73)\u306a \u3068\u304d\u3092\u0398=0\u3068 \u3059 \u308b.\n(a) The xm-direction\n(b) The y7 direction\n3\u30fb2 \u554f \u984c \u8a2d \u5b9a \u30d7 \u30ec\u30fc \u30c8\u306e\u904b \u52d5 \u306b \u3088\u3063\u3066,\u30d7\n\u30ec\u30fc \u30c8\u304b \u3089\u898b \u305f\u5bfe \u8c61 \u7269 \u306e \u76f8 \u5bfe\u4f4d \u7f6e \u30fb\u59ff \u52e2 \u306b\u95a2 \u3059 \u308b\n3\u6210 \u5206(mxB,nyB,n\u03b8B)\u3092 \u64cd \u4f5c \u3057,\u4efb \u610f \u306e \u521d \u671f \u4f4d \u7f6e \u30fb\u59ff\u52e2(nxsB,mysB,m\u03b8sB)\u304b \u3089\u4efb \u610f \u306e \u76ee\u6a19 \u4f4d \u7f6e \u30fb\u59ff\u52e2\n(mxGB,myGB,m\u03b8GB)\u306b\u5230 \u9054 \u3055\u305b \u308b\u554f\u984c \u3092\u8003 \u3048\u308b.\n4. \u30d7 \u30ec\u30fc \u30c8\u52d5\u4f5c \u3068\u64cd \u308a\u6226\u7565\n\u5bfe\u8c61\u7269 \u306e\u904b\u52d5\u3092xn,ym\u5404 \u65b9\u5411\u306e\u4e26\u9032\u904b \u52d5,ZB\u8ef8 \u56de \u308a\u306e\u56de\u8ee2\u904b\u52d5 \u306e3\u3064 \u306e\u6210 \u5206\u306b\u5206 \u3051\u3066\u8003 \u3048,\u5404 \u8ef8\u65b9 \u5411\u3078\n\u306e\u904b\u52d5\u751f\u6210 \u3068\u305d\u308c \u3089\u3092\u7d44\u307f\u5408\u308f\u305b \u305f \u76ee\u6a19\u4f4d \u7f6e \u30fb\u59ff\u52e2\u3078 \u306e\u79fb\u52d5\u8a08 \u753b\u306b\u3064\u3044 \u3066\u8003 \u5bdf\u3059 \u308b.\n4\u30fb1 xm\u65b9 \u5411\u306e \u4e26\u9032\u904b \u52d5\u306e\u751f\u6210 \u5bfe\u8c61\u7269 \u304a \u3088\u3073 \u30d7\n\u30ec\u30fc \u30c8\u304c\u9759 \u6b62 \u3057\u3066\u3044 \u308b\u72b6\u614b \u3092\u8003 \u3048 \u308b(X=0,\u0398=0).\n\u56f34(a)\u306b \u793a \u3059 \u3088 \u3046\u306b,\u30d7 \u30ec\u30fc \u30c8\u306e\u4e26 \u9032\u904b\u52d5X\u306b \u3088\u3063 \u3066\u5bfe\u8c61\u7269 \u306b\u306f\u6163\u6027\u529bmBX\u304c \u4e0e \u3048 \u3089\u308c \u308b.\u6700 \u5927\u6469\u64e6\u529b\n\u03bcmBg\u3092 \u8003\u616e\u3059 \u308b\u3068,\u9759 \u6b62\u72b6\u614b\u306b \u3042\u308b\u5bfe\u8c61\u7269 \u306b\u5bfe \u3057\u3066, \u4e26\u9032\u52a0\u901f\u5ea6mxB\u3092 \u751f\u6210\u3059 \u308b\u305f\u3081 \u306b\u306f,\n(1)\n\u3092\u6e80\u8db3\u3059 \u308b\u3088 \u3046\u306a \u6587 \u3092\u4e0e \u3048\u308c \u3070 \u3088\u3044.\u3053 \u3053\u3067,X>0 \u304a \u3088\u3073X<0\u304c,\u305d \u308c\u305e\u308c\u8ca0 \u65b9\u5411(mxB<0)\u304a \u3088\u3073 \u6b63\u65b9\u5411(mxB>0)\u306e \u4e26\u9032\u904b\u52d5 \u751f\u6210\u306b\u5bfe\u5fdc\u3059 \u308b.\u306a \u304a, \u5bfe\u8c61\u7269\u304c\u30d7 \u30ec\u30fc \u30c8\u306b\u5bfe \u3057\u3066\u76f8\u5bfe\u904b\u52d5 \u3057\u3066\u3044 \u308b\u72b6\u614b\u3067\u306e xm\u65b9 \u5411\u306b\u95a2\u3059 \u308b\u904b \u52d5\u65b9\u7a0b\u5f0f \u306f,\n(2)\n\u3067\u4e0e \u3048 \u3089\u308c \u308b.\u30d7 \u30ec\u30fc \u30c8\u306e\u4e26\u9032\u904b\u52d5 \u8ef8 \u3068ym\u8ef8 \u304c\u76f4\u4ea4 \u3057\u3066\u3044 \u308b\u3053 \u3068\u304b \u3089,xm\u65b9 \u5411\u306e\u4e26\u9032\u904b\u52d5 \u306f,ym\u65b9 \u5411\u306e\nFig. 3 Model for analysis\nFig. 4 Translational manipulations" ] }, { "image_filename": "designv11_61_0000090_0141-0229(81)90010-7-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000090_0141-0229(81)90010-7-Figure1-1.png", "caption": "Figure 1 Kinetics of inversion of sucrose in a flow reactor (dependence of In 1/(1 --x) on retention time, 7). Conditions: adsorbent la (Table 1), pH 5, 30\u00b0C, sucrose concentration (wt %): 1,64.2; 2, 53.7; 3, 37.8; 4, 19.5", "texts": [ " This relation holds only for low concentrations of the substrate (< 10 wt %), when the reaction is not inhibited. In an opposite case, the corresponding equation becomes more compficated. 8 Boundrant and Cheftel found, however, that the kinetics of inversion of sucrose with enzyme adsorbed on to :338 Enzyme Microb. Technol., 1981, Vol. 3, October Inversion of sucrose in a continuous process: J. Hradil and F. Svec Amberlite IRA 93 could be described in terms of the kinetic equation of a first-order reaction: As demonstrated by Figure 1, the dependence of in 1/(1 - x) on the reaction time, r, is linear up to a conversion >50% in the system used too, and equation (1) adequately describes the existing conditions. Kinetic parameters of the process are summarized by Table 4. The rate constant markedly decreases (by as much as two orders of magnitude) if the concentration of the substrate solution is raised three times. This finding demonstrates that simple overall firstorder reaction kinetics are valid only at substrate concentrations up to 50%" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002656_1.3058625-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002656_1.3058625-Figure1-1.png", "caption": "FIG. 1. Relevant details of the core and clamping configuration under consideration.", "texts": [ " During each epoch, every particle is accelerated toward its own personal best as well as the global best position according to: vpi k = w vpi k + c1 rand 1 pi k \u2212 Ui k + c2 rand 2 pi g \u2212 Ui k , 8 Ui k = Ui k + vpi k, 9 where k=1, 2 , . . . ,M, c1 and c2 are positive constants, rand 1 and rand 2 are random functions in the range 0,1 , and w is the inertia weight refer to Ref. 10 . The proposed approach has been numerically implemented and magnetostatic flux density distribution computations were carried out for a single-phase transformer core subject to different mechanical clamping configurations. Relevant details of the transformer under consideration are shown in Fig. 1. Windings were assumed to have height, thickness, core-to-winding spacing, and winding-to-winding spacing of 0.9, 0.15, 0.05, and 0.10 m, respectively. Throughout the simulations, J was assumed to correspond to a magnetization current density of 19 104 A /m2, and the typical nonclamped core B-H relation was approximated by expression 2 with ni and C of 11 and 0.748, respectively. For mechanically clamped ferromagnetic core zones the flux density dependent B-H relation was approximated by Eq. 5 having C of 0", " Different B-H relations for all n B\u0304 values under consideration are given in Fig. 2. Sample results for the computed flux density distributions are shown in Figs. 3 and 4. In order to check the accuracy of the proposed approach, analogous finite-element FE simulations were carried. Within these FE simulations, local core nonlinear B-H characteristics were set by Eq. 10 for corresponding local flux density values computed by the proposed approach. A comparison between computed peak flux density values for the different clamping configurations identified in Fig. 1 using both the FE and proposed PSO approaches is shown in Fig. 5. This figure clearly suggests a reasonable qualitative and quantitative agreement between results computed by both computation methods. The suggested approach was also utilized to compute the expected variation in magnetization current Im and core loss corresponding to the different clamping configurations for the case of voltage excitation. In accordance with Refs. 11 and 4 and from Eq. 11 and 4 , the core loss Wc was estimated from Wc r=1 Pmc Ca rx 2 + ry 2 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001442_135065005x34080-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001442_135065005x34080-Figure1-1.png", "caption": "Fig. 1 A schematic diagram of a ball bearing Fig. 3 Inner race to ball in pure rolling contact", "texts": [ " All balls in the races are separated by an equal angular gap with the help of cage. The interaction of cage between the races and balls is isolated and neglected. 5. The bearing is assumed to operate under isothermal condition, thereby variation of lubricant properties with temperature is ignored. 6. The flexibility of the bearing rings is neglected. Proc. IMechE Vol. 219 Part J: J. Engineering Tribology JET53 # IMechE 2005 at University of Ulster Library on March 24, 2015pij.sagepub.comDownloaded from Figure 1 shows a schematic diagram of a ball bearing containing balls, inner race, outer race, and cage. Points \u2018A\u2019 and \u2018B\u2019 are the centres of inner and outer races, respectively, under loaded condition. The total energy of this system is considered to be the sum of kinetic energy, potential energy, and strain energy of the springs representing contact and dissipation energy due to contact damping. Figure 2 shows the contact model of the ball on races represented by non-linear springs and dampers. Kinetic energy of the system is the sum of individual kinetic energies of each element and can be formulated separately" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002168_isie.2007.4374753-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002168_isie.2007.4374753-Figure3-1.png", "caption": "Fig. 3. Rotors: with double cage A. Spiral sheet rotors B-C", "texts": [], "surrounding_texts": [ "Forming a rotor with spiral shape sheets, distributed in a radial disposition around the shaft, it is possible to generate magnetic fields stay more in the rotor's periphery, inducing peripheral e.m.f, and currents along the same sheets, that are only active in their periphery. The peripheral currents of this rotor have more section to circulate, compared with a normal cage rotor's current The figure 4 shows a developed plain representation of the disposition of the sheets. Instead of being shaped in an angular shape, in order to make a difference between the both zones, one where active currents go through, and the other which is used to receive the possible returning currents (A returning currents proposal). In spite of this, the returning currents can be established in two manners: 1) Option A: Through short-circuit rings. With that kind of construction the only rotor resistance that must be considered, is the one corresponding to the outside of the iron sheets that form the rotor. This is due to the short-circuit ring can theoretically be built in a big section, being its resistance totally despised comparing with the resistance offered by the superficial sheets' layers. This solution is not suitable for manufacturing, however it gives the best results. 2) Option B: This second option, consists of a part of the same sheets being a pathway for back currents that do not generate torque. In this model, the zone of active currents that generate torque are placed in the sheets' periphery. Because of these ones are affected by the magnetic field, leaving enough sheet section as a return way, they do not generate torque currents, without increasing, the resistance of the rotor winding, being the last one of a very high section" ] }, { "image_filename": "designv11_61_0001872_978-1-4020-6366-4_15-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001872_978-1-4020-6366-4_15-Figure5-1.png", "caption": "Fig. 5. The Watt parallel motion linkage.", "texts": [ " Watt combined his straight-line linkage with a pantograph, one link becoming a member of the pantograph. This pantograph mechanism [16], denoted as ABEG, is shown in Figure 4. With this design, the length of each oscillating link of the straight-line linkage was reduced to one-fourth instead of one-half the beam length. The entire mechanism could then be constructed so that it would not extend beyond the end of the working beam. This arrangement soon came to be known as Watt\u2019s parallel motion linkage, denoted as O2ABO4 in Figure 5. Through insight we can detect in this straight-line linkage the birth of a very ordered and advanced synthetic process. The kinematic analysis of the Watt four-bar linkage, see Figure 6a, and the geometry of the path of point M fixed in the coupler link AB (link 3) can be investigated using the method of kinematic coefficients [17]. The vectors that are required for the kinematic analysis of the Watt fourbar linkage are shown in Figure 6b. Modern Interpretation of Main Contribution to Mechanism Design The vector loop equation for the four-bar linkage can be written as \u221a I R2 + \u221a ", " (23b) In general, an arbitrary coupler point of a general four-bar linkage will trace a curve which is described as a tricircular sextic [18\u201320]. However, coupler point M of the Watt four-bar linkage traces a special curve which is best described as a figure-eight-shaped curve, as shown in Figure 8. This curve is commonly referred to as a lemniscate and has two segments that approximate straight lines [21]. By means of the pantograph mechanism (see Figure 4), the path traced by point M \u2032 (see Figure 5) is similar to the path traced by coupler point M. The Watt four-bar linkage was employed some 75 years after its inception by Richards when, in 1861, he designed his first high-speed engine indicator. The Richards indicator, which was introduced into England the following year, was an immediate success, and many thousands were sold over the next several decades. In considering the order of synthetic ability required to design the straight-line linkage and to combine it with a pantograph, it should be kept in mind that this was the first one of a long line of such mechanisms" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001874_ijcat.2007.015267-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001874_ijcat.2007.015267-Figure3-1.png", "caption": "Figure 3 Schematic drawing of core-type P-transducer: 1: steel reflex house; 2: piezoelectric ceramic chap; 3: foil", "texts": [ " When the device revolves to realise online follow-up pressure, we drill a hole in the centre of the transducer and pass a steel roof bar through it. We call the new transducer as the core-type P-transducer. Doing this must guarantee that it does not affect the joint bolt intensity. Considering the disadvantages of M-transducer, e.g. it makes the magnetism turbulent flow easy, we adopt the P-transducer. When we add the output electrical energy from the generator on both ends of the P-transducer, it will cause tension or compress in the corresponding direction, creating mechanical vibrations (Figure 3). Ultrasonic transducer is a device that realises the conversion from sound energy to electrical energy or reverse. One kind of ultrasonic transducer is the M-transducer and the other is the P-transducer. Supposed that the input-end areas of the steel reflex house, piezoelectric ceramic chap and A-alloy sonic head are A2, radius are I2; output-end area of A-alloy sonic head is A3, radius is I3; thickness of steel reflex house is I1, thickness of A-alloy sonic head is I2, thickness of each piezoelectric ceramic chap is I0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002676_20090916-3-br-3001.0025-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002676_20090916-3-br-3001.0025-Figure2-1.png", "caption": "Fig. 2. Line following.", "texts": [ " Steering Equation Steering equation is often described in literature with (1) where r is yaw rate, \u03c8 is heading, \u03c4N commanded yaw torque, and parameters to be identified are yaw inertia, Ir , and drag kr|r|. Ir r\u0307 = \u2212k\u0303r|r|r|r| + \u03c4N \u03c8\u0307 = r . (1) For Charlie ASV, the yaw torque control is achieved by controlling the rudder angle \u03b4 while propeller revolution rate n is kept constant, i.e. \u03c4N = n2\u03b4. The dynamic parameters in (1) have been identified in Caccia et al. (2006). The identification experiments have also shown that the sway speed can be neglected. Line Following Equations The line following approach is shown in Fig. 2. The aim is to steer the vehicle moving at surge speed ur in such a way that its path converges to the desired line. If \u03b3 is orientation of the line that should be followed, a new parameter \u03b2 = \u03c8\u2212\u03b3 (vehicle\u2019s orientation relative to the line) is defined. Having this in mind, the line following equations (2) - (5) can be written, where \u03bd is drift due to sea current. r\u0307= \u2212 kr|r| Ir r|r| + 1 Ir \u03c4N (2) \u03c8\u0307= r (3) \u03b2\u0307 = r (4) d\u0307= ur sin\u03b2 + \u03bd (5) The nonlinearities of the line-following model appear in (2) and (5)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002739_s1052618808040092-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002739_s1052618808040092-Figure4-1.png", "caption": "Fig. 4.", "texts": [ " Under conditions of individual manufacturing, both sides of the wheel tooth are processed, as a rule, from one setting. In this case, the rolling ratio ir and tool radial setting U are calculated. The parameters of the rough setup of the machine for pinion tooth processing are calculated in the same way. For finishing processing of each side of a tooth, several subsetups are included in the rough setup: axial (\u2206A) and hypoidal (\u2206E) displacements of the piece, table displacement (\u2206B), variation of the radial setting (\u2206U) of the cutter head (Fig. 4: the scheme of machine setups in bevel gear processing). The rolling ratio ir = \u03c9t/\u03c91 = z1/zm is also changed; where \u03c9t is the angular velocity of rotation of the generating tool wheel, \u03c91 is the angular velocity of rotation of the pinion, and zm is the machine tooth number of the generating tool wheel (Fig. 4). If it is necessary to cut the teeth of the optimized bevel gear using standard cutter heads with a half-sum of the profile angles of the external and internal cutters of the cutter head \u03b1n = 20\u00b0, in technological synthesis the concept of manufacturing cones must be introduced. By manufacturing cones are meant those cones that are coaxial with the initial cones over which the plane generating tool wheel with a standard cutter head ha* haw* 376 JOURNAL OF MACHINERY MANUFACTURE AND RELIABILITY Vol. 37 No. 4 2008 AKIMOV should roll during tooth processing. The angles of the manufacturing cones \u03b4si (i = 1, 2) are related with the angles of the initial cones of the gear \u03b4i (i = 1, 2) by the relationship (Fig. 4) When the teeth are processed on machines with a noninclinable spindle, i.e., on machines operating according to the scheme of a plane-vertex generating tool wheel, the ratio of rolling motion isr is determined by the relationship (5) where \u03b8if is the angle of the dedendum of the pinion and the wheel (i = 1, 2). Using the rolling ratio determining according to Eq. (5) the both sides of the wheel teeth are processed and the subsetups needed for processing of each sides of the pinion teeth are determined by the known means, for example, with the help of the EXPERT software [4, 5]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002953_ijsurfse.2009.026607-Figure15-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002953_ijsurfse.2009.026607-Figure15-1.png", "caption": "Figure 15 Engagement of an involute gear pair with radial eccentricity", "texts": [ " Even though the difference of amount of grating disk eccentricity was over 30 \u00b5m, two measured TE curves agreed with each other very well. It is clearly found that the elimination of the influence of grating disk eccentricity by the opposite reading was successfully performed in high accuracy. To validate the measurement, TE should be calculated and compared with each other using the exact analytical solutions. Here, the equations, which can calculate TE curves with gear eccentricity, are derived shortly. Figure 15 shows an involute gear pair, which has the radial eccentricities of each gear in two dimensions. After rotating each gear about the centre of rotation F by some angles A , two dashed base circles move to the corresponding solid base circles, respectively. To derive the analytical equations more easily, we can redraw this relation. Figure 16 shows gear engagement in the different Cartesian coordinate; the x-axis is always corresponding to the line of action. For the engagement of an involute gear pair, it is easier to understand movement of the gear pair through the line of action" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003865_ultsym.2013.0217-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003865_ultsym.2013.0217-Figure1-1.png", "caption": "Figure 1 the sketch of the definitions of angles and .", "texts": [ " This powerful filter estimates the past, present and even the future state of a system. It is therefore an excellent candidate to predict the position of a needle in 3D US. To fix a line in 3D space, three parameters are needed. According to the RANSAC algorithm, the parameters chosen are two direction angles (the angle between the needle and 844978-1-4673-5686-2/13/$31.00 \u00a92013 IEEE 2013 Joint UFFC, EFTF and PFM Symposium the plane xoz), (the angle between the needle and the positive direction of the z-axis) and the tip position (needle end). Figure 1b gives the definition of these two angles. During the insertion procedure, the needle moves along its axis direction. However, since the US probe is held and manipulated manually, there might also be a rotational movement and even a movement along the z-axis. As a result, the two angles , are included in the measurement vector. Moreover, for the same reasons, the inserting speed might change from time to time, so the velocity of the needle tip is also included in the measurement vector. Because the needle is actuated by hand, there is no control vector or control matrix, so the updated equations for the Kalman filter in this case are: 1 \u02c6 \u02c6 k k X FX (1) 1 T k k P FP F Q (2) with ,[ , , , ]Tt tv v X p v , 2 2 2 2 2 6 2 2 2 2 2 6 3 4 3 3 3 3 3 4 3 3 3 3 dt dt I I 0 0 I 0 F 0 I I 0 0 I , [ , , ]t t t tx y z p is the coordinate of the needle tip, [ , , ]t tx ty tzv v v v is the velocity vector of the needle", " The axis accuracy axis is defined as the maximum Euclidean distance from the two true end point of the needle to the estimated axis, as in 1 tip 2max N Q , N Qaxis (8) To evaluate the proposed method, the mean error and the standard deviation (STD) of the error of RANSAC and ROI + RANSAC + Kalman algorithms are compared together. The mean and STD of error are calculated as: \u02c6[ ]rX E X X (9) 1 2 2 1 1 \u02c6( ) 1 n i i STD X X n (10) Figure 4 shows the mean error together with the STD of the error of the tip accuracy, the axis accuracy, and the two angles 846 2013 Joint UFFC, EFTF and PFM Symposium , (defined in Figure 1b) with 90 , 73 . The mean error and the STD compared with the ground truth value are calculated. From these sub-figures, it is obvious that the ROI-based RANSAC and Kalman method is more robust than RANSAC algorithm only. Both the mean error and the error STD remain small with the proposed method. The smaller STD supports the stabilization of the proposed method. Table 1 shows the ameliorative percentage of the average STD of the different needle lengths (6 \u2013 25 mm) of the proposed method compared to the RANSAC algorithm alone" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003195_978-3-540-89393-6_21-Figure21.7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003195_978-3-540-89393-6_21-Figure21.7-1.png", "caption": "Fig. 21.7 Brushed DC pager motors (a), from left to right: Seiko 2.8 mm (discontinued) (0.16 g), 4 mm (0.51 g), 6 mm (1.58 g), 7 mm (2.70 g). Small brushless motor (1.4 g) from WES-Technik (b), Martin Newell\u2019s custom-built brushless motor (45 mg) (c). Figure (c) reprinted with permission from M. Newell", "texts": [ " It is unlikely that they will be miniaturized below the 1 g range, however, since they require a reservoir for hydrogen that has to be small yet still refillable. Miniature flying robots generally require a main power source to remain in the air, and most of them use rotary motors for this task, either to spin a propeller or to flap wings. Different types of rotary motors exist today, including DC, brushless, or ultrasonic (piezo) motors. Traditional brushed DC motors are the most common type of electric motor in the hobbyist market, can be very efficient (providing up to 300 W/g), and are commercially available in a variety of sizes (Fig. 21.7a). Their main advantage is their ease of use, since they can be driven and regulated simply by using a constant voltage. It should be noted, however, that these mass-produced motors are meant for pagers, not flying robots. To produce significant thrust they must be run continuously at overvoltage, which quickly damages the motors and reduces their useful lifetime. These motors also do not scale down well below around 5 mm of diameter. The latest advances in magnets have improved their performance, but short rotor-stator gaps 8 See: http://www", " Brushed motors perform at maximum efficiency around 10,000 rpm, and thus a gearbox is often used to reach the desired propeller speeds of below 1,000 rpm (see Sect. 21.3.3.1). Gearboxes add some weight and sink some of the power produced by the motor, but allow for a bigger, more efficient rotor to be used, and can thus increase the overall efficiency of the powertrain (see Sect. 21.3.3). Finding an optimum between motor size, gear reduction, and propeller size is important when using DC motors, which requires testing and can be limited by the material that is commercially available. Brushless DC (BLDC) electric motors (Fig. 21.7b) have recently revolutionized the hobbyist market due to their generally increased power to weight ratio compared with their brushed counterparts (up to 40 times higher as shown in Chap. 20). The lack of a mechanical commutator translates to higher efficiency due to the lack of friction between motor and stator. Better cooling and thicker wires allow more current in the coils, while multiple poles reduce the rotation speed which permits BLDC motors to be used without a gearbox, saving additional weight", "3 g, however, and given the trend in miniaturization of electronics, the weight of the control electronics will likely not be a bottleneck in the development of BLDC motors. BLDC motors\u2019 second disadvantage is their initial cost and difficulty in manufacturing. Current motors on the market are mostly hand-wound, and motors below 5 g are difficult to manufacture and suffer in precision and efficiency due to inaccuracies in construction. Current state of the art includes commercially available motors in the 1 g range11 (up to 20 g thrust, Fig. 21.7b), though Martin Newell\u2019s home-made 45 mg motor12 (Fig. 21.7c) demonstrates what can be achieved with some tricks such as gluing the controller to the motor and not using a casing. New manufacturing techniques should continue decreasing the size of BLDC motors. The predominant power source used in large aircraft is not the electric motor but the gasoline combustion engine or turbine. Attempts have been made at miniaturizing this technology in the form of microturbines 10 Examples can be found at: http://www.microinvent.com. 11 Examples can be found at: http://www" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002674_imece2009-11222-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002674_imece2009-11222-Figure3-1.png", "caption": "Figure 3: \u03b8 for each face is determined by the line from the centroid and perpendicular to the face relative to the x-axis. From this the rotation tranformation is determined for the normal and tangent directions.", "texts": [ " After new stresses for each face have been obtained, application of Newton\u2019s 2 nd Law leads to semidiscrete differential equations governing the cell\u2019s velocity change. A suitable temporal discretization of these equations yield the final rule set. The new displacements are stored until all cells have been updated in what amounts to a double buffering technique suitable for parallelization. As mentioned each face has a unique orientation that the strains are calculated for. To facilitate strain computations in the twodimensional case, an angle \u03b8 is set as the angle between the xaxis and the face normal (-\u03c0 \u2264 \u03b8 \u2265 \u03c0) as shown in Figure 3. The direction perpendicular (normal) to the face (defined as en) is in the \u03b8 direction. The direction parallel (tangent) to the face (et) is in the \u03b8 + \u03c0 /2 direction. Note that the calculation for the strains and stresses are like that of a square element as per typical mechanics; however, each face is independent. Hence, the shear stress along the face does not relate to the shear stress along any other face. However, in order to satisfy equilibrium it is a requirement that the shear and normal stress on a face be the same for both neighbors that share the face" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003646_s12283-012-0108-5-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003646_s12283-012-0108-5-Figure3-1.png", "caption": "Fig. 3 Free body diagram of the experimental setup", "texts": [ " The asphalted bottom plate was built and stored at room temperature for 2 months before the measurements were executed. An asphalt surface can be very aggressive and rough when it is new. Measurements on such a surface do not reflect the conditions that roller skiers most frequently train on. Thus, the surface was mechanically machined with a grinding tool to remove the most aggressive areas. 2.2 Mechanics of the roller skis There is a schematic sketch of the experimental setup in the free-body diagram in Fig. 3. Horizontal equilibrium, for the Camber-Ski, when measured without any contact between the Rs and the forward wheel, shows that: S Ff Fr Fl \u00bc 0 \u00f01\u00de In the situation when the Rs came into contact with the forward wheel, the force registered in the load cell gives the equation: S 0 F 0 f Fr Fl \u00bc 0 \u00f02\u00de With the static friction coefficient (lS) defined as the ratio of the resisting force to the normal force (N) on the wheel with the grip function (Nf or Nr), the following relationship was established: lS \u00bc S 0 S N \u00f03\u00de The individual normal forces of the forward (Nf) and rear (Nr) wheel were calculated as: Nf \u00bc mg\u00f0l1 l3\u00de \u00fe F\u00f0l1 l2\u00de S 0 h l1 \u00f04\u00de and Nr \u00bc mg l3 \u00fe F l2 \u00fe S 0 h l1 \u00f05\u00de where the moment of F1 was negligible" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001311_bf00882589-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001311_bf00882589-Figure1-1.png", "caption": "Fig. 1", "texts": [ "4) i t i s n e c e s s a r y to d e t e r m i n e all the s ingu la r points of the complex func - t ion u . . I t is evident f r o m the e x p r e s s i o n s (2.3) that the in t eg rand funct ion has a s e c o n d - o r d e r pole at the point p = 0 and b r anch points at p = 0 and p = - , o . Since the i n v e r s i o n t h e o r e m is appl icable to mul t ivalued funct ions only on the f i r s t shee t of a R i e - mann su r f ace (0 < a r g p ~ 2~r), we m u s t choose the c losed con tour of in tegra t ion as shown in Fig. 1. By J o r d a n , s l e m m a the in t eg ra l s along the cu rves c R tend to z e r o fo r R --- ~o provid ing that t > (2kl + x )c~ 1, t > [2(k + 1 ) / - x ] c ~ . Using the fundamenta l r e s idue t heo rem, we can wr i te the e x p r e s s i o n fo r u(x, t) in the f o r m Here F k = exp [s (2tet + x) c~tR v cos 0 v - - st] sin [s (2kl + x) c~.'R v sin 0vl; G k = exp is 12 (k + 1) l - - xl c~ l i l~ , COS 0~. - - St]' \u2022 sin {s [2 (k + 1) 1 - - x] c~'R v sin 0v}; yk = t - - ( 2 k l + x ) ~ t ; z, = t - - [ 2 ( k + 1) / - - x ] ~ ; 1 l 1 R, = r " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002650_14644193jmbd202-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002650_14644193jmbd202-Figure4-1.png", "caption": "Fig. 4 Phase space partition: (a) impacting chatter and (b) stick motion", "texts": [ " The corresponding separation boundaries for stick motion are defined as \u2202 (i) 12 = \u0304 (i) 1 \u2229 \u0304 (i) 2 = \u23a1 \u23a2\u23a2\u23a3(x(i), x\u0307(i) ) \u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223 \u03d5 (i) 12 \u2261 x\u0307(i) \u2212 Rx\u0307(i\u0304) 2+(tm) = 0 x(i) \u2212 Rx(i\u0304) 2+(tm) = d 2 , tm \u2208 (0, \u221e) \u23a4 \u23a5\u23a5\u23a6 \u2202 (i) 21 = \u0304 (i) 1 \u2229 \u0304 (i) 2 = \u23a1 \u23a2\u23a2\u23a3(x(i), x\u0307(i) ) \u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223 \u03d5 (i) 21 \u2261 x\u0307(i) \u2212 Rx\u0307(i\u0304) 2\u2212(tm) = 0 x(i) \u2212 Rx(i\u0304) 2\u2212(tm) = d 2 , tm \u2208 (0, \u221e) \u23a4 \u23a5\u23a5\u23a6 \u2202 (i) 23 = \u0304 (i) 2 \u2229 \u0304 (i) 3 = \u23a1 \u23a2\u23a2\u23a3(x(i), x\u0307(i) ) \u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223 \u03d5 (i) 23 \u2261 x\u0307(i) \u2212 Lx\u0307(i\u0304) 2+(tm) = 0 x(i) \u2212 Lx(i\u0304) 2+(tm) = \u2212d 2 , tm \u2208 (0, \u221e) \u23a4 \u23a5\u23a5\u23a6 \u2202 (i) 32 = \u0304 (i) 2 \u2229 \u0304 (i) 3 = \u23a1 \u23a2\u23a2\u23a3(x(i), x\u0307(i) ) \u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223 \u03d5 (i) 32 \u2261 x\u0307(i) \u2212 Lx\u0307(i\u0304) 2\u2212(tm) = 0 x(i) \u2212 Lx(i\u0304) 2\u2212(tm) = \u2212d 2 , tm \u2208 (0, \u221e) \u23a4 \u23a5\u23a5\u23a6 \u23ab\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23ac \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23ad (19) where \u0304(i) \u03b1 is the closure of (i) \u03b1 (i = 1, 2) and (\u03b1 = 1, 2, 3). In a similar fashion, for the i\u0304th gear, the domains (i.e. (i\u0304) 3 , (i\u0304) 2 , and (i\u0304) 1 ) and correspond- ing boundaries (i.e. \u2202 (i\u0304) 32, \u2202 (i\u0304) 23, \u2202 (i\u0304) 21, and \u2202 (i\u0304) 12) can be defined. The partitions of phase plane for impacting chatter and stick motions in the absolute frame are sketched in Fig. 4, and the domains and boundaries are presented. In Fig. 4(a), the shaded domain is sketched and labelled as (i) 2 for the free-flying and impacting chatter motions of the two gears. The two non-passable boundaries for the chatter impacts are presented by the dash-dot curves, labelled by L\u2202 (i) 2\u221e and R\u2202 (i) 2\u221e. The impacting time tm represents the impact location on the boundary, and such boundaries are determined by the left and right sides of the i\u0304th oscillator. For the motion continuity of the gear transmission system, the transport law is utilized. Herein, the transport law is the simple impact law. In Fig. 4(b), the two domains for the stick motion are (i) 1 and (i) 3 . The subdomain for free-flying motion is still (i) 2 . The boundaries for the onset and vanishing of stick motion are sketched by the two dash-dot lines, and the switching times tm mark the locations for the appearance and disappearance of stick motions. The hollow and solid circular symbols represent the onset and vanishing of stick motion, respectively. For these boundaries, under certain conditions, the motion can pass through the boundary from one domain to an JMBD202 \u00a9 IMechE 2009 Proc" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003917_icsem.2010.22-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003917_icsem.2010.22-Figure1-1.png", "caption": "Figure 1 the division of tooth surface of toroidal worm", "texts": [ " Based on the helical ne of the primal worm, the deviation between helical line of the current worm and the primal worm is the modification object. In other words, based on the primal worm, expanding the helical line of the current worm shapes a modification Through the study on the meshing relation between modification curve and the worm, the modification nciple of toroidal worm can be studied. The tooth surface of worm along the direction of tooth ght consists of numerous modification curves. Fig. 1 shows that every modification curve represents the tooth file of worm on the special position. So, to study odification curve is actually to discuss the tooth profile ong the helical line. The reference circle lies on the middle he working tooth, so its modification curve is most representative and usually regarded as the representative of the tooth surface of worm. In the following sections, the positive direction? Fig. 2 shows that the point M and N all lie on the original tooth surface. To point N, there may be two directions after modifications which are de and \u2013de" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001095_iros.2004.1389778-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001095_iros.2004.1389778-Figure1-1.png", "caption": "Fig. 1. HlROBO SF40 autonomous helicopter and bodyfixed coordinates: main rotor diameter = 1,790 mm, maximum lift = 17.5 kgf, quipped with 4 k c gasoline engine.", "texts": [ " Good consistency between experimental data and simulation data demonstrates the high accuracy of the models and the adequacy of the modeling method. I. Introduction A small-scale unmanned helicopter offers many advantages, including low weight and the ability to fly within a narrow space. From the viewpoint of flight outside the operator's range of direct observation and reduced labor: autonomous control technolom is indispensable, and research of this type has been popular in recent years. Our research group has been conducting research on autonomous flight control of a hobby-class small-scale unmanned helicopter (shown in Fig.1) for four years, and we have already completed development of basic control system hardware [l]. In this paper we describe autonomous hovering control and guidance control of the above-mentioned small-scale unmanned helicopter by identified mathematical model-based H , control theory. Modeling methods are roughly classified into three categories: 1) the anall-tical approach, 2) the system identification approach, and 3) the numerical approach. Each of the above-described approaches has advantages and disadvantages" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002210_iecon.2007.4460261-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002210_iecon.2007.4460261-Figure2-1.png", "caption": "Fig. 2. Position propagation of the mobile robot.", "texts": [ " On the two dimensional X -Y Cartesian coordinates, position of the mobile robot is described by XR (t) and YR (t) while the orientation is represented as 0R (t). P =KR YR R]'T where p = [XR YR OR ] Now the kinematics model of the represented as YR =cos ORV2, YR =sin2RV2 OR = V2 - (3) mobile robot can be (4-a) (4-b) (4-c) Kinematics analysis aims at the proper velocity assignment to each wheel to drive the mobile robot to a desired position and orientation [10, 12]. B. Position Propagation In previous session, we studied that the states of the mobile robot with differential driving mechanism are changing according to the two wheel velocities. In Fig. 2, when the mobile robot is moving from A where the robot is located on robot = [XR YR 6R ]kT at time k to C where the position is on p =+n [XR YR TR]k+n at time k + n As shown in Fig. 1, a mobile robot with differential driving mechanism has two wheels on the same axis, and each wheel is The state transition of the mobile robot can be described in terms of currents state and inputs as follows: XR = XR+T 2R COR s6k (5-a) (5-b) YR YR + T R sink y} = y} + T s'nwhere2 Or OR - R' I where T is the sampling period" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000295_2005-01-0016-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000295_2005-01-0016-Figure2-1.png", "caption": "Figure 2 - Overview Truck Layout", "texts": [ " The paper is organized in three sections: an electrification overview, an integration overview, and results. The first will provide details on the components selected and installed. The second focuses on control, communication, and coordination between the subsystems. The third gives details of test results and analysis of energy savings. Original Equipment Manufacturer (OEM) mechanical engine loads were removed and replaced with electric components. The engine cooling pump and the AC compressor are the main loads removed. Figure 2 illustrates the layout of the truck. The fuel cells and hydrogen tanks are enclosed in weather-tight boxes with appropriate ventilation to minimize hydrogen accumulation. Engine Cooling Pump A high-efficiency 42V brushless DC pump replaced the belt driven engine cooling pump. The electric pump offers variable speed and flow control based on the engine thermal load, not engine speed. The cooling system also utilizes a three way diverter valve replacing the wax pellet controlled thermostat. A coolant bypass flow circuit was added between the diverter valve and the inlet to the water pump" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000184_bf00794932-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000184_bf00794932-Figure1-1.png", "caption": "Fig. 1", "texts": [ " This means that the p r o c e s s is no longer a x i s y m m e t r i c a l , which hinders theore t ica l ana lys i s and compute r s imulat ion, as well as in te rpre ta t ion of exper imenta l data . The re has not been any s y s t e m a t i c study of the pene t ra t ion of p la tes at angles to the no rma l . He re this gap is filled in r e l a - tion to h igh-speed impact . 1. Cons ider the col l is ion of a cyl inder of height s i m i l a r to the d i a m e t e r of the base with a plate of th ickness 5. Let the impac t veloci ty v 0 coincide in d i rec t ion with the axis to the cyl inder and const i tute an angle with the su r f ace of the pla te (Fig. l a ) . Pa r t s b and c of Fig. 1 r e p r e s e n t the in te rac t ion with the pla te . The broken l ines in Fig. lb show the posi t ions of the longitudinal shock waves in the plate and cyl inder , whose p a r a m e t e r s a r e de te rmined by the normal component of the coll is ion veloci ty, which is v 0 sin ~; if ~ =90 ~ the wave phenomena caused by penet ra t ion of the plate and disrupt ion of the injected pa r t and of the cyl inder i t se l f have been examined in detai l [1, 3]. If ~ ug0 ~ the phenomena a re compl ica ted by the high shea r s t r e s s e s occur r ing in the body and the plate , whose speeds a r e much lower than those of the longitudinal waves , and t he re fo re they have a substant ia l effect on the final s tages of the p r o c e s s . This m a y r e su l t in changes in the c h a r a c t e r i s t i c s of the f r agmen t s re la t ive to those in normal incidence. The longitudinal shock wave r eaches the r e a r f r ee su r f ace of the plate f i r s t at a point opposite the point of f i r s t contact between the cyl inder and the plate (Fig. 1). The resul t ing cen te red n e g a t i v e - p r e s s u r e wave gives r i s e to mot ion of the pla te normal to the su r face . If the resul t ing tens i le s t r e s s e s a r e l a rge enough and exis t fo r a sufficient t ime , then pa r t of the plate will b reak away. On the o ther hand, the deforming and de - composing pa r t i c l e continues to move in a d i rec t ion c lose to that of the incident veloci ty . Consequently, the flow of f r agmen t s fo rmed f rom the plate moves main ly no rma l to the su r face , and this is s epa r a t e f r o m the flow of f r agmen t s f r o m the impact ing body, which follows the or iginal l ine" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000842_ccece.2004.1349733-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000842_ccece.2004.1349733-Figure2-1.png", "caption": "Fig. 2: Stator flux linkage vector diagram", "texts": [], "surrounding_texts": [ "magnet synchronous motors is very important as it affects the machine pegormance. For steady state operation, it can help to decrease stator current or machine losses: however, in order to get a fast torque response during motor start up, one needs to change both stator flux angle and magnitude. In this paper, a method is presented to find the optimal voltage vectors to change both flux angle and its magnitude which results in high dynamics for torque response. The closed form formula given for optimal voltage vector is derived based on maximizing torque change in each sampling time. The simulation results show that the presented method is better than using conventional DTC. In other word, it is shown that the voltage vectors derived from switching table for conventional DTC are not the optimal voltage vectors for the situation in which there is a step change in torque command.\nKeywords: Interior Permanent Magnet (IPM) synchronous motors, Direct Torque Control (DTC).\n1. INTRODUCTION\nIRECT Torque Control (DTC) is considered as one of D the best alternatives for the motor drive designers in order to get a fast torque response; especially when torque control instead of speed or position control is control objective. Besides high torque dynamics, it is well known for being robust to the motor parameters change, except the stator resistance [l]. For the first time, its application in Permanent Magnet Synchronous (PMS) motor was discussed in [ 2 ] . The authors also investigated the effect of the flux reference on torque dynamics; a bigger flux reference results in a bigger achievable torque but a slower torque response. They also found a maximum value for the flux reference; however, they did not present any strategy to select the flux command during torque transitions, i.e. there is a step change in torque command like motor start up.\nThe flux command is kept constant and usually equal to the nominal value for the speeds under the nominal speed;\nhowever, it does not give the maximum torque to amp ratio. For the speeds over the nominal speed, flux weakening is applied to the vector control of an IPM synchronous motor [3]. An algorithm has also been proposed to select flux command for two different objectives: Minimization of the stator current and Minimization of the total motor losses [4]. The proposed algorithm is proper for steady state operation and still does not provide any improvement to torque dynamics because it determines a constant flux. A qualitative description is in [5] to achieve a fast torque response in PMS motors. However, no mathematical derivation or system synthesis is attempted.\nIn the current paper, a method, independent of the flux command, is proposed to get a fast torque response for Interior Permanent Magnet (IPM) motors start up from standstill. The proposed method looks for an optimal voltage vector which results in a maximum torque change in each sampling time.\nIn section 2, the model of an IPM machine besides fundamentals of DTC are presented. To show the relationship between the flux command and torque dynamics, the effect of choosing different flux commands for same machine is investigated in section 3. In section 4, a new strategy is presented to calculate the optimal voltage vector in order to get a rapid torque dynamics. The presented method is evaluated by simulating an IPM motor and the results are compared with the ones obtained by traditional DTC in section 5. Finally, section 6 summarizes the main contributions of this paper.\n2. IPM MACHINE MODEL AND DTC FUNDAMENTALS\nUsing d-q transformation, the voltage equations of an IPM machine in rotor reference frame are as follows:\n@ e 2, dad vd = R , id +-- dt\nd A dt v , = R , i , + ~ + w , A ,\nCCECE 2004- CCGEI 2004, Niagara Falls, May/mai 2004 0-7 803- 825 3-6/04/$17.00 02004 IEEE\n- 1673 -", "where A, = L, id + 1, , 2, = L, i, and stator flux linkage is\na, = (1; + 1: ) f . The corresponding equivalent circuits are\nshown in Fig. 1 It has been shown in [2] that the electromagnetic torque in an IPM machine can be regulated by controlling the magnitude and angle of stator flux linkage (or load angle 8 ) [see Fig. 21. This can be performed by applying proper output voltage vectors of an inverter to the machine. There are 6 nonzero voltage vectors and one zero voltage vector for a two level inverter as depicted in Fig. 3 . They can be represented by:\nSa, Sb and S, are used to show the state of each leg in the inverter which are either 0 or I . They are 0 when the leg is connected to zero and 1 when connected to the DC bus voltage VDc. It has also been proved that the torque dynamics is dependent on the speed of rotation of stator flux linkage with regard to magnet flux linkage. In other word, to get a fast torque response, one should increase the load anglesas quickly as possible. There are two different inverter switching tables introduced for IPM machines [6]. One of\nthe switching tables is suitable for low speed operation. Table I is such a switching table which does not use the zero voltage to decrease the machine torque. After sampling motor currents and voltages, the stator flux magnitude and machine torque are estimated as:\n3 T , = --P 2 [a, I , - a, I , ]\n(4)\nand compared to their set points a] and T,' respectively. D and Q are used to represent the quantities in stationary reference frame. Based on the estimated torque and flux errors and the stator flux region number n (n: 1-6 as depicted in Fig. 3 ) the switching table generates proper switching commands for the inverter.\n3. FLUX COMMAND EFFECT ON MACHINE PERFORMANCE\nThe authors in [2] have discussed the relationship between the amplitude of stator flux command and derivative of electromagnetic torque (or torque dynamics). They also showed that in order to reach high dynamics for torque response, the flux command should be selected properly\n- 1674 -", "during the torque transient. However, no attempt has been made to determine the method of flux selection. In this section the result of [2] are recalled and further analyzed to provide a basis for the optimal voltage selection presented in the next section. Fig. 4 shows the electromagnetic torque with respect to 8 of an IPM machine with the specifications as in appendix I for 4 different flux commands. It is obvious that the smaller values of the flux command result in linear and fast torque responses with respect to 6. However, the maximum achievable electromagnetic torque is smaller which is a disadvantage for the machine especially in the case machine is heavily loaded and needs bigger electromagnetic torque to start up besides a fast torque response. So it is important to vary the stator flux properly during torque transition to overcome the mentioned shortcoming. In the next section an effective strategy is proposed to find the optimum voltage vector to vary both stator flux angle and its amplitude in order to obtain maximum torque increase in each sampling period.\n4. OPTIMAL VOLTAGE VECTOR\n4.1 Optimal Voltage Vector Selection\nThe torque of (5) can also be represented in rotor\n(6)\nThen the torque differential with respect to id and i, is obtained as:\nreference frame as: 3 2 T, =- P [Am + (Ld - Lq ) i d ] i q .\n~ T ~ = - - P { ( L , 3 - L q ) i q d i d +[A,+(L, - ~ , ) i , ] d i ~ } (7) 2\nDuring each sampling period T,, (7) can be approximated by: where: AT,=AAi , + BAiq (8)\nA=--P(L , 3 - L q ) i q ' (9) 2 3 2\nB =--P [A, + ( L , - L , ) id I' By neglecting the voltage drop across the stator resistance, the following equations are easily obtained from the machine voltage equations (l), (2) along the d and q axes:\n~i~ - ' d + E d T, 9 Ed =we aq 9 (11)\n(12) Ld\nL,\nv - E Aiq = U T s 7 E, =We a d *\nSubstituting (1 1) and (12) into (8) yields: AT, = f (vd ,vq)=avd +bvq +c (13)\nwhere:\nA very fast torque response requires that A T , has a maximum value all the time. This is possible by applying an optimal voltage vector to the machine in each sampling period. This voltage is found here in terms of its d and q components where:\nvq = J+ 2 >' - v,' .\nSubstituting (15) into (13), the optimal voltage is the solution of the following equation:\na ATe ( V d 1 - af (Vd 1 -o (16) a ' d a ' d\nwhich in connection with (15) yields:\nEquations (17) and (18) represent the d and q components\nof the optimal voltage vector V , which has been derived based on maximizing torque change in each sampling time T,. In the next section, Space Vector Modulation (SVM) is used to apply the calculated optimum voltage to IPM motors.\n-*\n4.2 SVM to Apply Optimum Voltage Vector\nThe optimal voltage vector is obtained in rotor reference frame above. However, DTC is basically performed in a stationary reference frame. In this section, a reference transformation from the stationary reference to the stationary frame is proposed with a minimal requirement to the rotor position.\nBy using (17), (1 8) and the initial rotor position 6,. , the optimal voltage vector components in a stationary reference frame are found as follows:\n(19) COSB, -sine, [:]=[sin e, coser j[ a] During torque development duration which typically takes shorter than IO rnsec, the rotor position 0, does not change too much and can be assumed as a fixed angle. Therefore, by knowing the initial position of rotor, an encoder is not required. In practice, once the coefficient matrix in (19) is calculated, it can be stored in the memory and used in each sampling time; thus the dq to DQ transformation does not take much time for processor. Now, SVM which is basically\n- 1675 -" ] }, { "image_filename": "designv11_61_0002042_j.engfracmech.2007.03.025-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002042_j.engfracmech.2007.03.025-Figure4-1.png", "caption": "Fig. 4. Finite element model of the zone around the crack.", "texts": [ " In order to obtain the displacement field under the tooth surface, that pressure distribution is firstly schematized as a set of finite number of point forces normal to the free-surface of the half-space; then, using the Boussinesq theory [18,19], the displacement components induced by each of those point loadings are analytically computed; by adding the contribute of each point loading, the displacement everywhere in the half-space is known. With the aim to evaluate the SIF for Modes I, II and III along the crack front, the displacements obtained by the previous step are applied as boundary condition to a 3D finite element model of the zone surrounding the crack with radius equal to a (Fig. 4). Brick elements with 20 nodes and second-order shape functions were used. Special elements were also used to simulate the contact between the crack faces, avoiding their overlapping. The friction between the crack faces can be included in the model. This model allows a very refined mesh near the crack front, where the 1/4 point technique was used to better simulate the stress singularity according to LEFM. The height and the radius of the model (H and R of Fig. 4) were chosen to avoid that the displacement field of the un-cracked half-space was influenced by the presence of the crack, that is to say to have exact boundary conditions for the FE analyses. After a trial and error procedure H and R were slightly modified with respect of the results included in [20] and were chosen equal to 2a and to 4.5a without appreciable modification of the results (compared in terms of displacements with the ones of a simple, not crack focussed FE model). In this way, it was possible to consider cracks closer to the free-surface than in [20]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003282_978-0-387-74244-1_11-Figure11.5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003282_978-0-387-74244-1_11-Figure11.5-1.png", "caption": "FIGURE 11.5. A two-wheel model for a vehicle with roll and yaw rotations.", "texts": [ "103) Assuming small angles for slip angles \u03b2f , \u03b2, and \u03b2r, the tire sideslip angles for the front and rear wheels, \u03b1f and \u03b1r, may be approximated as \u03b1f = 1 vx (vy + a1r \u2212 zfp)\u2212 \u03b4 \u2212 \u03b4\u03d5f = \u03b2 + a1 r vx \u2212 C\u03b2f p vx \u2212 \u03b4 \u2212 C\u03b4\u03d5f\u03d5 (11.104) \u03b1r = 1 vx (vy \u2212 a2r \u2212 zrp)\u2212 \u03b4\u03d5r = \u03b2 \u2212 a2 r vx \u2212 C\u03b2r p vx \u2212 C\u03b4\u03d5r\u03d5. (11.105) 11.3.3 F Body Force Components on a Two-wheel Model Figure 11.4 illustrates a top view of a car and the force systems acting at the tireprints of a front-wheel-steering four-wheel vehicle. When we consider the roll motion of the vehicle, the xy-plane does not remain parallel to the road\u2019s XY -plane, however, we may still use a two-wheel model for the vehicle. Figure 11.5 illustrates the force system and Figure 11.6 illustrates the kinematics of a two-wheel model for a vehicle with roll and yaw rotations. The rolling two-wheel model is also called the bicycle model. The force system applied on the bicycle vehicle, having only the front wheel steerable, is Fx = 2X i=1 (Fxi cos \u03b4 \u2212 Fyi sin \u03b4) (11.106) Fy = 2X i=1 Fyi (11.107) Mx = Mxf +Mxr \u2212 wcf \u03d5\u0307\u2212 wkf\u03d5 (11.108) Mz = a1Fyf \u2212 a2Fyr (11.109) where \u00a1 Fxf , Fxr \u00a2 and \u00a1 Fyf , Fyr \u00a2 are the planar forces on the tireprint of the front and rear wheels" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003015_cae.20257-Figure11-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003015_cae.20257-Figure11-1.png", "caption": "Figure 11 Procedure for wiring circuit of head light\u2014practical mode. (a) Question and single-choice item; (b) the question for next step; (c) the question for another step; (d) right answer and finish this job. View this article online at wileyonlinelibrary.com.", "texts": [ " There are several steps arranged in this test and the learner has to select an option in each step (three options are listed in general) and finish the procedure. Figure 10a shows the title is \u2018\u2018which one of the correct clipping position of venire?\u2019\u2019. The options are \u2018\u2018left-leaning,\u2019\u2019 \u2018\u2018mean,\u2019\u2019 and \u2018\u2018right-leaning.\u2019\u2019 If the learner selects any wrong step of all, a corresponding error message is displayed as in Figure 10b and requests the learner to do it again. If all steps are right, a message signaling the completion of the inspection will be shown along with the measurement result that the dimension of length is 42.50 mm (as shown in Fig. 10c). Figure 11 shows the head light circuit-wiring practice in which learners have to select an option in each step and finish the procedure. The light is bright when wiring is exact; otherwise, an alarm message is automatically shown and requests learners to do it again. The right wiring procedure is positive terminal of battery! fuse (Fig. 11a)!multi-function switch (Fig. 11b)! relay (Fig. 11c)! terminals of head light (Fig. 11d)! ground. Take another example of practice mode for testing the performance of alternator (includes maximum current and a specified voltage in the vehicle lights are turn on, crankshaft speed is 2500 rpm as well as air conditioning is open). To measure charging-system output, the exact process is illustrated in Figure 12. Connect an ammeter (positive lead, \u2018\u2018\u00fe\u2019\u2019) at the alternator BAT terminal (as shown in Fig. 12a). Follow the operating instructions for the ammeter you are using, link the negative lead of ammeter and positive terminal of battery (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001978_aero.2007.352757-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001978_aero.2007.352757-Figure1-1.png", "caption": "Figure 1. Inertial and helicopter's body frame and the applied aerodynamic forces and moments", "texts": [ " In this research, a six degrees of freedom (DOF) dynamic model has been considered for the helicopter. The fly-bar dynamics, which have an essential effect on stabilizing the model helicopter manoeuvres, are also considered in the modelled aerodynamic forces based on the approach presented in [8]. The position and attitude of the helicopter are determined by the application of four input commands: The lift force TM, the thrust force TT, the roll moment MO, and the pitch moment M0, respectively. These commands are shown in Fig. 1. The body coordinate system of the helicopter is attached to helicopter's Center of Gravity (CG) with x, y and z axis pointing to helicopter's front, right-hand side, and downward, respectively. Neglecting the effect of the tail rotor, helicopter's plane of symmetry is defined as the x -z plane of the body coordinate system. As indicated in Fig. 1, ir and lr are the distances of the main and tail rotor shafts from the helicopters center of gravity in the direction of body coordinate x axis. Helicopter's equations of motion can be written as follow: d PI1 = -tPI (1)dt P PI [RIB (fdB + faB) + fil] (2) C;-B = t - B X IWB (3) WB =Ee (4) In these equations, pI is the position vector ofthe helicopter's CG. The vectors e and -B consist ofthe Euler angles and angular velocities of the helicopter's body, respectively. Also, E represents the transformation matrix between vectors e and AB" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000371_1-4020-3741-4_3-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000371_1-4020-3741-4_3-Figure2-1.png", "caption": "Figure 2: Schematic representation of laminar Taylor-vortex flow (redrawn from Schlichting, 1987).", "texts": [ " Here, secondary flow phenomena are superimposed on the Couette flow and a series of counter-rotating toroidal vortices that are periodic in the axial direction form in the annulus. Each vortex has equal size and equal rotational speed. According to Savas (1985), the dark areas that are observed are the sinks and indicate radial inward motion of the particles. The thin dark lines are the sources and indicate radial outward motion of the particles. The light areas indicate motion in the azimuthal (circumferential) direction. Laminar Taylor-vortex flow is represented schematically in Figure 2 where Ui is the rotational speed of the inner cylinder. With further speed increments, travelling waves that are periodic in the circumferential direction develop and become superimposed on the laminar Taylor-vortex flow, thereby producing a doubly periodic or wavy-vortex flow phenomena (c). With still further speed increments, a modulated vortex flow ensues as a result of higher modes being generated by harmonics of the two fundamental frequencies in the doubly periodic wavy-vortex flow. A discrete spectrum to higher frequencies is achieved with continued speed increments and more complex stable flow structures form, however, on reaching this point these structures are now no longer stable (d)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002487_s102319350810008x-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002487_s102319350810008x-Figure1-1.png", "caption": "Fig. 1. Diagram of the electrochemical cell.", "texts": [ " In all measurements, an Ag (s) | AgCl(s) | KCl (aq, 1 M) reference electrode was used. The auxiliary electrode was made of a Pt wire, 1cm length and 0.5 mm in diameter. Flow Injection Setup The equipment for flow injection analysis included a 10 roller peristaltic pump (UltrateckLabs Co., Iran) and a four\u2013ways injection valve (Supelco Rheodyne Model 5020) with a 50 \u00b5 l sample injection loop. Solutions were introduced into the sample loop by means of a plastic syringe. The electrochemical cell used in flow\u2013injection analysis is shown in Fig. 1. The volume of the cell was 100 \u00b5 l. In all experiments described in this paper, the flow rate of eluent solution was 100 \u00b5 l/s. All of the electrochemical experiments were done using a setup comprised of a PC PIV Pentium 900 MHz microcomputer, equipped with a data acquisition board (PCL\u2013818HG, Advantech. Co.), and a custom made potentiostat. All data acquisition and data processing programs were developed in Delphi 6 program environment. In Fig. 2, the diagram of applied waveform potential during cyclic voltammetric measurements is shown" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003282_978-0-387-74244-1_11-Figure11.3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003282_978-0-387-74244-1_11-Figure11.3-1.png", "caption": "FIGURE 11.3. The force system at the tireprint of tire number 1, and their resultant force system at C.", "texts": [ "39) in (11.30) and (11.33) results in the force system. \u23a1\u23a3 Fx Fy 0 \u23a4\u23a6 = m \u23a1\u23a3 v\u0307x \u2212 \u03c9zvy v\u0307y + \u03c9zvx \u03c9xvy \u23a4\u23a6 (11.40) \u23a1\u23a3 Mx 0 Mz \u23a4\u23a6 = \u23a1\u23a3 \u03c9\u0307xI1 \u03c9x\u03c9zI1 \u2212 \u03c9x\u03c9zI3 \u03c9\u0307zI3 \u23a4\u23a6 (11.41) 11.3 F Vehicle Force System To determine the force system on a rigid vehicle, we define the force system at the tireprint of a wheel. The lateral force at the tireprint depends on the sideslip angle. Then, we transform and apply the tire force system on the rollable model of the vehicle. 11.3.1 F Tire and Body Force Systems Figure 11.3 depicts wheel number 1 of a vehicle. The components of the applied force system on the rigid vehicle, because of the generated forces at the tireprint of the wheel number i, are Fxi = Fxwi cos \u03b4i \u2212 Fywi sin \u03b4i (11.42) Fyi = Fywi cos \u03b4i + Fxwi sin \u03b4i (11.43) Mxi = Mxwi + yiFzi \u2212 ziFyi (11.44) Myi = Mywi + ziFxi \u2212 xiFzi (11.45) Mzi = Mzwi + xiFyi \u2212 yiFxi (11.46) where (xi, yi, zi) are body coordinates of the wheel number i. It is possible to ignore the components of the tire moment at the tireprint, Mxwi , Mywi , Mzwi , and simplify the equations" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002138_iros.2007.4398956-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002138_iros.2007.4398956-Figure1-1.png", "caption": "Fig. 1. Treatment Robot", "texts": [ " INTRODUCTION Protontherapy is a radiotherapy modality that uses protons in order to have a better conformation of the dose around the treated volume. In the considered case, the clinical proton beam line is horizontal and fixed in treatment rooms. Thus, contrarily to conventional radiotherapy and radiosurgery linear accelerators (Linacs) [8] where the photon beam turns around the tumor, the patient has to be moved towards a fixed beam in order to achieve all required fields incidences. This is achieved thanks to a 6DOF robot which can move a bed or a chair where the patient is place, see Fig. 1. Treatments considered here, are fractionated into several ses- sions requiring a daily positioning of the patient. Moreover, to reduce the irradiation of the safe tissues around the tumor, multiple incidences fields are used as illustrated in Fig. 2. Usually, a treatment comprises approx. thirty sessions spread out over 5 to 7 weeks, during which the patient receives two S. Pinault and G. Morel are with the Universit\u00e9 Pierre et Marie Curie - Paris6, Institut Syst\u00e8mes Intelligents et Robotique (ISIR), CNRS (FRE2507), BP 61, 18 route du Panorama, 92265 Fontenay-aux-Roses, France {pinault, morel}@robot" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000777_icmlc.2004.1382286-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000777_icmlc.2004.1382286-Figure1-1.png", "caption": "Figure. 1 The ball-and-beam system", "texts": [ " The population \u201cevolves\u201d by genemting new solutions, mostly based on the better solutions in the population. GAS are par!icularly suited for the evolution of controllers for complex systems, since they only require feedback on the performance of a possible solution. In the present research we apply a genetic algorithm to the development of a feedforward neural controller for the ball-and-beam system. We evolve both the weights and the architecture of the network. 2. The ball-and-beam system The ball-and-beam system, used in [4,5], is illustrated in figure 1. It consists of a beam that pivots at the center point, and a ball that is free. to move along the beam in a vertical plane. The task of the system\u2019s controller is to apply a sequence of torques to dynamically balance the ball from any initial position with any initial speed (within a f 25-degree beam angle). An optimal controller can get the ball stationary and the beam at a horizontal position within a short period of time. Supposing no friction exists in the system, and the ball mass and the beam mass are both uniformly distributed, the ball-and-beam system is described by the following set of Lagrangian equations of motion: 0-7803-8403-~04/$20" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002654_s0263574708004244-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002654_s0263574708004244-Figure4-1.png", "caption": "Fig. 4. Error ellipse contribution converted to a line.", "texts": [ " The ellipse parameters (4a, b) are an estimate of the total deadreckoning induced error contribution to the AGV location measurement a(l) = \u2211 a(l), a(l) \u223c= a(R1, R2, \u03b81, \u03b82, \u03c6e) (4a) b(l) = \u2211 b(l), b(l) \u223c= b(R1, R2, \u03b81, \u03b82, \u03c6e) (4b) and in the following rationale, these expressions will be evaluated (5a\u2013e), where K a(l), K b(l) are error constants, \u03c6 = \u03c60 + 1 2b (R1 \u03b81 \u2212 R2 \u03b82) (5a) \u03c6odom = \u03c60 + R 2b ( \u03b81 \u2212 \u03b82) (5b) \u03c6e =\u03c6odom \u2212\u03c6 = 1 2b [ \u03b81(R \u2212 R1) \u2212 \u03b82(R \u2212R2)] (5c) a(l) = K a(l)[R cos(\u03c6odom)( \u03b81 + \u03b82) cos(\u03c6) +R sin(\u03c6odom)( \u03b81 + \u03b82) sin(\u03c6) \u2212 cos(\u03c6odom)(R1 \u03b81 + R2 \u03b82) cos(\u03c6) \u2212 sin(\u03c6odom)(R1 \u03b81 + R2 \u03b82) sin(\u03c6)] = K a(l) cos(\u03c6e)(R \u03b81 + R \u03b82 \u2212 R1 \u03b81 \u2212R2 \u03b82) (5d) b(l) = K b(l)[\u2212R cos(\u03c6odom)( \u03b81 + \u03b82) sin(\u03c6) +R sin(\u03c6odom)( \u03b81 + \u03b82) cos(\u03c6) + cos(\u03c6odom)(R1 \u03b81 + R2 \u03b82) sin(\u03c6) \u2212 sin(\u03c6odom)(R1 \u03b81 + R2 \u03b82) cos(\u03c6)] = \u2212K b(l) sin(\u03c6e)(R \u03b81 + R \u03b82 \u2212 R1 \u03b81 \u2212R2 \u03b82). (5e) Analysing Eq. (5e), if \u03c6e = 0 rad, then b(l) = 0 m and the ellipse becomes a line (Fig. 4). So the line that passes at point (x1odom, x2odom) with slope \u03c6 is given by Eqs. (6a\u2013b). x2 \u2212 x2odom = tan(\u03c6)(x1 \u2212 x1odom) (6a) \u03c6 = \u00b1\u03c0 2 \u21d2 x1 = x1odom (6b) Nevertheless, for changing heading angle \u03c6, the error ellipse becomes rotated as in (7a, b). The error ellipse contribution becomes a line, but the corresponding overall error ellipse, in general, is not a line: { a \u03c6(l) = a(l) + a(l) cos( \u03c6) + b(l) sin( \u03c6) b \u03c6(l) = b(l) + b(l) cos( \u03c6) \u2212 a(l) sin( \u03c6) (7a) \u03c6 = \u03c6 \u2212 \u03c60 (7b) where \u03c6 is the heading angle change and a \u03c6(l) and b \u03c6(l) are the new error ellipse parameters" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003966_ijvsmt.2012.049432-Figure12-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003966_ijvsmt.2012.049432-Figure12-1.png", "caption": "Figure 12 Obstacles used in simulation (Unit: mm) (see online version for colours)", "texts": [ " This means that, under the harmonic road excitation, road profile acceleration has great effects on the tyre force response which agrees with the results in ShaoPu et al. (2009). From Figure 11, one can notice that the maximum \u0394Fz has a non-linear relationship with velocity (power function, \u0394Fz = aVb). When the velocity is increased, the slope (rate of change) of relationship between z \u0394F and velocity is increased. The simulations were performed by rolling the tyre over obstacles while maintaining a constant axle height. Three different shapes of obstacles are used in the simulation, as shown in Figure 12. Also, three different velocity values (0.5, 25, and 60 km/h) are used. The simulation at 0.5 km/h is performed to represent the quasi-static condition. The tyre has an initial vertical load 4,000 N. Then, the force response is calculated and compared with available data from the literature. The measured quasi-static force response is digitised from Zegelaar and Pacejka (1996) and Schmeitz et al. (2004), then the data were curved-fitted by using piecewise cubic curve fitting method as given in Figures 13 and 14" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003577_iet-pel.2010.0319-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003577_iet-pel.2010.0319-Figure3-1.png", "caption": "Fig. 3 Continuous converter virtual flux vector and the continuous converter voltage vector divided into d- and q-components", "texts": [ " The modelling of modulation with continuous functions is based on determining the modulation in terms of continuous functions a(t), b(t) and zv(t) instead of time-discrete switchings ta, tb and t0. The converter virtual flux vector corresponding to a continuous voltage SV is c cont = cconte jvt = \u222b ucontdt (14) where v is the angular frequency of the continuous converter virtual flux vector. To derive equations for a, b and zv, a synchronous reference frame fixed to the continuous converter virtual flux vector is introduced in Fig. 3. The continuous converter voltage vector is divided into d- and Rotating synchronous dq frame and the stationary ab frame are shown 479 & The Institution of Engineering and Technology 2012 q-components as ucont = ucont,d + jucont,q (15) In this particular synchronous frame, we have c cont = ccont = \u222b ucont,d dt (16) and ucont,d = dccont dt (17) The angular frequency of the virtual converter flux vector is expressed as v = ucont,q ccont (18) The virtual grid flux vector c s is defined as c s = cse jvst = \u222b usdt (19) where us is the grid voltage vector and vs is the grid angular frequency" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003094_iros.2009.5354487-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003094_iros.2009.5354487-Figure4-1.png", "caption": "Fig. 4. Configuration at instant of heel strike.", "texts": [ " (3) These matrices are described in detail elsewhere [2]. If we assume inelastic collisions for the stance-leg exchange and set suitable values for the physical parameters, the robot can exhibit passive dynamic walking on a gentle slope. Let E be the total mechanical energy of the robot, and relationship E\u0307 = \u03b8\u0307 T Su between the mechanical energy and the control inputs holds. The modeling of an inelastic collision is briefly described here. A more detailed explanation is given elsewhere [5]. We extended the configuration as shown in Fig. 4. We define the stance and swing legs immediately before impact as \u201cLeg 1\u201d and \u201cLeg 2\u201d and derive their dynamic models independently. We define qi = [ xi zi \u03b8i ]T as the extended coordinate vector for Leg i and define q = [ qT 1 qT 2 ]T as that of the whole system. The inelastic collision model is then derived as M\u0304(\u03b1)q\u0307+ = M\u0304 (\u03b1)q\u0307\u2212 \u2212 JI(\u03b1)T\u03bbI , (4) where M\u0304 \u2208 R 6\u00d76 is the inertia matrix corresponding to q and \u03b1 [rad] is the half inter-leg angle at impact and is defined as \u03b1 := \u03b8\u22121 \u2212 \u03b8\u22122 2 = \u03b8+ 2 \u2212 \u03b8+ 1 2 > 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002473_s12206-009-0315-6-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002473_s12206-009-0315-6-Figure4-1.png", "caption": "Fig. 4. Animation of the ball bearing model.", "texts": [ " In our case we have ( ) ( ) ( ) ( ) 2 2 2 22 2 2 2 2 2 2 2 2 , , 4 , , , , A A A A B B f x y z R x y x y z R r f x y z x y z R = + \u2212 + + + \u2212 = + + \u2212 where rA is the toroidal pipe radius, RA is the radius of the circle being an axis of that toroidal pipe, RB is the ball radius. The geometry of the contacts inside the ball bearing can be seen in Fig. 3. The outer and the inner rings are assumed to be connected rigidly with outer and inner shafts in our case, attached one with another by the bearing. In the example under consideration, the body connected with the outer ring rests w. r. t. AF while the body connected to the inner ring rotates uniformly about the z-axis of AF, both thus performing the prescribed motion (see the animation image in Fig. 4). To verify the quality of the Hertz and volumetric models\u2019 implementation, we compared the vectors \u03b3 and nA as functions of time. The computational experiments showed that their coordinates coincide with a very high accuracy. To compare two contact models under analysis, one can present for example the time dependences for the normal component of the contact force. The Hertz and the volumetric cases corresponding to the same contact object inside the ball bearing model showed a high degree of a coincidence" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003500_9780857094537.9.623-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003500_9780857094537.9.623-Figure3-1.png", "caption": "Fig. 3: Three-Disk rotor scheme [15]", "texts": [ " The objective is to compare the influence of four following types of contact models on the rotordynamics [12], [13] and [14]: a- Damped penalty, b- Smooth penalty, exponential law, c- Smooth penalty, power law, dLagrange Multiplier. Particular attention is paid to two different behaviours generated by the contact: partial and full annular rubs in forward precession. Spectrograms of each simulation are then shown. The academic three-disk rotor in bending, presented in [15], is made of a steel shaft with a circular constant section and a 1.3m length. It is supported by two linear bearings, see Fig. 3, which their damping and stiffness parameters do not depend on the speed of rotation. Disks and bearing properties are described respectively in Tables 2 and 3. The operating rate is 6000 rpm (100 Hz). The stator is assumed to be rigid with no motion. Tab. 2: Disks properties Disk Mass [kg] Polar inertia [kg.m\u00b2] Diametral inertia [kg.m\u00b2] D1 14.58 0.123 0.0646 D2 45.95 0.976 0.498 D3 55.13 1.176 0.602 Tab. 3: Bearing properties kxx=kzz [N.m-1] cxx=czz [N.s.m-1] 7x107 1x103 The bending FE model is composed with 13 identical Euler-Bernoulli shaft elements and 56 DOF, (two lateral displacements and two slopes at each node)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003564_iccda.2010.5541299-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003564_iccda.2010.5541299-Figure4-1.png", "caption": "Figure 4. Static pressure profile at mating ring", "texts": [ " Because the size in Z-axis is so smaIl, the model was divided into two main zones (detailed in Fig 3) to get higher grid quality, and near the area of conjunction, finer grid is needed. In the calculation process, the X-axis direction nodes from 60 to 90, he Y-axis direction nodes from 50 to 60, and the Z-axis direction nodes from 30 to 60, ensure the V5-228 Volume 5 difference of opening force less than 1 %, meshing quantity is about 220000. Use SIMPLC discretization methods, the flow fields are described vividly, and the working principle also analyzed. III. NUMERICAL SIMULATIONS A. Pressure and path lines The numerical results in Fig 4 show the distribution of static pressure on whole mating ring face. At the intersection of groove and land, there's clear change of gas film pressure, at the gas flow out, pressure will increased, hydrodynamic pressure is formed, at the place gas flow into, pressure will decreased, negative pressure (relatively) zone formed, gas film stiffuess and opening force will be weakened. At the area without grooves, pressure change gradually along radial direction, and remain slightly higher than the pressure at inner radius" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003612_s11044-010-9190-2-Figure9-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003612_s11044-010-9190-2-Figure9-1.png", "caption": "Fig. 9 Absolute orbits, described in a complex coordinate system", "texts": [ "3 Absolute orbits A complex coordinate system (z; j \u00b7y) is introduced to describe the orbits of the shaft centre point W , the centre of the stator mass S, the centre of the shaft journal V, and the centre of the bearing housing B. It is positioned at the centre of the stator mass coordinate system (zs;ys), which is also the centre of the rotor mass coordinate system (zw;yw). But also the orbits of the shaft journal V and the bearing housing B can be described in this complex coordinate system (Fig. 9). The orientation of the complex coordinate system is linked to the direction of rotor rotation \u03a9 , so that the direction of rotor rotation is in the positive mathematical direction. For the mathematical description of the orbits in the complex coordinate (z; j \u00b7y) system, a new index \u03b4 is introduced: \u03b4 = w, s, v, b (20) The complex vectors r\u03b4,\u03ba (\u03b4 = w, s, v, b), which describe the orbit of shaft centre point W with \u03b4 = w, the orbit of stator S with \u03b4 = s, the orbit of journal shaft V with \u03b4 = v, and the orbit of bearing housing B with \u03b4 = b, can be described in the complex coordinate (z; j \u00b7 y) for each single excitation: r\u03b4,\u03ba = z\u03b4,\u03ba + y\u03b4,\u03ba \u00b7 j (21) With the components of the amplitude vectors q\u0302\u03ba,cos and q\u0302\u03ba,sin the complex vectors r\u03b4,\u03ba can be described by r\u03b4,\u03ba = z\u0302\u03b4,\u03ba,cos cos(\u03a9 \u00b7 t + \u03d5\u03ba) + z\u0302\u03b4,\u03ba,sin sin(\u03a9 \u00b7 t + \u03d5\u03ba) + j \u00b7 [y\u0302\u03b4,\u03ba,cos cos(\u03a9 \u00b7 t + \u03d5\u03ba) + y\u0302\u03b4,\u03ba,sin sin(\u03a9 \u00b7 t + \u03d5\u03ba) ] (22) With the EULER equations cosx = 1/2 \u00b7 (ejx + e\u2212jx ); sinx = \u2212j/2 \u00b7 (ejx \u2212 e\u2212jx ) (23) the complex vector r\u03b4,\u03ba becomes r\u03b4,\u03ba = 1 2 \u00b7 [z\u0302\u03b4,\u03ba,cos + y\u0302\u03b4,\u03ba,sin + j \u00b7 (y\u0302\u03b4,\u03ba,cos \u2212 z\u0302\u03b4,\u03ba,sin) ] \u00b7 ej (\u03a9\u00b7t+\u03d5\u03ba ) + 1 2 \u00b7 [z\u0302\u03b4,\u03ba,cos \u2212 y\u0302\u03b4,\u03ba,sin + j \u00b7 (y\u0302\u03b4,\u03ba,cos + z\u0302\u03b4,\u03ba,sin) ] \u00b7 e\u2212j (\u03a9\u00b7t+\u03d5\u03ba ) (24) So referring to [1], the complex vector r\u03b4,\u03ba can be described by a superposition of two rotating complex pointers" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001230_detc2005-84223-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001230_detc2005-84223-Figure2-1.png", "caption": "Figure 2: Two cases for distance calculation between a cylisphere and a plane", "texts": [ " QUADRILATERAL PLANES Quadrilateral planes are represented by three points, 1P , 2P , and 3P , and a half thickness (to give the plane some volume), t . A fourth corner point, 4P , is calculated to close the quadrilateral such that 4 2 3 1P P P P= + \u2212 . (24) Because quadrilateral planes are only used to model obstacles and all robots are modeled using cylispheres, a witness point, oP , on a quadrilateral plane fits into two possible categories. It is either a projection of one of the two endpoints of the associated cylisphere onto the plane (see Figure 2a), or it is calculated just as for a cylisphere using one of the edges of the quadrilateral to represent the obstacle cylisphere (see Figure 2b). If oP is a cylisphere endpoint projection, then 4 The third and fourth terms in Equation (23) could be combined, but writing the equation as shown will prove useful later in the derivations. ownloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?u ( )( ) ( )( )1 1 1 \u02c6 \u02c6 \u02c6 \u02c6 o c cP P P r r P P s s P= \u2212 + \u2212 +i i , (25) where cP is the nearest cylisphere endpoint and r\u0302 and s\u0302 are orthogonal unit vectors in the plane. Assuming the plane is fixed, differentiating Eq. (25) results in ( ) ( )\u02c6 \u02c6 \u02c6 \u02c6 o c cP P r r P s s= +i i " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002715_02533839.2008.9671426-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002715_02533839.2008.9671426-Figure1-1.png", "caption": "Fig. 1 Physical configuration of the cross section at the mid-plane z* = 0 of a short journal bearing", "texts": [ " The continuity equation and momentum equation of an incompressible couple stress fluid neglecting body forces and body couples, derived by Stokes (1966), are \u2207 . V* = 0, (1) \u03c1DV* Dt = \u2013 \u2207 p* + \u00b5\u2207 2V* \u2013 \u03b7c\u2207 4V* , (2) where V* is the velocity vector, \u03c1 is the density, p* is the pressure and \u00b5 is the classical viscosity coefficient. The constant \u03b7c denotes a new material constant responsible for the couple stress fluid and has the dimension of momentum. In this study we consider the physical configuration of the cross section at the midplane z* = 0 of a short journal bearing lubricated with an incompressible Stokes couple stress fluid, presented in Fig. 1. The journal rotor of radius R is rotating with angular velocity \u03c9* within the bearing shell. To obtain the non-Newtonian couple-stress dynamic Reynolds-type equation the process is described as follows: (a) solving for the velocity component in the x*-direction from momentum Eq. (2), (b) integrating the continuity Eq. (1) across the fluid film, (c) obtaining the Reynolds-type equation. From the work of Lin (2001) the non-dimensional non-Newtonian dynamic Reynolds-type equation for a short journal bearing is \u2202 \u2202z [g(h, l) \u2202p \u2202z ] = 24\u03b22[(1 \u2013 2 d\u03d5 d\u03c4 )\u2202h \u2202\u03b8 + 2\u2202h \u2202\u03c4 ] " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001008_kem.291-292.163-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001008_kem.291-292.163-Figure1-1.png", "caption": "Fig. 1 illustrates the mechanical movements both of the crankshaft rotation and the horizontal wheel head travel with the wheel head oscillating type CNC crankshaft pin grinder. The wheel head moves horizontally along X axis according to the journal rotation around C axis.", "texts": [ " Therefore, in order to make the grinding mechanism clear, the substantial surface speed of both the workpiece and the grinding wheel are investigated. Then, the grinding results such as grinding force, surface roughness and residual stock are estimated by simulation method, and effectiveness for precision under the controlled speed ratio grinding method is compared to the one under the constant journal rotation speed grinding method. Analysis of the crankshaft pin grinding mechanism Fig. 2 shows a model of grinding mechanism of the pin. When the journal rotates, then the pin separates from the wheel surface. In order to Fig. 1 Crankshaft pin grinding with wheel head oscillating type CNC crankshaft pin grinder 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-13/04/15,08:26:42) maintain the grinding condition, the wheel head moves horizontally. However, the journal rotational center position moves in this figure, instead of the grinding wheel horizontal movement" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002818_nme.2282-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002818_nme.2282-Figure4-1.png", "caption": "Figure 4. Non-linear pendulum oscillation about point A and forces.", "texts": [ " The second input is directly related to term R \u00b7 y3 appearing in the equation, and then thirdly, the second-order derivative of the unknown function is expressed by multiplication of matrices (A(1) i j \u00b7A(1) i j ) \u00b7 y. A comparison among the CM, DQ solution to [6] and Equation (17) is given in Table II. Observing the table reveals that the CM method captures more accurate results as compared with the solutions of Reference [6]. For large angular displacements, the system becomes non-linear. As shown in Figure 4, the equation of motion of the pendulum can be established by applying Newton\u2019s principle of moment equilibrium about point A d2 (t) dt2 + g L \u00b7sin( (t))=0 (23) where 2=g/L is the natural frequency of the pendulum. Therefore, Equation (23) can be restated as follows: d2 (t) dt2 + 2 \u00b7sin( (t))=0 (24) with the initial conditions 0= 4 and 0=0 (25) The analytical solution to Equation (23) is required in order to compare the present work\u2019s results. For this reason, we use the invariance of the unknown function with time because it is assumed that there is no damping factor in the system" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002468_msec2009-84012-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002468_msec2009-84012-Figure4-1.png", "caption": "Fig. 4: A cell volume showing the geometric factors that affect laser energy delivery.", "texts": [ " The thermal models must therefore accept a number of expressions to define the laser intensity across a complex surface. There are several factors to consider in determining the laser intensity at a given location on the complex surface. These include the laser incident angle \u03b8, the local surface angle \u03c8, the surface area of the given cell volume as seen from the perspective of the laser l\u03b8, and the surface area of the cell as seen from the vertical axis, lrad. These factors are represented on a surface cell diagram shown in Fig. 4. To adjust for these factors, a multiplier function M was used in the thermal models, given by Equation 7 ( ) radl l xIM \u03b8= (7) where I is the normalized intensity of a given laser profile (i.e. Gaussian or top-hat) as a function of the distance x from the center of the beam, and l\u03b8 and lrad are defined in Fig. 4. Fig. 5 shows two multiplier functions corresponding to the surface geometry in Fig. 3. In this example, the beam has a top-hat profile. For a 0\u00b0 incident angle, one expression for the top-hat profile is sufficient, but a piecewise function must be used for an incident angle of 20\u00b0. 2.2 Phase Transformation Kinetic Model The kinetic model predicts the microstructural transformations in hypoeutectoid steels during laser hardening processes. It is solved simultaneously with the thermal model, and can therefore include the effect of latent heat of transformation in the energy equation" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001424_jmes_jour_1973_015_066_02-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001424_jmes_jour_1973_015_066_02-Figure2-1.png", "caption": "Fig. 2. Components of frictional force (wheel locked)", "texts": [ " A constant forward speed v will be assumed and the steering axis will be constrained to the vertical; therefore, gyroscopic couples can be ignored. Contact patch deformation is ignored in the first instance, i.e. \u2018hard\u2019 contact is assumed. Simple coulomb friction will be used, with frictional force proportional to normal reaction R, namely pR, with p the dynamical-friction coefficient. This force opposes directly the relative slip-velocity vector. 2 WHEEL LOCKED CASE Fig. 1 shows the basic geometry and Fig. 2 shows the resulting contact patch-force components, with V = dv2 + + 2vtd sin e. This gives the following equation of motion for a rotation l3 about the steering axis. t2&R vt sin epR V Id = --- V where Z is the moment of inertia about the steering axis. The MS. of this paper was received at the Institution on 6th November Simon Engineering Laboratories, University of Manchester, t References are given in the Appendix. I972 and accepted for publication on 19th June 1973. 33 Manchester MI3 9PL. For small 0 To interpret this, take the usual case in practice of v2 % (tt#)2, and Z8+(--)b+(pRt)0 pRt2 = 0 This is the standard form for viscously damped single degree of freedom vibration" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000089_rspa.2002.1105-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000089_rspa.2002.1105-Figure3-1.png", "caption": "Figure 3. The eye splice and core subrope.", "texts": [ " (a) The splice The splicing of these specific ropes is described in two parts: firstly, the core is considered, this being common for all three splices; and, secondly, the splicing of the outer subropes by the long (transmission), the Admiralty and the Liverpool splices is detailed. In order to describe the geometry of the splices, the following notation is used. The components from the rope are labelled R and those from the splice are S. (b) The core Since the rope being considered here is a six-round-one structure, the first point to be considered is the core subrope; for this work the core is assumed to be laid along itself as shown in figure 3 and, in more detail, in figure 4. Although there are migrating strands that entrap the incoming S subrope (from the splice), the main mechanism for keeping the two subropes together is friction. The important contribution to this is the contact force initiated by the entrapping strands from the rope R subrope and reinforced by the action of the outer layer on the core. A simplified but adequate theory for this friction splice is developed in the following section on the long splice. At the crotch of the splice, both subropes are subject to the same load; as the station is advanced to the end of the splice zone, the splice S subrope sheds its load to the rope R subrope, and at the end of the splice, it carries zero load" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001920_cpe.2007.4296534-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001920_cpe.2007.4296534-Figure1-1.png", "caption": "Fig. 1. Basic scheme of inverter controlled twin stator CDFIG", "texts": [ " Such twodoubly fed machine arrangement leads to a so-called cascaded doubly fed induction machine (CDFIM) [5, 9- 14] which can operate as cascaded doubly fed induction generator (CDFIG) [5 , 9, 11, 14]. Rotor terminals a], bl, cl and a2, b2,c2 of both machines are connected back to back in sequence [14]: a] - a2, b1 - c2, cl - b2 which ensure the generate torques act in the same direction. While the stator of the first machine is connected to the grid, the stator of the second machine can be fed from the power converter. We propose the PSIM simulation software [15] to steady state analysis of twin stator CDFIM. The basic power generation system with considered twin stator CDFIG is shown in Figure 1. 1-4244-1055-X/07/$25.00 (2007 IEEE. dir(1,2)y rOLm,nLs +3(OL2 RstL dtlw(]s(1,2)x + lr(1,2)x + Lis(,2)y +dt WC7 WG W7 RrLs. Ls Lm ir(1,2)y + Ur(1,2)y WcUs(1,2)y (5) where us [uSX uSY] and ur :urx Ury] denote stator and rotor voltage vectors, Rs, Rr - denote stator and rotor resistances, LS , Lr, Lm - denote stator, rotor and mutual inductances and wy=Ls r-Li. It assumed that all parameters are identical for both machines. For the electrically coupled rotors of induction machines IMI and IM2 from Figure 1 appears [13]: ir2 =rI '= ir II. DYNAMICAL MODEL OF TWIN STATOR CDFIG The twin stator CDFIM under consideration comprises two identical 3-pole pair wound rotor machines with their three-phase stator windings independently excited. It is assumed the two stator voltages are of different frequency. Rotor terminals cross-coupling ensures that two stator excited fields co-operate that the CDFIM is operating in a 'synchronous' mode producing the same frequencies for the rotor currents. The synchronous speed of CDFIM is a half of synchronous speed of individual machine" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003356_978-3-642-33509-9_4-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003356_978-3-642-33509-9_4-Figure2-1.png", "caption": "Fig. 2. Structure of permanent magnet biased axial magnetic bearing", "texts": [ " The two auxiliary bearings - are normal ball bearings and they support the rotor in radial and axial direction when the load goes beyond the capacity of the magnetic bearings or the active magnetic bearing losses its stability. The axial magnetic bearing is an active magnetic bearing and its bias magnetic field is established by permanent magnets to reduce energy consumption. The main load of the axial magnetic bearing is the axial components of the wind force imposed on the rotor though the impeller. The structure of the permanent magnet biased axial magnetic bearing is given in Fig. 2. The axial stator and rotor core are made with solid soft magnetic material. The axial control coils are made with enameled wires. The permanent magnet ring is made from NdFeB rare earth material and magnetized in radial direction. The body parameters of the permanent magnet biased axial magnetic bearing are given in Table 1. According to the working principle of the active magnetic bearing system, the structure of single-degree-of-freedom magnetic bearing closed-loop control system is shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000910_iecon.2005.1569137-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000910_iecon.2005.1569137-Figure8-1.png", "caption": "Fig. 8. Two phase symmetrical components method", "texts": [ " [IMA] and [VMA] mean the actual main/auxiliary winding currents and voltages. By converting into [Ima] and [Vma], formula can be treated as if main and auxiliary turns are the same. This conversion enables it easy to construct mathematical model of single-phase LSPM. In the next step, method of symmetrical components is applied to [Ima] and [Vma] as follows. \u2212 = 2 111 I I jjI I a m , [ ] [ ][ ]12ISI ma = (10) \u2212 = 2 111 V V jjV V a m , [ ] [ ][ ]12VSVma = (11) Fig.8 shows the mechanism of these equations. I1 , I2 means positive and negative components winding current. In this way, actual current and voltage are separated to positive and negative components. From (7), (9) and (8), (10) we get \u2212 = \u2212\u2212 2 1 11 I I jjI I A M \u03b2\u03b2 (12) \u2212 = 2 1 V V jjV V A M \u03b2\u03b2 (13) By solving the equivalent circuit of positive and negative components, [I12], [V12] are obtained. Once [I12], [V12] are given, actual current and voltage of [IMA] and [VMA] are obtained by substituting them into (3)-(6)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000453_13552540510589421-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000453_13552540510589421-Figure2-1.png", "caption": "Figure 2 EB-PVD chamber for RM at Penn State", "texts": [ " This extensive value-added processing contributes to high cost and a fairly limited range of commercial-shape components. Therefore, in this paper EB-PVD has been studied as an alternative to address these problems and is considered as an alternative process for RM technology of rhenium components. EB-PVD process is used for high-rate deposition where a high-energy and focused electron beam directly heats and evaporates a vapor-emitting surface of material ingots inside vacuum chamber as shown in Figure 2. Typical deposition rates of 10-300mm/min are obtained for metals with an evaporation rate of 10-15 kg/h, compared to deposition rates of about 1mm/min from high-rate magnetron sputtering. Figure 2 shows material ingots to be evaporated fed through crucibles at the floor of the chamber, which also act as collectibles for the molten material. The high-powered electron beams are focused on the material ingots and they are evaporated to a vapor plume in the chamber. Material has to be fed at a constant rate to compensate for this evaporated material. Workpiece to be produced is placed and moved in this region of vapor plume so that the evaporated material condenses on the workpiece. The resulting depositions are effective in providing protection against corrosion, oxidation, and wear, which in turn lead to greatly enhanced performance and extended life" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001607_robot.2005.1570128-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001607_robot.2005.1570128-Figure2-1.png", "caption": "Fig. 2. Model of gantry crane", "texts": [ " Using the inverse dynamics calculation and the obstacle detection method, a safe anti-sway control method is presented. Finally, experiments using the method are shown. 0-7803-8914-X/05/$20.00 \u00a92005 IEEE. 253 II. INVERSE DYNAMICS OF GANTRY CRANE This section presents inverse dynamics calculation considering the crane mechanism as a manipulator. In the current study, the inverse dynamics means calculation of rope length, trolley position and their actuator\u2019s torques to realize given trajectory of suspended load position. We here consider a gantry crane system shown in Fig.2. In the system, we assume point mass load, no bending moment and no elasticity for the rope. By expressing the position of the load with p = [px py pz]T and tension force of the rope by f = [fx fy fz]T, the motion of equation for the suspended load system is described as m(p\u0308 \u2212 g) = f (1) where m is mass and g = [0 0 \u2212 g]T is acceleration vector of the suspended load. The rope drawing point q is manipulated on x \u2212 y plane by controlling the positions of girder and trolley. Then, the point q is calculated using the geometrical relation of load position p and rope force vector f as q = px \u2212 pz p\u0308x p\u0308z + g py \u2212 pz p\u0308y p\u0308z + g 0 (2) Where the acceleration of p must be satisfied the following condition, p\u0308z > \u2212g (3) because the suspended load does not produce the acceleration over the gravitational acceleration" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000027_s00216-002-1714-z-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000027_s00216-002-1714-z-Figure2-1.png", "caption": "Fig. 2 The absorption spectra of the reagent solution at different acidity: 1. 1.0 mol L\u20131 HCI, 2. 1.0 mol L\u20131 H3PO4, 3. 0.1 mol L\u20131 acetic acid, 4. borax buffer solution (pH 9.0), 5. 0.1 mol L\u20131 sodium hydroxide. Other conditions: HNAR 7\u00d710\u20135 mol L\u20131, mixed surfactant 2\u00d710\u20133 mol L\u20131 CTMAB and 5\u00d710\u20133 mol L\u20131 Triton X-100", "texts": [ "0 mL of 1.34\u00d710\u20133 mol L\u20131 HNAR solution. Dilute to volume with water and stir well. Leave for 20 min and measure the absorbance at 480 nm against a blank. Determine the gold concentration using calibration graphs prepared in the same manner. The characteristics of HNAR HNAR is a salmon pink powder, melting point 215\u2013216 \u00b0C. It is difficult to dissolve in water and dissolves easily in organic solvent such as alcohol and acetone. The effect of acidity on the absorption spectra of the reagent is shown in Fig. 2; it can be seen that the reagent has different colors at different acidity due to different ionization. In general, the reagent solution is yellow in acid or red in strongly basic media. Assuming the reagent to be H2L, four species are involved in the acid dissociation behavior, the equilibrium between these species can be written as: K1 K2 K3 H3L+ \u2192\u2190 H2L \u2192\u2190 HL\u2212 \u2192\u2190 L2\u2212 The respective ionization constants were determined in water, maintaining the ionic strength at 0.2 with potassium nitrate, by a spectrophotometric method [37], the results were pk1=2" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000732_0094-114x(78)90039-3-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000732_0094-114x(78)90039-3-Figure2-1.png", "caption": "Figure 2. Different generations of Watt's sextic.", "texts": [ " A decisive progress has been achieved in a recent paper of Folkeson [5], but as the author uses a computer program to calculate the coordinates of a sufficient number of points of the coupler curve, his paper contains no formulas which could serve for the synthesis of optimized Watt mechanisms. tProfessor of Geometry, Technical University of Vienna, Austria. 155 To approximate, by a Watt sextic, a segment of the x-axis of a cartesian frame (O; x, y), let us locate the fixed pivots at L ( - e , f ) and M(e , - f ) . Then we replace the 4-bar linkage L A B M by a 5-bar linkage L I C J M , whose arms L I = M J = b are parallel to the coupler A B and whose links IC = JC = a are parallel to L A and M B , respectively (Fig. 2). To provide a constrained motion of the 5-bar linkage, the arms L I and M J have to be driven simultaneously in such a way that they conserve their parallel position; this might be achieved by an auxiliary parallelogram. As the triangle I JC is isosceles, the radii vectores O C = r and O I = R are orthogonal and connected by r2+ R 2= a 2. This induces a remarkable correspondence I--* C which maps the circular path i of I onto the path c of C. This planar transformation, defined by O I . L O C and I C = a (Fig. 2), allows an elegant graphical construction of the coupler curve c. Denoting the coordinates of C with x and y and those of I with X and Y, the equations of the involutary algebraic two-to-two transformation I ~ C (of degree 4) read 2 2 ~ a - R y2. x = ~ t Y , y = - X X with ,~ = ~ , R2=X2+ (1.1) Applying the inverse formulas X = -/zy, Y = ~x with /z 2 a 2 _/,2 = ~ , r2= x2+ y 2 (1.2) for the mapping of the circle i, described by (X + e) 2 + ( Y - f)2 = b 2, (1.3) we obtain the equation of the coupler curve c (X2+ y2)(X2 + y2_ d2)2 +4( fx + ey)2(x2 + y2_ a 2) = 0 with d 2 = a 2 - b2+ e 2 + f 2. (1.4) This equation shows that c is a sextic, i.e. an algebraic curve of order 6. The quantity d may be interpreted as the central distance of two isolated double points of c, situated on the base line L M (fx + ey = 0) which is an axis of symmetry for c. Representing the circle i (1.3) in parametric form by X = b c o s u - e , Y = b s i n u + f , (1.5) and introducing the auxiliary angles v and w of Fig. 2 by tan v = X / Y , cos w = R / a , (1.6) we obtain by application of (1.1) a polar representation of c x = r c o s v , y = - r s i n v with r = a s i n w . (1.7) This simple procedure to calculate points of the coupler sextic is similar to that used by Folkeson [5], but perhaps somewhat more convenient. 2. Approximating Quintic For small values of y we may neglect higher powers of y in eqn (1.4) of Watt 's coupler sextic c. Thus we get a quintic x ( x 2 - d2) 2 + 4 f ( f x + 2ey)(x 2 - a s) = 0. (2.1) It approximates the flat branch of c. The coincidence of the curves c and E is good, as they have in common five precision points Co, C~ . . . . . C4 on y = 0 (Fig. 2). t t =2, t : 5 y I p = 3 , q = l O f a = l O - . - t = 2 p ..~ 1 = 2 t : 4 p ' i t~ Rgure 3. O p t i m i z e d a p p r o x i m a t i n g q u i n t i c . The rational quintic ~, symmetric with respect to the origin O, is now better apt for applying Chebyshev's optimization principle. The fundamental condition that 6 has to remain between two parallel lines y = -+h within an interval - t _-__ x _-__ t, touching them at points with abscissas -+h and ---t2 (Fig. 3), requires that eqn (2.1) for y = h is equivalent to (x - t ) ( x - t02(x + t2) 2 = 0 (2", " To guarantee the full length I of the approximation also for the sextic c it would have been better to choose p somewhat greater than 1/4, say p = 26. In any case the obtained approximation is essentially better than that of the not optimized linkage in[6]; corresponding to the choice of the pivot coordinates e = 29.692 and f = 145.710 in [6], the deviation amounts there to Ymax = 0.077. By the way, Watt 's symmetric 4-bar linkage always might be replaced by another doublerocker linkage, derivable by means of Roberts ' theorem and sometimes called the grasshopper mechanism of Evans [3]. In Fig. 2 this is achieved directly if the midpoint H of the bar I C is connected to the origin O by a single bar of length a/2. Then the replacing mechanism consists in the 4-bar linkage OHIL which also leads the coupler point C along Watt 's curve c. Taking the opportunity to compare the qualities of Burmester 's and Chebyshev's optimization principles, let us finish by considering the classical mechanism of Fig. l a = f , b = e . (4.6) The eqn (1.4) of the generated coupler curve c- -which now has a higher inflection point at O-- takes the form (X 2 + y2) (X2 + y2 _ 2 a 2 ) 2 + 4(ax + b y ) 2 ( x 2 + y2 _ a 2) = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000056_robot.1992.220297-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000056_robot.1992.220297-Figure1-1.png", "caption": "Fig. 1 Two robot arms holding an object.", "texts": [ " The force ellipsoid [13] represents the characteristics of the end effector forces that correspond to all unit norm joint torques, and since the joint torque vectors of unit norm are equivalent to joint torque vectors of unit effort , the concept of force ellipsoid can be utilized in the optimal load distribution problem when the performance index is minimum effort. This paper is organized as follows. In section II, the optimal load distribution problem is formulated, and in section 111, the solution procedure is presented. In section IV, a numerical example is used to illustrate the proposed solution, and the internal force effect is 46 I 0-8186-mo-4/92 $3.00 01992 IBEE studied in section V. Finally, conclusions are drawn in section VI. 11. Problem Formulation The model of the two robot arms and the object is shown in Fig. 1 , where the end effectors of the two robots are grasping an object. The object is held rigidly so that no relative motion is allowed between the object and the end effectors. Two robots are working in the undistinguished mode, so that no distinction is made regarding the master or slave status. For the purpose of convenience, the superscripts i = 1, 2 are used to indicate the two robots. Let 'q = joint position vector of robot i, 'q e R\", 'J = manipulator Jacobian matrix,.'J E R m x n , m s n, 'T = joint torque vector of robot i, 'T E R\", 'F = Cartesian force" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000714_ed081p1628-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000714_ed081p1628-Figure1-1.png", "caption": "Figure 1. The heart of the experiment (A) is a stationary optical mouse that is vertically aligned and a reference surface that moves according to the extension of the polyethylene film. (B) A schematic of the setup.", "texts": [ " 11 November 2004 \u2022 Journal of Chemical Education 1629 Materials \u2022 Two C-clamps \u2022 A loading fixture comprising (a) an upper portion\u2014a clamp for specimen attachment, and (b) a lower portion\u2014a clamp for specimen attachment, a cord to hang a weight, and a reference surface \u2022 A vertically mounted optical mouse: the mouse used here has a universal serial bus (USB) connection and possesses a resolution of 0.0635 mm \u2022 A screwdriver \u2022 Low-density polyethylene films of 20-mm width and 0.1-mm thickness from three different manufacturers; two horizontal lines are marked 100-mm apart \u2022 A 1.5-kg mass \u2022 A stopwatch It is important to ensure that the polyethylene film is properly fixed to the loading fixture (Figure 1). To do this the screws of the clamps at the upper and lower portions of the loading fixture are first loosened. The film is slipped inbetween the clamps and adjusted so that the two marked lines correspond to the edges of the clamps. The heads of the screws on the upper and lower portion of the loading fixture should all be on one side. After careful adjustment, the screws are firmly tightened. This is an important procedure as poor tightening can result in erroneous measurements. The vertically mounted optical mouse is then secured on any table using a C-clamp" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001872_978-1-4020-6366-4_15-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001872_978-1-4020-6366-4_15-Figure6-1.png", "caption": "Fig. 6a. The Watt four-bar linkage.", "texts": [ " With this design, the length of each oscillating link of the straight-line linkage was reduced to one-fourth instead of one-half the beam length. The entire mechanism could then be constructed so that it would not extend beyond the end of the working beam. This arrangement soon came to be known as Watt\u2019s parallel motion linkage, denoted as O2ABO4 in Figure 5. Through insight we can detect in this straight-line linkage the birth of a very ordered and advanced synthetic process. The kinematic analysis of the Watt four-bar linkage, see Figure 6a, and the geometry of the path of point M fixed in the coupler link AB (link 3) can be investigated using the method of kinematic coefficients [17]. The vectors that are required for the kinematic analysis of the Watt fourbar linkage are shown in Figure 6b. Modern Interpretation of Main Contribution to Mechanism Design The vector loop equation for the four-bar linkage can be written as \u221a I R2 + \u221a ? R3 \u2212 \u221a ? R4 + \u221a\u221a R1= 0, (1) where the first symbol above each vector indicates its magnitude and the second symbol indicates its direction. The known quantities are denoted by\u221a the unknown variables are denoted by ?, and the independent variable is denoted by I. Without loss in generality, the independent variable is assumed to be the angular position of link 2 and the unknown variables are the angular positions of the coupler link 3 and the side link 4" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001682_inmic.2004.1492982-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001682_inmic.2004.1492982-Figure2-1.png", "caption": "Figure 2: Helicopter process with h = 0", "texts": [], "surrounding_texts": [ "1. Introduction and Problem Formulation\nSome nonlinear control problems can be tackled nicely using feedback linearization. There are a few well known feedbnck linearization methods. 1) Exact linearization via sfatic state feedback 2) Approximate linearization via static state feedback 3 ) Exact linearization via dyiiamic state feedback.\nA two step exact linearization of the twin rotor modcl is described in [I]. The idea is to divide the dynamics of the system into two subsystems, exact state feedback linearization of subsystem 1 is done and desired state for subsystem 2 is obtained. Then, a servo controller is designed to track the desired state. The important factor over here is delay caused by servo controller and some nonlinearity left in the system. We will consider two scenarios. In first scenario, we will use an approach similar to described in [I], In second case, subsystem 2 will be considered in steady state and dynamics of subsystem 1 will be modified accordingly. This will allow exact state feedback linearization of whole system.\nIn this paper, a simplified model o f the twin rotor system will be considered. Firstly, we will design the feedback linearizing law and then, based on resulting linear system, state feedback controller will be designed. Two scenarios are considered 1) Partial linearization of exact model 2) Exact linearization of simplified model. Simulations results are shown for both cases.\n2. Plant Dynamics and Description\nThe dynamics for the twin rotor system, considered here, are derived in [Z] for the ETH helicopter process using Euler-Lagrange approach. A similar model is also derived in [3]; and control using LQG approach is designed. A schematic of the helicopter process configuration is shown in figure I .\nThe helicopter consists of a vertical axle (A), on which a lever arm (L) is connected by a cylindrical joint. The helicopter has two degrees of fieedom: the rotation of the vertical axle (angle ) with respect to the fixed ground, and the pivoting of the lever arm (angle ) with respect to the vertical axle. Two rotors\n0-7803-8680-9/04/$20.00 02004 IEEE. INMIC 2004", "are mounted on the lever arm: R, and R,, with the resultant aerodynamic forces giving rise to moments in the B and q directions respectively. The voltages U, and u2 to the rotor motors are the inputs to the system.\nThe dynamics for ETH helicopter model are: d . - lb=d dt\nU\" - -B =e dt\nWhere\n(4)\nL, = cos28J , -2hcosBsinBml,+ h2sin2Q8m+J,\nL5 = J , h2m\nThe values for different parameters are given in [ 2 ] . Due to complexity of nonlinear terms, exact state feedback linearization of (1) to (6) is not possible. Therefore, the model i s simplified by reducing the\nAfter inserting values of various parameters, the resulting dynamics of twin rotor system are:\ni2 = 1.16 x I O - ~ X : sec(x3) + 1 . I X ~ O - ~ X ; sec(x3)\n+2x2x, tan(+) ( 8 )\nWe can divide the dynamics in two subsystems. Subsystem I contains equations (7) to (10) whereas subsystem 2 consists of equations ( 1 1 ) and (12). Subsystem 1 represents the position of twin rotor system whereas subsystem 2 represents the velocity of main and tail rotor.\n3. Analysis and Feedback Linearization\nLooking at the dynamics reveals, subsystem 1 is the only nonlinear part of the system. A linearizing feedback law is derived as described in [4].", "Where\nf(x) =\nr o 0 1\n1 1.998 0.0705 1\n3.1 Vector Relative Degree\nA system is said to have vector relative degree\nr,, , T ~ at a point x\"if\n1) L,.L>~~(X) = o for all 1 5 j < m , for ail\n1 < i 5 m for all k 5 5 - I and for all x in the\nneighborhood of ?CO.\n2) The Xm matrix\nis nonsingular at x = x o . In this regard, subsystem I has veclor reIative\ndegree {2,2]at equilibrium point [xI 0 0.5836 01 for arbitrary x I . Here \"L \" represents the Lie derivative.\n3.2 Feedback Linearizing Law\nThe feedback lineanzing law is, then, given as\nv = A(x)+ b(x)u.\nU = A-' (x)[-b(xj+ V] (14) Where A(x) is as defined in (1 3 ) with m = 2 and\nThe resulting law has r 1 0.11 6 sec x3 1.1 sec x,\n1.998 0.0705\n2x,x4 tan x3\ni A(x)=\nb(x)= --x2 2 cosx3 sinx,\nThis feedback results following linear system\n0 1 0 0 0 0\nA=[' 0 0 0 1 l , B = l 1 0 0\n0 0 0 0 0 1\n.=[I 0 0 1 0 O o ] , D = 0 , 2\nA state feedback controller is designed using LQR approach for system in (18).\nThis would work fine if states x5 and x6 were available for feedback, but this is not the case here. So, one of the suggestions is to treat subsystem 1 under steady state. Is it permitted? Of course \"yes\", because, for positive U, and U?, x5aand x6 will always be positive. Equations (1 1) and (12) can be solved independently for given input U, and u2 respectively. Their solution for constant u1 and u2 is given below:" ] }, { "image_filename": "designv11_61_0000332_acc.1994.752362-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000332_acc.1994.752362-Figure1-1.png", "caption": "Figure 1: The Acrobot.", "texts": [ "6 The proposed control (36) is identical to that proposed in pseudolinearization [2], [6], with the exception that Tn(a) is more complex in this case. The additional term in Tn(=) (34) can be differentiated in Mathematica, although spline functions approximating F\u2019(s) and a : ( s ) are required. These are both computed using methods outlined in Remark 2.2. 3 The Acrobot Example In this section we derive the approximate state-feedback linearizing control law for the remotely driven Acrobot illustrated in Figure 1, and then present simulation results of its implementation. The Acrobot is a coplaner two-link robotic manipulator. Only one torque input is available: between the links. The mechanism is discussed at length in [5], [6] [8] and [9], and we omit the dynamic equations because of space limitations. Using Mathematica we have automated the process of 1. symbolically computing the pair (F(s ) ,G(s)) using (5); 2. symbolically computing the matrix O(s), defined in (8), along with 3. evaluating the expressions for al(s), a 2 ( s ) , a s ( s ) , and a q ( s ) according to (35) at the knot points, and interpolating a cubic spline through the resultant data; 4" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002883_icems.2009.5382964-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002883_icems.2009.5382964-Figure1-1.png", "caption": "Fig. 1. A stator of a two pole machine with a fractional conductor winding with q=3 and q2=1.", "texts": [ " This can be achieved, among other possibilities, by tuning the number of slots per pole per phase, q, and allowing it to be a non-integer. These windings are called in the literature fractional slot windings [1]. The number of conductors per slot is conventionally constrained to be an integer number. If this constraint is relaxed by allowing it to take a fractional value, more possibilities are introduced regarding the tuning of the machine parameters. Fractional conductor windings were defined and described in [2]. An example of a fractional conductor wound stator is given in figure 1 where q = 3 and one (= q2) of the coils in each phase belt has N2 turns as compared to the other two (= q1) coils in each phase belt having N1 turns. As the winding is not conventional anymore, the analytical expressions of all the classical parameters used in the equivalent circuits such as magnetizing inductance, leakage inductance, stator and rotor resistance have to be adapted. In [2], a new winding factor has been introduced together with the investigation of the air gap m.m.f (magnetizing inductance)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003013_robot.2009.5152257-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003013_robot.2009.5152257-Figure8-1.png", "caption": "Fig. 8. Object with one branch", "texts": [ " In [13], we assumed that the rib line at point P(ub) coincides with line BF referred to as a connecting line and defined separately variables of the left part ABFH and those of the right part BCDF to simplify integration of the potential energy. However, it was found that such assumption is not valid by comparison between the computational result and the experimental result. Applying our proposed model in this paper, the rib angle at the central part BEFG including point P(ub) is determined by minimizing the total potential energy of the object. In the case of a branched belt object as shown in Fig.8, by assuming that the central part BEH deforms so that it forms a part of a cylindrical surface and considering the continuity of the surface on line BE, EH, and BH, we can derive its deformed shape. Thus, our proposed model can represent deformation of various belt objects. In this section, the computation result will be experimentally verified by measuring the deformed shape of an angled belt object. Computation algorithm to derive the deformed shape of a straight belt object was mentioned in Section IIC" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003372_978-3-642-25486-4_28-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003372_978-3-642-25486-4_28-Figure2-1.png", "caption": "Fig. 2. Structure of kinematic analysis of parallel indexing cam mechanism", "texts": [ " Numerical integration is applied to obtain the responses. Through the numerical results we analyze effects of clearance on dynamics response and the reasons for impact vibration on turret. It provides basis for obtaining the fault characteristic of indexing cam and implementing dynamic model-based fault diagnosis. The basic structure of parallel indexing cam mechanism is shown in Fig. 1. Profile of parallel indexing cam can be generated by the method of Exponential Product Formula. The general frame is built in Fig. 2, in which origin of frame is located in center of turret, axis z is the rotation direction of turret and axis x is defined as line from center of turret to cam center. In order to describe the motion of parallel indexing cam, skew is defined by rotation vector and coordinate of any point in rotation axis. The skew parameters are shown as table 1. where a is the center distance between cam axis and turret axis; r is the radius of roller in turret. Suppose Af (0) is a point in the center line of a roller, and its coordinate is T f rA )00()0( 1= According to exponential product formula and screw theory initial cam theoretical profile ),( 21 \u03b8\u03b8cA can be obtained by following equation ),()0( 21 \u02c6)(\u02c6 221101 \u03b8\u03b8\u03b8\u03be\u03b8\u03b8\u03be cf AeAe =+ (1) Where 1\u03b8 is rotation angle of turret, 2\u03b8 is rotation angular of cam\uff0c 1\u03be , 2\u03be are skew value of turret and cam shaft" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000297_s0076-6879(04)81045-0-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000297_s0076-6879(04)81045-0-Figure3-1.png", "caption": "Fig. 3. Diagram of position of electrodes in the sensor body; the end of the Ag/AgCl reference electrode is slightly offset from the end of the Pt/Ir working electrode, as shown in the inset.", "texts": [ " Care should be taken to carefully cut the Pt/Ir wire with a clean, square cut so that the surface area of the working electrode is as reproducible as possible from sensor to sensor. 4. The reference electrode and working electrodes are then threaded through the sensor body, which is a 75-mm length of Silastic medical grade tubing (0.91 mm i.d. 1.2 mm o.d.). The electrode ends are aligned and pushed through the tubing until about 3 cm of wire extends past the end of the tube. After the electrodes are in place, retract the Ag/AgCl electrode slightly so that it is offset and shorter than the Pt/Ir working electrode (see Fig. 3). This avoids the potential problem of the electrodes directly touching one another. The length of the wires employed to make the electrodes and the sensor body can be adjusted depending on the size of the catheter used to implant the sensor. The dimensions suggested here are appropriate for use with 1.16-in.-long catheters. 5. Using a 30-gauge needle and a disposable syringe, fill the sensor sleeve with 0.15 M KCl/1.5%(wt) Methocel solution. Gentle heating is required to fully dissolve the Methocel to prepare the solution, but avoid overheating and boiling the solution" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000983_icit.2003.1290263-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000983_icit.2003.1290263-Figure7-1.png", "caption": "Fig. 7: The Robot Model of HVC", "texts": [ "1 Hopping Velocity Controller(HVC) Velocity command for the desired hopping height is generated by HVC. And, shortage energy of or excess energy of pre-hopping is adjusted. Velocity command value \"u&Ti is derived from \"the law of conservation of kinetic energy\". (8) w uGzh cmd =J2g (w&md - ' \"hm) cmd = cmd w cmd w cmd vGh [ ' G z h 'Gyh ' G z h 1' \" h S d : The desired height of COG PhCh : The takeoff height of COG Moreover, velocity command \"v&!: is generated so that the acceleration of COG at takeoff may become greater than gravitation acceleration. the robot model of HVC is shown in Fig. 7. The used characters in Fig. 7 are as follows. 1, : Leg length 0, : Angle of ground and the line which connects COG and the tip of leg ZG = &sin R, (11) (12) w w . zG = 1, sin 0, - 1, cos 0,0, Leg length lg is changed in order to acquire '\"u&?f. Eq. (13) expresses the relation between 1, and \"'vz:. Furthermore, energy Eer\" is the difference of desired hopping height \" h T d and a last actual one \" ' h y . (14) =w hcmd -w F e G hG E\"' < 0. . . Excess energy equivalent to E\"' Ee\" 2 0 . . . Shortage energy equivalent to E\"' EeC' is changed into the vertical velocity in Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003422_asjc.598-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003422_asjc.598-Figure1-1.png", "caption": "Fig. 1. Flexible joint electric drive system.", "texts": [ " According to Theorem 1, if there exist matrices Pa > 0, La, K and W satisfying ( ) ( ) ( ) ( ) , ( ) A L C P P A L C H KC H KC P B UC a a a T a a a a a a a T a a a a a + + + \u2212 + + < = \u03bc 0 T, \u23a7 \u23a8 \u23aa \u23a9\u23aa (35a) (35b) where U has the same form as in (15), B B E B E a a a2 2 0 0 [ ] = \u23a1 \u23a3\u23a2 \u23a4 \u23a6\u23a5 , C C H C H Ir a a a \u23a1 \u23a3\u23a2 \u23a4 \u23a6\u23a5 = \u23a1 \u23a3\u23a2 \u23a4 \u23a6\u23a5 0 , then the signal ucps(t) given in (16) can be used to stabilize the error dynamics. Further, the parameters La, K and W also can be designed using Algorithm 1, since the conditions (35a) and (35b) have the same form as (13) and (14). What is more, since UC B UCB a a = , Lemma 3 can be used directly to verify whether (35b) can be satisfied or not. VI. SIMULATION EXAMPLES Example 1. The fault reconstruction scheme is illustrated on a flexible joint electric drive system depicted in Fig. 1. The load and the motor are connected by a torsional spring. The dynamic model of the system is given by: \u00a9 2012 John Wiley and Sons Asia Pte Ltd and Chinese Automatic Control Society \u03b8 \u03c9 \u03c9 \u03ba \u03b8 \u03b8 \u03c9 \u03c9 \u03b8 \u03c9 \u03c9 \u03ba \u03b8 \u03b8 l l l l m l v l l l m m m m l m = = \u2212 \u2212 \u2212 = = \u2212 \u2212 , ( ) ( ), , ( ) J K J J \u03a82 K J K J K J u fv m m v m m m \u03c9 \u03c9 \u03c4\u2212 + + \u23a7 \u23a8 \u23aa \u23aa \u23aa \u23a9 \u23aa \u23aa \u23aa \u03a81( ) ( ), (36) where ql(t) and wl(t) are the load position and velocity, respectively. qm(t) and wm(t) are the motor position and velocity, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003949_j.ast.2011.09.010-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003949_j.ast.2011.09.010-Figure6-1.png", "caption": "Fig. 6. Views of the refined mesh model near the FS482 frame crown.", "texts": [ " It was determined that the hub mass in the fuselage model has to be reduced from 2100 to 400 lb in order to prevent duplication (Rhoads [13]) This mainly affects the resonance frequencies (see Table 3 for examples). Mode shapes are very similar for the three aircraft weights. The fuselage model mesh has been refined at the crown of frame 482 thus providing enough detail to capture the actual loca- tion of strain gauges in this region. Stress results are derived from the whole fuselage model with refined mesh at FS 482. Fig. 6 illustrates the mesh refinements in the region of strain gauge 2. No separate component model is used in the present study. The dynamic fuselage properties used in the aeromechanics model are derived from the fuselage model including this refinement. Frequency results for the refined fuselage model at various aircraft weights are presented in Table 3. Note that frequencies have changed significantly by weight changes and the reduction of hub mass (compare with Table 2). The mode shapes remain very similar" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000920_iecon.2005.1569144-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000920_iecon.2005.1569144-Figure8-1.png", "caption": "Fig. 8 shows the fan motor under test and the drive PCB. The fan prototype is the improved one based on the", "texts": [], "surrounding_texts": [ "We have known that the resting rotor has two possible positions, as shown in Fig.2. If the controller supplies an appropriate current pulse to the winding for a sufficient period of time and then decreases the current to zero slowly, it is achieved that the rotor will stop at the fixed position determined by the current direction through the following analysis. When the controller supplies such a large positive current pulse to drive the rotor rotating, the stator teeth of A1 and A2 have south and north polarity separately in spite of the rotor initial position, as shown in Fig.3 (a) and Fig.4 (a). Now what analyzed is the motion of the rotor under the different initial conditions. (a) Motor state while current exciting winding at Position 1 (b) Stator MMF and rotor MMF at Position 1 Fig.3 Motor state while current exciting motor at Position 1 (a) Motor state while current exciting winding at Position 2 (b) Stator MMF and rotor MMF at Position 2 Fig.4 Motor state while current exciting motor at Position 2 (a) Motor state while current exciting motor at zero position (b) Motor state while no current exciting motor at zero position Fig.5 Motor state while rotor at zero position If the rotor initial position is Position 1 (shown in Fig.2 (a)), Fig.3 shows the state of the motor when a positive current pulse is supplied to the winding. From Fig.3, it is seen that the rotor pole S has the same polarity as the tooth A1. The magnetic motive forces (MMF) relationship between stator (Fs)and rotor (Fr) is shown in Fig.3 (b). Obviously, the current pulse results in rotor\u2019s counterclockwise rotation. The pulse keeps on, and the rotor will rotate an electrical angle of )( \u03b8\u03c0 \u2212 degree. Then the rotor will stop after vibrating for a moment at the special position (shown in Fig.5 (a)), where the polarities of the stator tooth and rotor pole are opposite and the rotor is attracted by the stator. Fig.6 illustrates the rotor position and speed during this procedure. In the figure, S2 is the rotor position curve whose zero position is shown in Fig.5 (a), W2 is the rotor speed curve, and i is the current curve. Fig.6 shows that the rotor stops after t2. Then if controller lets the current decrease to zero slowly at t3 (t3>t2), the rotor yields a counterclockwise rotation, as shown in Fig.5 (b), and eventually offsets an electrical angle of \u03b8 degree and stop at Position 2 shown in Fig.2 (b) under the interaction between cogging torque and electromagnetic torque. That is, after the positive current pulse acts, the rotor stops at Position 2 if the rotor initial position is Position 1. If the rotor initial position is Position 2 (shown in Fig.2 (b)), Fig.4 shows the state of the motor while a positive current pulse is supplied to the winding. Fig.4 shows the rotor pole N and the tooth A1 have the opposite polarities. And the relationship between the two MMFs is shown in Fig.4 (b). Obviously, the current pulse results in rotor clockwise rotation. The pulse keeps on, and the rotor will rotate an electrical angle of\u03b8 degree. Then the rotor will stop after vibrating for a moment at the special position (shown in Fig.5 (a)), where the polarities of the stator tooth and rotor pole are opposite and the rotor is attracted by the stator. Fig.6 illustrates the rotor position and speed during this procedure. In the figure, S1 is the rotor position curve, whose zero position is shown in Fig.5 (a), and W1 is the rotor speed curve. From Fig.6, it is seen that the rotor will stop after t1. Then if controller lets the current decrease to zero slowly at t3 (t3>t2>t1), the rotor yields a counterclockwise rotation, as shown in Fig.5 (b), and eventually offsets an electrical angle of \u03b8 degree and stops at Position 2 shown in Fig.2 (b) under the interaction of cogging torque and electromagnetic torque. That is, after the positive current pulse acts, the rotor also stops at Position 2 if the rotor initial position is Position 2. Similarly, if the controller uses such a negative current pulse to excite the winding, the rotor will stop at Position 1shown in Fig.2 (a) in spite of the rotor initial position. It is emphasized that the current pulse should be decreased to zero slowly after t3 shown in Fig.6. Otherwise, the rotor may vibrate seriously, even deviate from the expected position and stop undesiredly, which eventually results in failure of start-up. All above tells that the rotor stops at the special position determined by the polarity of the current pulse. After fixing the rotor position, the controller can control the motor to rotate in the desired direction easily. For example, if controller knows that rotor position is Position 2, it can use a negative current to let rotor yield counterclockwise rotation or a positive current to let rotor yield clockwise rotation. Of course, the said clockwise rotation may finish quickly if without commutating. But the time is always long enough for the controller to detect the back EMF, which is the key during commutating, and commutate the motor. The key of this method is the time length and the magnitude of the current pulse. The criterion of the former is that the time should be longer than t2 shown in Fig.6. And the criterion of the latter is that the current can drive the rotor rotating in the different rotor initial condition. Obviously, different motor needs different current pulse to start the motor. IV. STARTING METHOD BASED ON MOTOR COGGING TORQUE The starting method based on motor cogging torque is simpler. But it just can make the motor run in one direction determined by the structure of the asymmetrical air gap. From Fig.6, with the positive current pulse exciting the winding, the time t1 (in which, the rotor rotates from Position 2 shown in Fig.2 (b) to the zero position shown in Fig.5 (a)) is much shorter than t2 (in which, the rotor rotates from Position 1 shown in Fig.2 (a) to the zero position). So, if the controller lets the current decrease to zero suddenly as soon as t1 arrives, it is found that the rotor will rotate in the Fig.6 Waveforms of the current, rotor position and speed while starting the motor based on fixing rotor position same direction in spite of the rotor initial position as the following analysis: Fig.7 illustrates the waveforms of current, rotor position and speed during this operation. In this figure, time point t5 is just little longer than t1. At this time point, with the positive current pulse having excited the winding, the rotor has stopped at the special position (\u03b8 =0), as shown in Fig.5 (a) if the rotor initial position is Position 2 (shown in Fig.2 (b)). Then if current decreases to zero suddenly, the cogging torque will drive the rotor rotate counterclockwise, as shown in Fig.5 (b). S1 and W1 illustrate the rotor position and speed waveforms during this operation in Fig.7. If the rotor initial position is Position 1 (shown in Fig.2 (a)), the rotor will be still rotating in the counterclockwise direction at t5. So when current decreases to zero suddenly at t5, the rotor will keep on rotating counterclockwise because of the rotor inertia. S2 and W2 illustrate the rotor position and speed waveforms under this condition in Fig.7. That is to say, in spite of the rotor initial position, after the positive current pulse finishes its action at t5, the rotor will rotate in the same direction, a counterclockwise direction. Because of time t1 is very short, this positive current pulse is called as short positive current pulse. Similarly, after using a short negative current pulse exciting the motor, the rotor will also keep on rotating in the counterclockwise direction in spite of the rotor initial position. The effect is the same as using a short positive current pulse. But the reasons are just exchanged. That is, here, the reason for the counterclockwise motion is the cogging torque when rotor initial position is Position 1, while the rotor inertia is the reason when rotor initial position is Position 2. Nevertheless, in spite of the rotor initial position, this kind of short current pulse results in the rotor\u2019s rotation also in the counterclockwise direction, which is only determined by the structure of the asymmetrical air gap. If left air gap is made larger than the right air gap under the stator tooth A1 (shown in Fig.2 (a)) and right air gap is made larger than the left air gap under the stator tooth A2 (shown in Fig.2 (a)), the rotor will run in the clockwise direction after action of short current pulse. After the short current pulse has succeeded in making rotor run in the same direction with a speed, the controller can realize the commutation by analyzing the information of the motor back EMF, which is observable only when motor speed is not zero. V. EXPERIMENTS AND RESULTS AFC0912DE series made by DELTA ELECTRONICS, INC. Its rated voltage and current are 12V and 2.5A respectively. The MCU applied here is PIC12F675, which has an analog-to-digital converter of 4 channels embeded. Fig.9 and Fig.10 show the current waveforms with different starting methods when using Hall-less commutation methods. The starting method used in Fig.9 is Fig.7 Waveforms of the current, rotor position and speed while exciting the motor with short current pulse Fig.9 Current waveform under starting based on fixing rotor position based on fixing rotor position with the first current pulse in this figure decreasing to zero little by little. In Fig.10, the starting method is based on the cogging torque, obviously, the first current pulse here decreasing to zero suddenly. The experiments prove that both the starting methods could solve the motor starting problems successfully." ] }, { "image_filename": "designv11_61_0002542_00029890.2009.11920919-Figure17-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002542_00029890.2009.11920919-Figure17-1.png", "caption": "Figure 17. Special cases of Example 2 with positive k, and \u03bb = 0.3 in (a)\u2013(c).", "texts": [ " The dashed line in Figure 16a shows the case k = 1 with A = 2, B = 3. In this case the envelope \u03b5 is the parabola described by \u221a x/2 + \u221a y/3 = 1. The symmetric case A = B (Figure 16d) yields a known parabolic envelope [6, p. 76]. Figure 16b shows the case k = 2 (an ordinary ellipse with A = 2, B = 3) whose envelope is an asymmetric astroid. February 2009] THE TRAMMEL OF ARCHIMEDES 129 Example 2. G(x, y) = y \u2212 xk (kth power function). In this case the trace is another kth power function given by y = \u03bc ( x \u03bb )k . Figure 17a shows the case k = 1, in which the governor is the line y = x , the trace curve is the line y = \u03bcx/\u03bb, which is never tangent to the trammel, and there is no envelope. If k = 1, the envelope is yet another kth power function given by y = ( k\u22121 k )k 1 \u2212 k xk . Moreover, in this case the envelope touches all the trace curves only at the origin. The line through the trammel is always tangent to the envelope in the fourth quadrant. Figure 17d shows a special feature that holds for any exponent k. The trammel and the trace have the same slope at their point of intersection when (19) is satisfied. If n = mk then dn/dm = kmk\u22121 = kn/m, and (19) implies n m ( \u03bck \u03bb + 1 ) = 0. (26) Relation (26) is satisfied if n/m = 0 (when the trammel is horizontal) but (26) also holds for a nonhorizontal trammel when the point of subdivision (\u03bbm, \u03bcn) on the trammel satisfies \u03bb/\u03bc = \u2212k, or \u03bb = k/(k \u2212 1), k = 1. For this choice of \u03bb, the single trace \u03c4(\u03bb) coincides with the envelope of the family of traces, and a line through the trammel is tangent to both in the fourth quadrant as shown in Figure 17d. Figure 18 shows special cases of Example 2 when the exponent k is negative. In Figure 18a, b, c we have k = \u22121, k = \u22122, k = \u22123, respectively. The exceptional case 130 c\u00a9 THE MATHEMATICAL ASSOCIATION OF AMERICA [Monthly 116 (26) is shown in Figure 18d. For the special choice of ratio \u03bb/\u03bc = \u2212k, the trace \u03c4(\u03bb) coincides with the envelope and the trammel is tangent to both in the first quadrant as indicated in Figure 18d. The governing curve, each trace, and the envelope in Figure 18a are rectangular hyperbolas with the axes as asymptotes" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000828_iros.2003.1248883-Figure15-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000828_iros.2003.1248883-Figure15-1.png", "caption": "Fig. 15. fwt Joint angles and the reference trajectory of the", "texts": [ " APPENDIX: PLANNING THE REFERENCE TRAJECTORY AROUND THE PITCH AXIS The reference trajectories of joints 3, 4 and 5 are determined by the position of the foot. Let x and z be the position of the foot in the plane XZ which is perpendicular to the pitch axis, the reference trajectoly of the foot is given by, XF == :COS(@F), ZF = -H+/csin(@'), B xs = --cos(I$S), 2 U = -H, where ( X F , Z F ) and (xs,zs) are the positions of the foot in the free and support phase, respectively, H is the length from the ground to the joint 3, p is the step length, and h is the maximum height of the foot from the ground (Fig. 15). When the position of the foot is deterniml, the angle of each joint io be realized is calculated by the inverse kinematics as follows, k e, = - + atan2(z,x) - a t a d ( k , j ? +z2 + L: - 9) e, = -(&+e4), 2 e, = o r a n l ( k , l + t - L : - G ) where k is given by the following equation, k = \\ / ( x 2 + zZ + L: +L:)z - 2{(x2 + z ~ ) ~ +Li +q} In this research, the value of each parameter is set as follows: H = I8S[mmj, h = 81mm1, W = 13[degl, L1 = 100[m], L2 = loo[\"]." ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002626_1.2890382-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002626_1.2890382-Figure1-1.png", "caption": "FIG. 1. Example 1.", "texts": [ " Example 1: A homogeneous disk of mass m and radius R is moving in a plane and inelastically collides with a smooth line. The space-time bundle V is a four-dimensional manifold described by local admissible coordinates t ,x ,y , where x ,y are the Cartesian coordinates of the center of the This article is copyrighted as indicated in the article. Reuse of AIP content is subject to the terms at: http://scitation.aip.org/termsconditions. Downloaded to IP: 84.88.136.149 On: Tue, 02 Dec 2014 09:21:29 disk in the plane of the motion and is the orientation of the disk see Fig. 1 . The vertical scalar product on V is described by the matrix gij =diag m ,m ,A with A = 1 2mR2. The constraint S is given by the condition y=R or, choosing admissible coordinates t ,x , in S, by the injection i:S \u2192 V such that t,x, t,x,R, . 15 Then codim S =1. The velocity and vertical spaces are i* J1 V = p = t + x\u0307 x + y\u0307 y + \u0307 , i* V V = U = x\u0307 x + y\u0307 y + \u0307 , i* J1 S = q = t + x\u0307 x + \u0307 , i* V S = V = x\u0307 x + \u0307 , i* V S = W = w y . 16 Given a left velocity pL= / t+ x\u0307L / x + y\u0307L / y + \u0307L / , we have P S pL = t + x\u0307L x + \u0307L , VS pL = y\u0307L y " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003240_1.5061616-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003240_1.5061616-Figure1-1.png", "caption": "Figure 1 Schematic diagram of coaxial nozzle structure under investigation in this report with gas flows indicated.", "texts": [ " Considering the complexity and difficulty involved in solving a coupled multi-physical model, a series of independent models using various numerical tools (CFD/FLUENT; Analytical/MATLAB; FEA/ABAQUS) have been designed to simulate a decoupled system, with the objective of gaining a more fundamental understanding of the processes involved. With such a simplification, the relationship between the major LDMD process parameters and the physical interactions in play can be derived, so as to provide information for a coupled multiphysical model. In this work, a specific, commercially available, coaxial powder delivery nozzle, with its principle of operation resembling that shown in Figure 1, was chosen for the modelling. In this nozzle, the focusing laser beam and the \u2018nozzle\u2019 gas (argon), pass through the central axis of the nozzle assembly. The powder is delivered on a flow of \u2018carrier\u2019 gas (argon), to and through the nozzle. The powder flow is constrained by the dimensions of the nozzle to exit and come to a \u2018focus\u2019 at a particular (fixed) point with respect to the nozzle exit. The laser beam focus can be positioned with respect to the powder focus by moving the vertical position of the beam focusing lens" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003984_imece2013-62657-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003984_imece2013-62657-Figure1-1.png", "caption": "Figure 1 Cross section diagram of dual rotor system", "texts": [ " A parametric modeling method based on ANSYS software was proposed to model the spline joint. Then the test results of spool\u2019s stiffness were compared with two types of finite element models, contact element model and parametric model. At last, the distribution characteristics of the stiffness and the modal frequencies were analyzed. A typical rotor system, which consists of components with distinct functions and different materials, commonly used in high thrust ratio turbo fan engine is shown in Figure 1. Bolted joints and spline joint are used to assemble the components, meanwhile, they transmit torque and energy from the turbine to compressor. It is important to point out that the basic difference between the joint and the integral structure is the continuity of the interface. The interface could merely endure pressure stress, but tensile stress. And the contact stress is closely related to the contact state, which may be far open, near contact, sliding and sticky. For different design requirements, many types of spline joints are proposed" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003631_gt2012-69647-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003631_gt2012-69647-Figure1-1.png", "caption": "FIGURE 1: VIEW OF AN INTEGRAL GEAR COMPRESSOR (MAN DIESEL & TURBO SE).", "texts": [ " The present study seeks to detect and analyse correlations between input and output parameters and to understand the connection between the compressor and its gear, as well as to introduce meta models in the design process of radial gear compressors. Integral gear compressors [3] belong to the category of multi-shaft compressors. Characteristic for an integral gear compressor is the big gear - the bullgear - in the centre of the machine, which drives the circumjacent pinions, as can be seen in figure 1. On each shaft\u2019s end, a radial stage can be deployed. This design has the advantage, that the rotation speed of each pinion can be selected individually. Furthermore, it is possible to cool down the medium between two stages. On account of the variable rotation speeds, every stage can operate in a comfortable speed range. The intercooler leads to a convergence of the compression process with the isothermal compression. To enhance the global efficiency, either the compression process has to be improved or the losses of the gear box have to be reduced" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002541_1.2967884-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002541_1.2967884-Figure3-1.png", "caption": "Fig. 3 Thread root geome", "texts": [ " Assuming that the thread root radius is part of a single circle with radius , the thread profile can be defined by dividing the cross section into three parts, namely, \u201cA-B: root,\u201d \u201cB-C: flank,\u201d and \u201cC-D: crest.\u201d The cross section profile of the external thread perpendicular to the bolt axis is obtained by expanding the thread configuration of one pitch height into the plane, as shown in Fig. 2. It is worth noting that the cross section profile perpendicular to the bolt axis is identical at any position along the axis and can be expressed in terms of radius r shown in Fig. 1. Figure 3 a shows the geometry around the thread root in detail. The coordinate of a point on the arc BAB that forms the thread root radius is expressed as r , P /2 . The root radius has an upper limit restricted by a minor diameter d1. Figure 3 b displays the case of the maximum root radius max. The radial coordinate of the points B or B , where the thread flank becomes the tangent of the arc, coincides with that of minor diameter d1. Considering the symmetry between 0 and \u2212 0, the cross section profile shown in Fig. 2 is expressed as follows: r = d 2 \u2212 7 8 H + 2 \u2212 2 \u2212 P2 4 2 2 0 1 H + d 2 \u2212 7 8 H 1 2 d 2 2 1 1 = 3 P , 2 = 7 8 , 3 12 P, H = 3 2 P where d, P, and H represent the nominal diameter, pitch, and height of a fundamental triangle" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000112_j.bmc.2003.06.004-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000112_j.bmc.2003.06.004-Figure1-1.png", "caption": "Figure 1. Sketch of the proposed release system (not to scale).", "texts": [ " If we are to approach ionic signal propagation in an artificial system, we must have controlled insertion of a defined amount of compound into a defined patch of membrane with preservation of orientation. This means a short diffusion step through the aqueous phase to allow the hydrophobic-hydrophilic balance in the released compound to produce the self-assembled active structure with the correct orientation in the membrane. Consequently we are interested in ways to produce a pulse of a few thousand molecules at a defined point within a few microns of the membrane surface. Our conceptual design is illustrated in Figure 1. We envisage a gold microelectrode on a micromanipulator placed in proximity to the bilayer. The active transporter would be immobilized on the surface of the electrode via a gold\u2013thiol interaction. At the time of release, the Au\u2013S bond would be reductively cleaved, releasing the transporter into solution. Diffusion from this high local concentration should be fast and some fraction of the released material should insert and produce the required oriented and functional ion channel. Reductive release from gold\u2013thiol self-assembled monolayers (Au\u2013S SAMs) has been studied electrochemically", "2 V versus. SCE. The monolayer packing, the electrolyte, and the gold crystal face all influence the reduction potential, so for polycrystalline gold a range of potentials is common. The released thiolate can reabsorb once the electrode potential shifts below a reduction threshold, and there is evidence that poorly soluble thiols/thiolates remain with the electrode after release.20 Apart from reports of a \u2018thiol odor\u2019, the nature of the products has not been investigated. With respect to the proposal of Figure 1 there a number of unknowns: What species is (are) released into bulk solution (thiol, thiolate, disulfide)? What type of control is available (electrode area, reduction time)? What are the quantitative aspects (current efficiency, does released material simply reattach)? In related work on reductive cleavage of thiobuytrate esters, we have shown that the release step could potentially be made irreversible through intramolecular thiolysis of the ester.24 This \u2018traceless linker approach\u2019 would have the advantage that the ion channel would not need to be a thiol. Our goals in the current study are to focus on the principal unknowns as listed above. We examine the electrochemistry of an Au\u2013SAM in a \u2018macroscopic\u2019 system in order to detect, identify, and quantify the products of reductive cleavage. We then turn to a \u2018microscopic\u2019 experiment in which a gramicidin derivative is released in a proof-of-concept experiment as sketched in Figure 1. Some quantitative detail will distinguish the \u2018macroscopic\u2019 and \u2018microscopic\u2019 experiments. A well-ordered SAM from an alkyl thiol on single-crystal Au(111) has 1014 molecules per cm2.25 The macroscopic experiment uses an electrode area of 1.5 cm2 in contact with an electrolyte volume of 1.0 mL. Full release to the bulk solution is expected to give solution concentrations in the range of 10\u2013100 nM. Provided a suitable fluorophore is present in the release product(s), this level is appropriate for analysis by HPLC with fluorescence detection", " There are two possibilities: 6 is completely inactive, and we observe only the residual gramicidin in this experiment, or the channels formed by 6 and gramicidin are indistinguishable under these conditions. Other gramicidin derivatives modified as 6 at the reduced C-terminal position produce channels,32,33 so it is likely that the channels formed by 6 are sufficiently similar to those from native gramicidin as to be indistinguishable by conductance measurements. Finally, the experiment envisaged by Figure 1 was done using a 60-mm diameter Au wire as a microelectrode. Open-circuit deposition of 6 on the electrode tip was followed by a thorough washing sequence including a period of sonication in ethanol. We are quite confident that native gramicidin would have been removed from any surface under these washing conditions. The electrode was mounted on a micromanipulator and positioned within 50 mmof a bilayer formed in a bilayer clamp cell as assessed visually (binocular microscope) through a comparison of the electrode diameter and the distance to the bilayer" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002868_robot.2008.4543257-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002868_robot.2008.4543257-Figure1-1.png", "caption": "Fig. 1. Leader-follower setup", "texts": [ " IV the leader-to-follower range estimator is presented. In Sect. V an input-to-state feedback control scheme is designed and the stability of the closedloop system is analytically proved via Lyapunov arguments. In Sect. VI simulation experiments confirm the effectiveness of the proposed designs. Finally, in Sect. VII the major contributions of the paper are summarized and future research lines are highlighted. 978-1-4244-1647-9/08/$25.00 \u00a92008 IEEE. 504 The setup considered in the paper consists of two unicycle robots (see Fig. 1). One robot is the leader, whose control input is given by its translational and angular velocities, uL = [vL \u03c9L]T . The other robot is the follower, controlled by uF = [vF \u03c9F ]T . Each robot is equipped with an omnidirectional camera, which constitutes its only sensing device. Using well-known color detection and tracking algorithms [15], the leader is able to measure from the image, both the angle \u03b6 (w.r.t. the camera of the follower) and the angle \u03c8 (w.r.t. the colored marker P placed at a distance d along the follower translational axis) (see Fig. 1). Analogously, the follower can compute the angle \u03bd using its panoramic sensor. Note that the measurement of both the angles \u03b6 and \u03c8 by the leader, could not be a trivial task in practice, especially when the robots are distant. This problem has been addressed and solved in [9], where only the angle \u03b6 needs to be computed. As first shown in [6], the leader-follower kinematics can be easily expressed using polar coordinates [\u03c1 \u03c8 \u03d5]T , where \u03c1 is the distance between the leader and the marker P on the follower and \u03d5 is the relative orientation between the two robots, i.e the bearing. It is easy to verify that \u03d5 = \u2212\u03b6 + \u03bd + \u03c0. (1) Proposition 1 ([8]): Consider the setup in Fig. 1. The leader-follower kinematics can be expressed by the driftless system, \u03c1\u0307 \u03c8\u0307 \u03d5\u0307 = cos \u03b3 d sin \u03b3 \u2212 cos\u03c8 0 \u2212 sin \u03b3 \u03c1 d cos \u03b3 \u03c1 sin \u03c8 \u03c1 \u22121 0 \u22121 0 1 vF \u03c9F vL \u03c9L (2) where \u03b3 \u03d5 + \u03c8. In order to simplify the subsequent derivations and without losing in generality, we will only consider formations with a single follower (nevertheless, the results of this paper can be immediately extended to the general case of n followers [8]). The information flow between the robots is now briefly described" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003215_iecon.2008.4758119-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003215_iecon.2008.4758119-Figure3-1.png", "caption": "Fig. 3. Fault #2, hole in the outer race.", "texts": [ " This condition is close to the complete breaking of the bearing, so it cannot be considered as an incipient fault, but as an actual fault. Therefore, if a diagnostic tool does not succeed in detecting this fault, probably it will not able to detect any cyclic fault in the outer ring. Second, since in other researches one of the most investigated fault is the hole in the outer race [6] [9], we have reproduced also this single-point defect in order to have a comparison with the results obtained in the literature . In fact, a \u201cperfect\u201d fault like that shown in Fig. 3 cannot occur during the working of a bearing, but a similar defect can be caused by circulating currents, so this hole represents a magnification of defects that can really happen in the bearings. Then, we have investigated a fault which has not be considered in previous researches, i.e. a deformation of the protective shield (Fig. 4). This fault can be produced by errors during the assembly and can be considered as a cyclic fault, even if it does not produce effect like air-gap eccentricity. So, it is expected not to show particular changes in the current spectrum" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001753_cdc.2005.1582254-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001753_cdc.2005.1582254-Figure1-1.png", "caption": "Fig. 1. Dead-zone model", "texts": [ " These nonlinear characteristics have break-points so that they are non-differentiable (nonsmooth) but can be parameterized as in [22]-[24]. The parameterized model of the dead-zone characteristic DZi(.) can be unified as DZi(xi) = \u23a7\u23a8 \u23a9 mri(xi \u2212 bri) xi \u2265 bri 0 bli < xi < bri mli(xi \u2212 bli) v(t) \u2264 bli (2) where bri \u2265 0, bli \u2264 0 and mri > 0,mli > 0 are constants and xi is the input. In general, neither the break-points bri, bli nor the slopes mri,mli are equal. A graphical representation of the dead-zone is shown in the following Fig. 1. To parameterize the dead-zone, we introduce the following new smooth continuous functions \u03c6ri(xi) = exi/e0 exi/e0 + e\u2212xi/e0 (3) \u03c6li(xi) = e\u2212xi/e0 exi/e0 + e\u2212xi/e0 (4) where e0 is any positive constant. As e0 \u2192 0, \u03c6ri approaches a step transition from 0 at xi = \u2212\u221e to 1 at xi = +\u221e continuously and \u03c6li approaches a step transition from 0 at xi = +\u221e to 1 at xi = \u2212\u221e continuously. Remark 1: Note that the functions \u03c6ri and \u03c6li are continuous and differentiable. Compared with the nonlinearity functions in [22]-[24], the inverse indicator functions are not differentiable (non-smooth) so they cannot be used in the recursive backstepping design which requires the function to be continuous differentiable" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002513_detc2008-49084-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002513_detc2008-49084-Figure2-1.png", "caption": "Fig. 2 Pencil of meridians with slope angle \u03b5.", "texts": [ " 4). L = set of all the RS great circles that are normal, at the contact point, to two slipping curves of any slipping contact of an SM; Pij = IP of the relative motion between the i-th link and the j-th link of an SM; d From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?ur So = sequence which the secondary IPs can be determined with; Sp = set of the primary IPs of an SM; Ss = set of the secondary IPs of an SM; \u03b5 = slope angle (this angle identifies a pencil of meridians, see Fig. 2); \u03b8 = signed dihedral angle that locates one great circle in a pencil of meridians with assigned slope angle (see Fig. 2 and \u00a7\u00a7 2.1); (\u03be, \u03b6) = pseudo-Cartesian coordinates (these coordinates locate a point on the positive hemisphere of the RS, see Fig. 3); In spherical mechanisms (SMs), the instantaneous motion between two links is a rotation around an instantaneous rotation axis (instantaneous pole axis). The instantaneous (first-order) kinematics of SMs can be fully described by analyzing the positions of these axes [1]. Instantaneous pole axes (IPAs) have properties that are the spherical counterparts of instant centers' ones", " The distance between two points of the RS is defined as the convex central angle formed by the two radius vectors that pass 1 The great circle arc that joins two points on a sphere is the geodesic line and, in spherical geometry, it plays the same role as, in plane geometry, the straight-line segment that joins two points. nloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?u through the two points. Such a distance measured in radians ranges from zero to \u03c0. Any great circle of the RS intersects the IGC at two diametrically opposite points which identify a straight line lying on the yz plane (see Fig. 2). Such a line is uniquely located by giving the angle, \u03b5, between the y axis and the line. The angle \u03b5 (slope angle) is taken positive if counterclockwise with respect to the x axis. All the great circles that can be drawn on the RS can be collected into pencil of meridians which share the same intersections with the IGC. Such a property will be stressed by saying that the great circles belonging to the same pencil of meridians have the same slope. Provided that the IGC's points are excluded, all the great circles passing through a point of the RS can be identified by means of their slope angle, since only one great circle with an assigned slope passes through any point of the RS (the IGC's points are excluded)", " If the position of one out of these secondary IPs is arbitrarily assumed on the great circle it has to lie on, other secondary IPs can be coherently determined through the A.-K. theorem until a mismatch occurs: a secondary IP that should lie on a known great circle is determined out of this circle, that is, the IP is over-determined. This mismatch can be solved by analytically solving a system of n equations in n unknowns where the number, n, and the type of equations to write can be determined as it will be explained in the following section. Figure 2 shows a pencil of meridians with slope angle \u03b5. The great circles which belong to the same pencil of meridians can be located by giving the signed dihedral angle, \u03b8, they form with the great circle, of the same pencil, which passes through the intersection between the RS and the x axis (see Fig. 2). The angle \u03b8 will be assumed positive if clockwise with respect to the radius vector pointing toward the common intersection of the great circles of that pencil on the positive shell. So doing, it belongs to the range ]\u2212\u03c0/2, \u03c0/2] (rad). The ordered pair (\u03b5,\u03b8) uniquely locate a great circle in the positive hemisphere, whereas the ordered pair (tan \u03be, tan \u03b6) uniquely locate a point on the US through its gnomonic projection onto the RP. In the RP, an Argand diagram, with the origin at the point of tangency between RP and US, real axis parallel to the y axis, and imaginary axis parallel to the z axis, is introduced; and, hereafter, the RP's point, (tan \u03be, tan \u03b6), is located by giving the complex number p = tan \u03be + j tan \u03b6 where j denotes the imaginary unit 1\u2212 " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000117_135065003322620246-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000117_135065003322620246-Figure3-1.png", "caption": "Fig. 3 Examples of the mechanical linkage arrangements used to extract signals from reciprocating machinery: (a) as used by Westbrook and Munro [40] to extract inductive displacement signals from a transducer mounted in the side of a piston; (b) as used by du Parquet and Godet [42] to extract thermocouple signals from a big-end bearing; (c) as used by Jones [52] to extract thermocouple signals from a big-end bearing", "texts": [ " Also in 1967, Westbrook and Munro [40] published details of instrumentation they used to measure piston lateral motion during engine operation. Two systems were described. One made use of a VHF/FM radio telemetry link in which the transmitter electronics were mounted on the connecting rod and signal cables from the transducers mounted in the piston were fed to the transmitter on the connecting rod, the cables \u00afexing at the small-end bearing during engine operation. The second method described utilized a mechanical linkage connecting the big-end bearing cap to the stationary engine block, as shown in Fig. 3a (also shown in Fig. 3 are mechanical linkage arrangements used by other researchers). The disadvantage of both of these experimental approaches, however, is that they involve a signi\u00aecant additional mass being added to the Proc. Instn Mech. Engrs Vol. 217 Part J: J. Engineering Tribology J02701 # IMechE 2003 at Kungl Tekniska Hogskolan / Royal Institute of Technology on June 30, 2015pij.sagepub.comDownloaded from reciprocating components, thus upsetting the forces generated in the mechanism and thereby affecting the quantities that are being measured", " Comparison of measured oil \u00aelm thickness throughout the engine cycle with computed values based on a short bearing model [10] showed general agreement, although the measured oil \u00aelm thickness values were generally higher than the calculated ones. The authors found that theoretical results for a fully grooved bearing were closest to the measured values (which were for a bearing with a feed groove in the end cap only), and attributed the discrepancies to bearing distortion. In the same year, du Parquet and Godet [42] published measurements of connecting rod big-end bearing temperature obtained using thermocouples, with their cables supported by a linkage mechanism similar to that used by Westbrook and Munro [40] (see Fig. 3b). Their work related to a small, high-speed, fourstroke petrol engine, whereas the work of Goodwin and Holmes [11] related to medium-sized diesel engines operating at medium speeds. The authors instrumented their engine with six thermocouples equally spaced around the bearing. Results showed a temperature variation around the bearing circumference of up to 8 K and a difference between sump temperature and bearing temperature of up to 30 K. The additional mass of the linkage mechanism was of no consequence to the investigation reported, since the aim of the work was merely to compare temperature rises obtained with different grades of oil, and not to validate theoretical models", " The authors also published theoretical oil \u00aelm thickness values obtained using the short bearing approximation coupled with the mobility method, results showed that theoretical oil \u00aelm thickness values were smaller than measured values. Measurements of engine crankshaft bearing temperature distribution for various bearing geometrical con\u00aegurations were published by Jones [52]. Jones carried out measurements on a 2.0 l, four-cylinder, in-line gasoline engine. To do this, he used thermocouples mounted in the connecting rod big-end bearing cap, and led wiring from these, via a mechanical linkage assembly (see Fig. 3c), to the engine block, and from there to a data recorder. Four measurement points were used, equally spaced around the bearing and at 458 to the bearing/cap split. Jones concluded that increased oil feed temperature, reduced bearing clearance and incorporating an oil feed groove in the highly loaded region of the bearing all lead to increased bearing temperature, although reducing the overall width can lead to a reduced bearing temperature. His results also showed that the temperature range around the bearing circumference can be up to 8 K, and the mean bearing temperature can be over 30 K higher than the oil feed temperature at high engine speeds", " Figure 2d is reprinted from STLE Tribology Transactions, Vol. 33, 2, by Bates et al. `Oil \u00aelm thickness is an elastic connecting rod bearing: comparison between theory and experiment\u2019, pp. 254\u00b1266, 1990, with permission from STLE. Proc. Instn Mech. Engrs Vol. 217 Part J: J. Engineering Tribology J02701 # IMechE 2003 at Kungl Tekniska Hogskolan / Royal Institute of Technology on June 30, 2015pij.sagepub.comDownloaded from Figure 2e is reprinted with permission from SAE 930694 # 1993 Society of Automotive Engineers. Figure 3a is reprinted from Trans. ASME, J. Engng Power, April, 1967 by Westbrook and Munro, `The telemetering of information from a working internal combustion engine\u2019, pp. 247\u00b1254, with permission from ASME. Figure 3b is reprinted with permission from SAE 780980 # 1978 Society of Automotive Engineers. Figure 3c is reprinted with permission from SAE 960989 # 1996 Society of Automotive Engineers. Figures 5a and 5b are reprinted with permission from SAE 892114 # 1989 Society of Automotive Engineers. Figure 5c is reprinted from STLE Tribology Transactions, Vol. 35, 1, by Masuda et al. `A measurement of oil \u00aelm pressure distribution in connecting rod bearing with test rig\u2019, pp. 71\u00b176, 1992, with permission from STLE. 1 Grente, Ch., Ligier, J. L., Toplosky, J. and Bonneau, D. The consequence of performance increases of automotive engines on the modelisation of main and connecting rod bearings" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001953_ijcmsse.2007.017926-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001953_ijcmsse.2007.017926-Figure6-1.png", "caption": "Figure 6 Liquid fraction contour for Model II, case 5 with scanning speed 20 mm/s, beam diameter 2 mm and laser power 2 kW (a) top surface and (b) mid-cross section", "texts": [ " Owing to the Marangoni convection, the high temperature fluid penetrates deeper into the work piece, which results in high depth in Model II than Model I. This Marangoni convection also lowers the temperature gradient in Model II. The molten metal is pulled from the relatively cooler pool edges towards the centre of the pool, resulting in counter-rotating vortices as shown in Figure 4. The weld pool is deeper and narrower when convection due to Marangoni effect is considered. The liquid fraction contour with scanning speeds 20 mm/s for Models I and II are shown, through Figures 5 and 6, respectively. Model II which includes the fluid flow (Figure 6) shows low weld pool width, low weld pool length and high weld pool depth than those without fluid flow (Figure 5). It indicates that the Marangoni flow plays an important role in transferring the heat from the top surface of the workpiece to the interior region of the domain. The heat is quickly transferred to the solid\u2013liquid boundary at the bottom of the of weld pool. Moreover, the molten metal in the rear of pool becomes cold and it solidifies more quickly due to dissipation of heat. Hence, the pool becomes deeper and narrower when Marangoni effect is considered. Figure 6(b) shows the molten pool profile of the cross section considering fluid flow. The simulation without considering flow (Figure 5) does not predict the weld pool profile in agreement with the results observed in practice. Especially, the simulated weld pool widths are too high with shallow depth. In Model I higher heat transfer takes place along the horizontal direction and in Model II higher heat transfer takes place along the vertical direction. The maximum pool depth for Model I is 0.55 mm and for Model II it is 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000646_006-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000646_006-Figure6-1.png", "caption": "Figure 6. Free-body diagram for the upper pulley. The upper pulley has negligible mass.", "texts": [ " Knowing the accelerations a2, a3 and a5, we can now calculate the motion of the centre of mass of the system. In one second the change of position of each mass is yi = 0.5ai(1 s)2. (12) Therefore in one second y2 = 1.1 m, y3 = \u22120.9 m and y5 = \u22120.1 m. At the initial time m2, m3 and m5 are all located at y = 0. Therefore the original position of the centre of mass is 0. After one second our calculated changes in position give a new centre of mass position of \u22120.1 m, a change of 0.1 m. Let us now look at the net forces acting on the entire system of pulleys and masses. Figure 6 shows the forces acting on the upper pulley, which rotates but does not translate. Therefore its acceleration is zero. We can write P \u2212 2T2 = 0 (13) which gives P = 2T2 = 96 N. (14) Figure 7 shows the forces acting on the entire system. There is an upward force P of 96 N and a total downward gravitational force of (10 kg)g or 98 N. This yields a net downward force of \u22122 N, meaning that the system is not in equilibrium. Its acceleration is \u22122 N/10 kg or \u22120.2 m s\u22122. Using equation (12) gives a change in the centre of mass position after one second of \u22120" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003917_icsem.2010.22-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003917_icsem.2010.22-Figure2-1.png", "caption": "Figure 2 the modified direction of toroidal worm", "texts": [ " The tooth surface of worm along the direction of tooth ght consists of numerous modification curves. Fig. 1 shows that every modification curve represents the tooth file of worm on the special position. So, to study odification curve is actually to discuss the tooth profile ong the helical line. The reference circle lies on the middle he working tooth, so its modification curve is most representative and usually regarded as the representative of the tooth surface of worm. In the following sections, the positive direction? Fig. 2 shows that the point M and N all lie on the original tooth surface. To point N, there may be two directions after modifications which are de and \u2013de. Are both directions satisfied with the meshing conditions? This is what the modification theory will solve. principle of the toroidal worm. Because the derivation is based on the curvature radio, the principle is also named as \u201cthe curvature modification\u201d. The meshing equation of the kinematic method is the necessary condition of the meshing of the conjugate hyperboloids on theory" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000717_kem.297-300.102-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000717_kem.297-300.102-Figure1-1.png", "caption": "Fig. 1 Normal vector, stress vector, strain increment vector and a (virtual) crack plane", "texts": [ " Thus, it is necessary to calculate not the energy stored in a structure but the energy released due to crack initiation or growth. In other words, the plane of a crack should be taken into account in the calculation of the released energy. CED is only a part of the strain energy density which is to be released as a crack initiates or grows, and it depends on the crack plane. Note that CED can be defined not only for an actual crack plane but also for a (virtual) plane where a crack is likely to occur. In Fig. 1, a stress vector, a strain increment vector, and a normal vector of a crack plane are shown. The stress vector (or the strain increment vector) can be readily obtained from a linear transformation of the normal vector by the stress tensor (or the strain increment tensor). Then, the CED increment can be calculated from the dot product between the stress vector and the strain increment vector as shown in Eq. (1). \u222b= = = \u03b5 \u03b5\u03c3 \u03b5\u03c3 \u03b5\u03c3 0 rr rr rr dW ddW rdrdW T (1) Here, \u03c3 is the stress tensor, \u03b5d is the strain increment tensor, \u03c3 r is the stress vector, \u03b5 r d is the strain increment vector, and r r is the normal vector" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000073_s0094-114x(03)00005-3-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000073_s0094-114x(03)00005-3-Figure8-1.png", "caption": "Fig. 8. The T 3 manipulator\u2013\u2013a manipulator without a spherical wrist\u2013\u2013at the initial configuration, i.e. qi \u00bc 0; i \u00bc 1:6. The coordinate system is fixed with the link 4 whose length is h.", "texts": [ " Another point is that \u2018\u2018fill-in\u2019\u2019 generation (elements that were null and become non-null, i.e. become filled in) is minimal when the matrix is its finest block-triangular form [7], which requires less storage capability. Secondly, the determinant is easier to obtain in Eq. (25) than in Eq. (22). Each diagonal block in Eq. (25) for one term of the product which composes the determinant. For instance, det J \u00bc det J \u00bc \u00f0x14\u00de\u00f0gs3\u00de\u00f0 h\u00de det b c4 s4 s4 c4 \u00f01\u00de At last, we apply the method to a manipulator without a spherical wrist, the T 3 robot, Fig. 8. The Jacobian of this manipulator is extracted from [17]: _x1 _x2 _x3 _x4 _x5 _x6 2 6666664 3 7777775 \u00bc s234 0 0 0 0 c5 0 1 1 1 0 s5 c234 0 0 0 1 0 0 z24 gs4 0 0 0 x23 0 0 0 h 0 0 x34 gc4 0 0 hs5 2 6666664 3 7777775 zfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl}|fflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl{JT3 _q1 _q2 _q3 _q4 _q5 _q6 2 6666664 3 7777775 \u00f029\u00de where x23 \u00bc fc3 \u00fe gc23; x34 \u00bc fc34 \u00fe gc4; z24 \u00bc fs34 \u00fe gs4. The same reference [17] obtains the inverse of the Jacobian, and hence the structure of this inverse, and links screw theory with the columns and rows of both matrices. Fig. 8 represents the manipulator at the initial position with the screws represented at each joint. Using Algorithm 1 and extracting only the terms relative to the forward kinematics, as in Section 7 for the PUMA robot, the hierarchical canonical form of the Jacobian of Eq. (29) in a linear system form, becomes: \u00f030\u00de The singularities are: x23 hc234 \u00bc 0 ! fc3 \u00fe gc23 \u00fe hc234 \u00bc 0, c5 \u00bc 0, s3 \u00bc 0 in this particular order, from bottom to top. The partial order among the singularities x23 hc234 \u00bc 0 c5 \u00bc 0 s3 \u00bc 0, and again the wrist singularity precedes all other singularities" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001353_s1672-6529(06)60012-7-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001353_s1672-6529(06)60012-7-Figure3-1.png", "caption": "Fig. 3 Illustration of tester for adhesion and friction.", "texts": [ "5 mm, Goodfellow, Cambridge, England), sample plate (P), and lower samples of polyurethane (S, supplied by Dow Chemical, Germany) with diameter of 60 mm and thickness of 4 mm and was cut according to the required dimensions. The sensor G was driven by a motor to move down and up to load and unload. The deflection of the sensor (proportional to force) was recorded and processed by the computer. The elastic modulus of target material samples S 1, S2, S3 and S4 of polyurethane were measured to be 1206f80, 55f1.2, 93f1.0 and 104M13.6 kPa ( n = 4 ) respectively. ' 2.2 Multipoint contact: rough PU plane to flat glass The experiments were carried out on a test machine of adhesion and friction as shown in Fig. 3. The upper sample (glass, 10 mm in diameter) was glued to a cantilever, which included two dimensional sensors and was driven by a step motor. The normal force was detected and as feedback was sent to the motor to control it. The lower sample of polyurethane 10 mm x 10 mm x 4 mm (lengthxwidthx thickness), was fixed to a sample plate. Two series of experiments were performed on this tester: (1) relationship between adhesive force and normal force; (2) relationship between negative normal force (adhesive state) and tangential force" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003789_ijtc2011-61146-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003789_ijtc2011-61146-Figure3-1.png", "caption": "Figure 3: New design concept of mesoscale foil gas bearing based on two-piece sintering process of elastic foundation (Design II)", "texts": [ " The elastic foundation was redesigned to allow full LIGA process involving the micro metal powder sintering process. To facilitate the manufacturing process of the sintering mold insert and subsequent sintering processes, minimum feature size (i.e., thickness of the arc structure) had to be increased. However, large thickness of the elastic foundation structure results in very high stiffness if their geometry follows the same shape as Design I. To maintain similar stiffness to that of Design I, the new design adopted an inclined brush-like structure with larger feature size as shown in Figure 3(c) (will be denoted as \u201cDesign II\u201d hereafter). In addition, one set of elastic foundation is made of three pieces to allow a two-piece sintering method. 1 Copyright \u00a9 2011 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 2 Copyright \u00a9 2010 by ASME 18m off-centered bearing sleeves were manufactured through a precision electro-discharge machining (EDM), and 50 m thick Inconel top foil was manufactured through coldforming and age-hardening", " Figure 9 shows cross section of the assembled micro motor with foil bearings made of Design II elastic foundations, and Figure 9 shows the micro motor mounted on a pedestal for test. Two fiber optic probes measured rotor vibration along the vertical and horizontal directions. Cooling air for the bearings was provided from the rear end. Figure 10 presents waterfall plots of the rotor vibrations along the vertical directions up to 350,000 rpm. No subsynchronous vibration was observed up to the test speed. For the manufacturing of sintered elastic foundation of air foil bearing shown in Figure 3(c), Fe-45wt% Ni nano-sized powder (Aldrich, <100nm, 97%) and binder system consisted of paraffin wax, carnauba wax, bees waxes, polypropyleneethylene vinyl acetate (PP/EVA) blends, and stearic acid were used as nano-powder feedstock. Compacting and forming of the green mold was done using a two-piece type, uniaxial mold insert made of X-ray lithography. The mixing ratio of about 75vol% between the nanopowder and binder for the feedstock resulted in the best result, and sintering was performed at 1000oC (10oC/min) for 1 hour" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003762_s40430-013-0059-1-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003762_s40430-013-0059-1-Figure1-1.png", "caption": "Fig. 1 Continuum-based axisymmetric shell element: displacement nodes (master and slave) and stress nodes", "texts": [ " The circumferential direction, perpendicular to the axisymmetric half-plane, is represented by eu \u00bc er ^ ez: In the one-dimensional element i, there are three master nodes and three corresponding directors, respectively, denoted as: xa \u00bc raer \u00fe zaez; a \u00bc 1 :3; \u00f029\u00de pa \u00bc cos haer \u00fe sin haez; a \u00bc 1 :3: \u00f030\u00de The respective underlying continuum element has six slave nodes: xa \u00bc raer \u00fe zaez; a \u00bc s1:s6: \u00f031\u00de For convenience, we use the following two alternative label sequences to identify the slave nodes of element i (see Fig. 1): s1 : s6 f1 ; 1\u00fe; 2 ; 2\u00fe; 3 ; 3\u00feg: \u00f032\u00de Using this notation, we can relate master and slave nodes by the following formulas xa \u00bc xa ha 2 pa; a \u00bc 1 :3; \u00f033\u00de xa\u00fe \u00bc xa \u00fe ha 2 pa; a \u00bc 1 :3; \u00f034\u00de where ha is the shell thickness at node a. A fiber through the master node a and parallel to the director pa (a pseudo normal) moves rigidly. Thus, the velocities va; va and va\u00fe of master and slave nodes are related by the angular velocity _xaeu of the director pa as follows va \u00bc va ha 2 _xaeu ^ pa; a \u00bc 1 :3; \u00f035\u00de va\u00fe \u00bc va \u00fe ha 2 _xaeu ^ pa; a \u00bc 1 :3: \u00f036\u00de Considering (33) and (34), the above leads to va \u00bc va \u00fe _xaeu ^ xa xa\u00f0 \u00de; \u00f037\u00de va\u00fe \u00bc va \u00fe _xaeu ^ xa\u00fe xa\u00f0 \u00de: \u00f038\u00de We can gather (37) and (38) in symbolic matrix notation as va va\u00fe \u00bc Ti;a va _xa ; \u00f039\u00de where Ta;i :\u00bc 12 eu ^ xa xa\u00f0 \u00de 12 eu ^ xa\u00fe xa\u00f0 \u00de ; \u00f040\u00de with 12 denoting the 2D identity tensor", " x\u00f0n\u00dejBi \u00bc X a\u00bcs1:s6 ga\u00f0n\u00dexa; \u00f051\u00de with the symbol ga\u00f0n\u00de denoting the Lagrange interpolation functions g1 \u00bc \u00f01=4\u00den\u00f0n 1\u00de\u00f01 g\u00de; g2 \u00bc \u00f01=4\u00den\u00f0n 1\u00de \u00f01\u00fe g\u00de; g3 \u00bc \u00f01=2\u00de\u00f01 n\u00de\u00f01\u00fe n\u00de\u00f01 g\u00de; g4 \u00bc \u00f01=2\u00de\u00f01 n\u00de\u00f01\u00fe n\u00de \u00f01\u00fe g\u00de; g5 \u00bc \u00f01=4\u00den\u00f0n\u00fe 1\u00de\u00f01 g\u00de and g6 \u00bc \u00f01=4\u00den\u00f0n\u00fe 1\u00de\u00f01\u00fe g\u00de: Note that (51) is intrinsic. The curvilinear coordinate system ~R\u00f0n\u00de \u00bc fe~x; e~yg; used to enforce the hypothesis of zero transverse normal stress, is obtained employing the same method used by Belytschko et al. [3]. Accordingly, we define the base vectors e~x tangent and e~y normal to the lamina (see Fig. 1) as e~x :\u00bc x;n kx;nk \u00bc r;ner \u00fe z;nez r2 ;n \u00fe z2 ;n 1=2 ; \u00f052\u00de e~y :\u00bc eu ^ e~x \u00bc z;ner \u00fe r;nez r2 ;n \u00fe z2 ;n 1=2 ; \u00f053\u00de where subscript \u2018\u2018,n\u2019\u2019 denotes the corresponding derivative. Change of basis from ~R\u00f0n\u00de to R is accomplished using the general formulas for vectors and second-order tensors a\u00bd R\u00bc R a\u00bd ~R or A\u00bd R\u00bc R A\u00bd ~RRT \u00f054\u00de with R\u00f0n\u00de :\u00bc er e~x er e~y ez e~x ez e~y : \u00f055\u00de The Jacobian of the geometry mapping is given by J\u00f0n\u00de \u00bc rT x\u00f0n\u00de \u00bc X a\u00bcs1:s6 xa rga\u00f0n\u00de \u00f056\u00de or r;n r;g z;n z;g R \u00bc X a\u00bcs1:s6 raga;n raga;g zaga;n zaga;g R : \u00f057\u00de Further, the chain rule gives ga;n \u00bc ga;rr;n \u00fe ga;zz;n; \u00f058\u00de ga;g \u00bc ga;rr;g \u00fe ga;zz;g; \u00f059\u00de or rnga \u00bc JTrxga; \u00f060\u00de and thus rxga \u00bc J Trnga: \u00f061\u00de Finally, the derivatives of the interpolation functions with respect to spatial coordinates are obtained by substituting the inverse of the Jacobian in the equation (61), i" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003610_piee.1963.0011-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003610_piee.1963.0011-Figure3-1.png", "caption": "Fig. 3", "texts": [ " In a previous paper3 the author drew attention to the fact that Hall effect arising from the propagation of an electromagnetic wave through a semiconductor is the electrical counterpart of what we have here called the 'internal' radiation pressure. The problem was approached from the point of view of interaction between the high-frequency current in the material medium and the magnetic field of the wave, setting up in a mutually perpendicular direction a mechanical force and a corresponding electric field acting on the mobile carriers or the bound charges, as the case may be, within the semiconductor. It will be helpful to repeat very briefly the argument given. Referring to Fig. 3, and considering an elemental volume Similarly, the elemental mechanical force dSF d on that part of the displacement component of current associated with the material medium is = [iHJ'ddxdzdy where (21) (22) 3FC acts on the mobile carriers associated with conduction current in the medium and is equivalent to a Hall electric field Ec, both forces operating along the x-axis such that = mqcNcEcdxdydz (23) where m is a factor which takes account of the fact that the electric force on the individual charges qc differs from the applied field and iVc is the number of mobile carriers per unit volume" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002100_jmes_jour_1969_011_071_02-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002100_jmes_jour_1969_011_071_02-Figure7-1.png", "caption": "Fig. 7 . Overcut at centre distance a < (r, ,2-rbl)", "texts": [ " z,(4,,+inv q,l-inv Elimination of 6, results in -z2(+,,+inv c:,,,-inv (r,\u201dII) = 01 (15) r b 2 - r b l cos a,,, = a, rb2- r b l a,, = - cos %,I 1 (19) , At the new centre distance a,, new overcut may be expected. The conjugation of centre distances seems similar to the story of Achilles a, pursuing the tortoise a,,, but in our case the tortoise a,, goes to a well-defined point of maximum overcut and then creeps back until he meets Achilles a, in the point described by existing theory. CONJUGATION OF CENTRE DISTANCES BELOW a = ( r b 2 - r b l ) The simplest way to describe the co-operation of gears is to start with centre distance a = (rb2-rbl) and to examine the influence of a decrease of a, Fig. 7. In this way the kinematics of gear manufacture are fully accounted for. Suppose rule (12)not satisfied and let the centre distance be a,, a, < ( r b z - r b l ) . An arbitrary tooth of the pinion injures a tooth of the internal gear. The knocked out piece is circumscribed by a straight line up to a point with radius r,, and a curve determined by the pinion tooth. The said point determines an involute of the internal gear which will mesh with the pinion if the centre distance is increased to air. Vol I1 No 6 1965 at NANYANG TECH UNIV LIBRARY on June 5, 2016jms.sagepub.comDownloaded from OVERCUT, A NEW THEORY FOR TIP INTERFERENCE IN INTERNAL GEARS 587 r,,($,+inv = rbl(9,+inv . (20) the discrepancy in two domains for centre distances vanishes. Afterwards it is easy to conclude that the method of Fig. 7 is also operative for a > ( r h 2 - r h , ) . ry2 sin (vh+4h) = ral sin #1 a, = rU2 cos (&2+A$2)-r( l , cos $l . (22) A+2 = - 1nv %I1 1 . (21) . t j . (23) CONJUGATION O F CENTRE DISTANCES z2 - 21. '2 i REPRESENTED FOR p.2 > p a l (ra2-rul) < 01 < (ru2-rbl) J ' (24) The representation of the conjugate centre distance in Without detriment to general validity a simplification Fig. 8 confirms statement (12). The point a, = uI I = , / [ ( f a 2 - P u 1 ) 2 + ( t b 2 - T b 1 ) 2 ] relates to the case of Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001230_detc2005-84223-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001230_detc2005-84223-Figure1-1.png", "caption": "Figure 1: Illustration of \u2018skeleton\u2019 and actual witness points. Here \u2018 ' \u2019 indicates a skeleton point.", "texts": [ " (10), (11), and (12), determining the first and second derivatives of minimum distances requires the first and second derivatives of the witness points rP and oP . This section will discuss these derivatives for three simple types of robot and obstacle modeling primitives, a sphere, a cylisphere, and a quadrilateral plane3. The representations given below are not for actual surface points but instead for points on the skeleton of the given primitive (i.e. a point on the surface of a cylisphere is represented using the corresponding point on the axial line of the cylisphere). Primitives are represented in this way in order to simplify calculations (see Figure 1). A proof of the validity of representing primitives by their skeletons is given in [5]. 4 Copyright \u00a9 2005 by ASME l=/data/conferences/idetc/cie2005/72592/ on 07/22/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use D 7.1. SPHERES Spheres are represented by a single central point, 1P , and a radius, r. Therefore, the nearest distance point is always the same point on the sphere since the \u2018skeleton\u2019 for a sphere is a single point. The first and second derivatives of this point are 1 1 and P P " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003172_kikaic.74.1825-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003172_kikaic.74.1825-Figure5-1.png", "caption": "Fig. 5 Rotational manipulation", "texts": [], "surrounding_texts": [ "\u30d4\u30b6\u8077\u4eba \u306e\u30cf \u30f3 \u30c9\u30ea\u30f3\u30b0\u30e1\u30ab\u30cb\u30ba\u30e0 \u306b\u7740 \u76ee \u3057\u305f\u52d5\u7684\u64cd \u308a 1827\n3. \u554f \u984c \u8a2d \u5b9a\n3\u30fb1 \u30e2\u30c7\u30eb\u5316 \u56f33\u306b \u793a\u3059 \u3088 \u3046\u306a \u30d7 \u30ec\u30fc \u30c8\u3068\u5bfe \u8c61 \u7269 \u3092\u8003 \u3048\u308b.\u7c21 \u5358\u5316 \u306e\u305f\u3081,\u4ee5 \u4e0b \u306e\u4eee \u5b9a\u3092\u8a2d \u3051\u308b.\n\u4eee \u5b9a1: \u5bfe\u8c61\u7269 \u3068\u30d7 \u30ec\u30fc \u30c8\u306f\u525b\u4f53 \u3068\u3059 \u308b. \u4eee \u5b9a2: \u5bfe\u8c61\u7269 \u306e\u8cea \u91cf\u5206\u5e03 \u306f\u4e00\u69d8 \u3068\u3057,\u539a \u307f\u306f\u7121\u8996 \u3067\n\u304d\u308b\u307b \u3069\u5c0f \u3055\u3044.\n\u4eee\u5b9a3: \u30d7 \u30ec\u30fc \u30c8\u306f,\u5bfe \u8c61\u7269 \u304c\u843d\u4e0b \u3057\u306a\u3044\u7a0b\u5ea6 \u306b\u5341\u5206\n\u5927 \u304d\u3044.\n\u4eee\u5b9a4: \u5bfe\u8c61\u7269 \u3068\u30d7 \u30ec\u30fc \u30c8\u3068\u306e\u63a5\u89e6\u306f,\u5e38 \u306b\u5747\u4e00 \u306a\u9762\n\u63a5\u89e6 \u304c\u7dad \u6301 \u3055\u308c \u308b.\n\u4eee\u5b9a5: \u5bfe\u8c61 \u7269 \u3068\u30d7 \u30ec\u30fc \u30c8\u9593\u306e\u6469\u64e6\u4fc2 \u6570 \u306f\u4e00\u69d8 \u306b \u03bc\n\u3068\u3057,\u9759 \u6469\u64e6 \u3068\u52d5\u6469\u64e6 \u306f\u533a\u5225 \u3057\u306a\u3044.\n\u4eee\u5b9a6: \u5bfe\u8c61\u7269\u304a \u3088\u3073\u30d7 \u30ec\u30fc \u30c8\u306e\u4f4d\u7f6e \u30fb\u59ff\u52e2 \u306f\u30bb \u30f3\u30b5\n\u7cfb\u304b \u3089\u5f97 \u3089\u308c \u308b.\n\u56f33\u5185 \u306b\u793a\u3059\u8a18 \u53f7\u306e\u610f\u5473\u306f,\u4ee5 \u4e0b\u306e\u901a \u308a\u3067 \u3042\u308b.\n\u03a3R: \u57fa \u6e96\u5ea7\u6a19 \u7cfb.\u305f \u3060 \u3057,xR-yR\u5e73 \u9762 \u306f\u6c34\u5e73 \u3068\n\u3059 \u308b.\n\u03a3m: \u30d7 \u30ec\u30fc \u30c8\u306b\u56fa\u5b9a \u3055\u308c \u305f\u5ea7\u6a19 \u7cfb. \u03a3B: \u5bfe\u8c61\u7269 \u306e\u91cd\u5fc3\u4f4d\u7f6e \u306b\u56fa\u5b9a \u3055\u308c\u305f\u5ea7\u6a19\u7cfb.\u305f\n\u3060 \u3057,ZB\u8ef8 \u306f\u30d7 \u30ec\u30fc \u30c8\u3068\u306e\u63a5\u89e6\u9762 \u306b\u76f4\u4ea4\u3059 \u308b.\nmxB ,myB: \u03a3m\u304b \u3089\u898b \u305f \u03a3B\u306e \u4f4d\u7f6e m\u03b8B: \u03a3 m\u304b \u3089\u898b \u305f \u03a3B\u306eZm\u8ef8 \u56de \u308a\u306e\u56de\u8ee2 \u89d2\u5ea6. Rx\nB,RyB: \u03a3R\u304b \u3089\u898b\u305f \u03a3B\u306e\u4f4d \u7f6e. R\u03b8B: \u03a3R\u304b \u3089\u898b\u305f \u03a3B\u306eZR\u8ef8 \u56de \u308a\u306e\u56de\u8ee2\u89d2\u5ea6 .\nmB: \u5bfe\u8c61\u7269 \u306e\u8cea \u91cf, AB: \u5bfe\u8c61\u7269\u306e\u63a5\u89e6 \u9762\u7a4d.\ng: \u91cd\u529b\u52a0\u901f\u5ea6.\n\u672c\u7814\u7a76 \u3067\u306f,\u30d4 \u30b6\u8077\u4eba \u306e\u64cd \u4f5c\u30a2\u30ca \u30ed\u30b8\u30fc \u3092\u8e0f\u307e \u3048\u3066, \u56f33\u306b \u793a\u3059 \u3088 \u3046\u306b,\u30d0 \u30fc\u5148\u7aef \u306b\u30d7 \u30ec\u30fc \u30c8\u304c\u88c5\u7740 \u3055\u308c\u3066 \u3044 \u308b\u3082\u306e \u3068\u3059 \u308b.\u4e21 \u8005 \u306e\u63a5\u7d9a\u90e8 \u306b \u03a3m\u304c \u56fa\u5b9a \u3055\u308c,xm \u8ef8\u304a \u3088\u3073Zm\u8ef8 \u306f,\u305d \u308c\u305e\u308c\u30d0 \u30fc\u306e\u9577\u624b \u65b9\u5411,\u30d7 \u30ec\u30fc\n\u30c8\u9762\u306e\u6cd5\u7dda\u65b9\u5411 \u306b\u4e00\u81f4 \u3057\u3066\u3044 \u308b\u3082\u306e \u3068\u3059 \u308b.X,\u0398 \u306f,\n\u305d\u308c \u305e\u308c \u30d7 \u30ec\u30fc \u30c8\u306e \u4e26\u9032\u904b \u52d5\u306e\u5909 \u4f4d,\u30d7 \u30ec\u30fc \u30c8\u306e\u50be \u304d\u904b\u52d5 \u306e\u5909\u4f4d \u306b\u5bfe\u5fdc\u3059 \u308b.\u305f \u3060 \u3057,\u30d7 \u30ec\u30fc \u30c8\u304cxR-yR \u5e73\u9762 \u306b\u5bfe \u3057\u3066\u5e73\u884c(\u6c34 \u5e73)\u306a \u3068\u304d\u3092\u0398=0\u3068 \u3059 \u308b.\n(a) The xm-direction\n(b) The y7 direction\n3\u30fb2 \u554f \u984c \u8a2d \u5b9a \u30d7 \u30ec\u30fc \u30c8\u306e\u904b \u52d5 \u306b \u3088\u3063\u3066,\u30d7\n\u30ec\u30fc \u30c8\u304b \u3089\u898b \u305f\u5bfe \u8c61 \u7269 \u306e \u76f8 \u5bfe\u4f4d \u7f6e \u30fb\u59ff \u52e2 \u306b\u95a2 \u3059 \u308b\n3\u6210 \u5206(mxB,nyB,n\u03b8B)\u3092 \u64cd \u4f5c \u3057,\u4efb \u610f \u306e \u521d \u671f \u4f4d \u7f6e \u30fb\u59ff\u52e2(nxsB,mysB,m\u03b8sB)\u304b \u3089\u4efb \u610f \u306e \u76ee\u6a19 \u4f4d \u7f6e \u30fb\u59ff\u52e2\n(mxGB,myGB,m\u03b8GB)\u306b\u5230 \u9054 \u3055\u305b \u308b\u554f\u984c \u3092\u8003 \u3048\u308b.\n4. \u30d7 \u30ec\u30fc \u30c8\u52d5\u4f5c \u3068\u64cd \u308a\u6226\u7565\n\u5bfe\u8c61\u7269 \u306e\u904b\u52d5\u3092xn,ym\u5404 \u65b9\u5411\u306e\u4e26\u9032\u904b \u52d5,ZB\u8ef8 \u56de \u308a\u306e\u56de\u8ee2\u904b\u52d5 \u306e3\u3064 \u306e\u6210 \u5206\u306b\u5206 \u3051\u3066\u8003 \u3048,\u5404 \u8ef8\u65b9 \u5411\u3078\n\u306e\u904b\u52d5\u751f\u6210 \u3068\u305d\u308c \u3089\u3092\u7d44\u307f\u5408\u308f\u305b \u305f \u76ee\u6a19\u4f4d \u7f6e \u30fb\u59ff\u52e2\u3078 \u306e\u79fb\u52d5\u8a08 \u753b\u306b\u3064\u3044 \u3066\u8003 \u5bdf\u3059 \u308b.\n4\u30fb1 xm\u65b9 \u5411\u306e \u4e26\u9032\u904b \u52d5\u306e\u751f\u6210 \u5bfe\u8c61\u7269 \u304a \u3088\u3073 \u30d7\n\u30ec\u30fc \u30c8\u304c\u9759 \u6b62 \u3057\u3066\u3044 \u308b\u72b6\u614b \u3092\u8003 \u3048 \u308b(X=0,\u0398=0).\n\u56f34(a)\u306b \u793a \u3059 \u3088 \u3046\u306b,\u30d7 \u30ec\u30fc \u30c8\u306e\u4e26 \u9032\u904b\u52d5X\u306b \u3088\u3063 \u3066\u5bfe\u8c61\u7269 \u306b\u306f\u6163\u6027\u529bmBX\u304c \u4e0e \u3048 \u3089\u308c \u308b.\u6700 \u5927\u6469\u64e6\u529b\n\u03bcmBg\u3092 \u8003\u616e\u3059 \u308b\u3068,\u9759 \u6b62\u72b6\u614b\u306b \u3042\u308b\u5bfe\u8c61\u7269 \u306b\u5bfe \u3057\u3066, \u4e26\u9032\u52a0\u901f\u5ea6mxB\u3092 \u751f\u6210\u3059 \u308b\u305f\u3081 \u306b\u306f,\n(1)\n\u3092\u6e80\u8db3\u3059 \u308b\u3088 \u3046\u306a \u6587 \u3092\u4e0e \u3048\u308c \u3070 \u3088\u3044.\u3053 \u3053\u3067,X>0 \u304a \u3088\u3073X<0\u304c,\u305d \u308c\u305e\u308c\u8ca0 \u65b9\u5411(mxB<0)\u304a \u3088\u3073 \u6b63\u65b9\u5411(mxB>0)\u306e \u4e26\u9032\u904b\u52d5 \u751f\u6210\u306b\u5bfe\u5fdc\u3059 \u308b.\u306a \u304a, \u5bfe\u8c61\u7269\u304c\u30d7 \u30ec\u30fc \u30c8\u306b\u5bfe \u3057\u3066\u76f8\u5bfe\u904b\u52d5 \u3057\u3066\u3044 \u308b\u72b6\u614b\u3067\u306e xm\u65b9 \u5411\u306b\u95a2\u3059 \u308b\u904b \u52d5\u65b9\u7a0b\u5f0f \u306f,\n(2)\n\u3067\u4e0e \u3048 \u3089\u308c \u308b.\u30d7 \u30ec\u30fc \u30c8\u306e\u4e26\u9032\u904b\u52d5 \u8ef8 \u3068ym\u8ef8 \u304c\u76f4\u4ea4 \u3057\u3066\u3044 \u308b\u3053 \u3068\u304b \u3089,xm\u65b9 \u5411\u306e\u4e26\u9032\u904b\u52d5 \u306f,ym\u65b9 \u5411\u306e\nFig. 3 Model for analysis\nFig. 4 Translational manipulations", "1828 \u30d4\u30b6\u8077 \u4eba\u306e\u30cf \u30f3 \u30c9\u30ea\u30f3\u30b0\u30e1\u30ab\u30cb\u30ba\u30e0\u306b\u7740 \u76ee\u3057\u305f\u52d5\u7684\u64cd \u308a\n\u4e26\u9032\u904b\u52d5\u304a \u3088\u3073\u56de\u8ee2\u904b \u52d5 \u3068\u30ab \u30c3\u30d7 \u30ea\u30f3\u30b0\u7121 \u3057\u306b\u751f\u6210\u3059 \u308b\u3053 \u3068\u304c\u3067\u304d\u308b.\u4e00 \u65b9 \u5411\u3078 \u306e\u9023\u7d9a \u7684\u306a\u904b\u52d5\u751f\u6210 \u306b\u3064 \u3044 \u3066 \u306f,Furutani\u3089 \u306b\u3088\u3063\u3066\u63d0\u6848 \u3055\u308c \u305f\u5727 \u96fb\u30a4 \u30f3\u30d1 \u30af \u30c8 \u99c6\u52d5\u6a5f\u69cb(17)\u306e \u8003 \u3048\u3092\u7528 \u3044\u3066,\u6b63 \u65b9 \u5411 \u3068\u8ca0 \u65b9\u5411\u3067\u7570 \u306a \u308b\u30d7 \u30ec\u30fc \u30c8\u52a0 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\u306f,\u30d7 \u30ec\u30fc \u30c8\u3092 \u99c6 \u52d5\u3059 \u308b\u30a2 \u30af\u30c1\u30e5\u30a8\u30fc \u30bf\u306e\u6027\u80fd \u306b\u4f9d\u5b58\u3059 \u308b.\u4f8b \u3048\u3070, xm\u65b9 \u5411\u3078\u306e\u5bfe\u8c61\u7269 \u4e26\u9032\u904b \u52d5 \u3092\u4f8b \u306b\u6319 \u3052\u308b \u3068,\u30d7 \u30ec\u30fc\n\u30c8\u306b\u5bfe \u3057\u3066\u5fae \u5c11\u6642\u9593 \u0394t\u306e\u9593\u3060\u3051|X|=\u03bcg+\u03b5l(\u03b5l\u306f\n\u6b63\u306e\u5fae \u5c11\u5024)\u306e \u52a0\u901f\u5ea6 \u3092\u4e0e \u3048\u305f\u5834\u5408,\u3053 \u308c \u306b\u5bfe\u5fdc \u3057\u305f\n\u5bfe\u8c61\u7269 \u306e\u4e26\u9032\u5909\u4f4d \u306f\u5f0f(2)\u3088 \u308a\u0394mxB=1/2\u03b51\u0394t2\u3068\u306a \u308b.\u6975 \u7aef \u306b\u30a2 \u30af\u30c1 \u30e5\u30a8\u30fc \u30bf\u306e\u5fdc\u7b54\u6027\u80fd \u304c\u512a\u308c \u3066\u3044\u308b\u5834 \u5408(\u4efb \u610f \u306e\u52a0 \u901f\u5ea6 \u3092\u7121 \u9650\u5c0f \u6642\u9593\u751f \u6210\u3067 \u304d \u308b\u3088 \u3046\u306a\u5834 \u5408),\u7121 \u9650\u5c0f\u306e \u03b51\u3092\u7121\u9650\u5c0f\u6642 \u9593 \u0394t\u306e\u9593\u3060 \u3051\u751f\u6210\u3059 \u308b\n\u3053\u3068\u304c\u3067 \u304d\u308b\u305f \u3081,\u5bfe \u8c61\u7269 \u306e\u5909\u4f4d \u5206\u89e3\u80fd \u306f \u0394mxB\u30130\n\u3068\u3057\u3066\u7121\u9650\u5c0f\u3067\u4e0e \u3048 \u3089\u308c,\u3053 \u306e\u3088 \u3046\u306a\u5fae \u5c11\u904b\u52d5 \u3092\u7e70 \u308a\n\u8fd4\u3059 \u3053\u3068\u306b \u3088\u3063\u3066\u5bfe\u8c61\u7269 \u3092\u4efb\u610f \u76ee\u6a19\u5024 \u3078 \u3068\u53ce\u675f \u3055\u305b \u308b \u3053 \u3068\u304c\u3067 \u304d\u308b.\u305f \u3060 \u3057,\u73fe \u5b9f\u7684\u306b\u306f,\u30a2 \u30af\u30c1 \u30e5\u30a8\u30fc \u30bf \u306e\u5fdc \u7b54\u6027\u80fd \u306b\u306f \u9650\u754c\u304c \u3042\u308b\u305f\u3081,\u6709 \u9650\u6642\u9593 \u0394t\u3068\u6709\u9650\n\u5024 \u03b51\u306b\u4f9d\u5b58 \u3057\u305f\u5bfe\u8c61\u7269 \u306e\u5909\u4f4d\u5206\u89e3\u80fd \u0394mxB>0\u304c \u5b58\u5728 \u3059 \u308b.\u3057 \u305f\u304c\u3063\u3066,\u76ee \u6a19\u5024mxGB\u304c \u4e0e \u3048 \u3089\u308c \u305f\u5834\u5408,\u76ee \u6a19\u5024 \u8fd1\u508d(mxGB-1/2\u0394mxB)\u3013mxB\u3013(mxGB+1/2\u0394mxB)\u306e \u7bc4 \u56f2\u3067\u306e\u53ce\u675f\u304c\u4fdd\u8a3c \u3055\u308c \u308b\u3053 \u3068\u306b\u306a\u308b.\u4ee5 \u4e0a\u306e\u8b70 \u8ad6 \u3092\u8e0f \u307e\u3048 \u308b\u3068,\u5f0f(1),(3),(14)\u306b \u57fa \u3065\u304d,\n(15)\n(16)\n(17)\n\u3092\u4e0e \u3048\u308b\u3053 \u3068\u306b \u3088 \u308a,\u305d \u308c \u305e\u308c\u5bfe\u8c61 \u7269\u306eXm\u65b9 \u5411\u306e\u4e26 \u9032,ym\u65b9 \u5411\u306e\u4e26\u9032,\u56de \u8ee2 \u306b\u3064 \u3044\u3066\u5fae\u5c11\u904b \u52d5 \u3092\u751f\u6210 \u3059 \u308c \u3070,\u5404 \u8ef8 \u306b\u3064\u3044\u3066\u305d\u308c\u305e\u308c \u76ee\u6a19\u5024\u8fd1\u508d \u3078\u53ce\u675f \u3055\u305b \u308b \u3053 \u3068\u304c \u3067\u304d \u308b.\u305f \u3060 \u3057,\u03b51\uff5e \u03b54\u306f\u6b63 \u306e\u5fae \u5c11\u5024 \u3067\u3042 \u308b. \u3053\u3053\u3067,\u6ce8 \u610f \u3057\u305f\u3044\u306e \u306f,\u5bfe \u8c61\u7269 \u306e\u56de\u8ee2\u904b \u52d5\u306f\u4e21\u4e26\u9032 \u904b\u52d5 \u3068\u30ab \u30be\u30d7 \u30ea\u30f3\u30b0 \u3057,\u4e00 \u822c\u7684 \u306b\u56de\u8ee2\u5f8c \u306e\u5bfe\u8c61\u7269\u4f4d\u7f6e \u304c\u56de\u8ee2\u524d\u306e\u4f4d \u7f6e \u3068\u4e00\u81f4\u3059 \u308b\u3053 \u3068\u304c\u4fdd \u8a3c \u3055\u308c\u306a\u3044\u306e \u306b\u5bfe\n\u3057\u3066,\u4e26 \u9032\u904b\u52d5xm\u304a \u3088\u3073ym\u306f \u5404\u8ef8\u72ec\u7acb \u3057\u3066\u751f\u6210 \u3067\u304d \u308b\u70b9\u3067 \u3042\u308b.\u3053 \u306e\u70b9 \u3092\u8e0f \u307e\u3048,\u56f36\u306b \u793a\u3059 \u3088 \u3046\u306b,\u306f \u3058\u3081\u306b\u56de\u8ee2\u904b\u52d5 \u306b\u3088 \u308a\u5bfe\u8c61\u7269 \u3092 \u76ee\u6a19 \u59ff\u52e2\u307e\u3067\u56de\u8ee2 \u3055\u305b (\u56f36(a)),\u7d9a \u3044\u3066xm(ym)\u65b9 \u5411\u306e\u4e26\u9032\u79fb\u52d5,ym(xm) \u65b9 \u5411\u306e\u4e26\u9032\u79fb\u52d5 \u3092\u9806 \u306b\u884c \u3048\u3070(\u56f36(b),(c)),\u4efb \u610f\u306e \u521d \u671f\u4f4d\u7f6e \u30fb\u59ff\u52e2(mxSB,mySB,m\u03b8SB)\u304b \u3089\u4efb \u610f\u306e \u76ee\u6a19\u4f4d \u7f6e \u30fb \u59ff\u52e2(mxGB,myGB,m\u03b8GB)\u3078 \u3068\u5230 \u9054\u3059 \u308b\u3053 \u3068\u304c\u4fdd \u8a3c\u3067 \u304d\u308b. \u3059 \u306a\u308f \u3061,\u5f0f(15)\uff5e(17)\u306e \u5fae\u5c11 \u904b\u52d5\u3092\u65ad\u7d9a\u7684 \u306b\u7e70\n\u308a\u8fd4 \u3059 \u3053\u3068\u306b\u3088\u3063\u3066,\u5916 \u4e71 \u304c\u5165 \u3089\u306a\u3051\u308c\u3070\u5bfe\u8c61\u7269\u306e\u4f4d\n\u7f6e \u304a \u3088\u3073\u59ff\u52e2 \u3092 \u76ee\u6a19\u5024\u8fd1\u508d \u3078 \u3068\u53ce\u675f \u3055\u305b \u308b\u3053\u3068\u304c\u53ef\u80fd \u3067\u3042\u308b.\u4ee5 \u4e0a\u306e \u3088 \u3046\u306a\u7c21\u4fbf \u306a\u64cd \u308a\u8a08\u753b \u306f,\u63a1 \u7528 \u3057\u305f\u30d7\n\u30ec\u30fc \u30c8\u306e2\u81ea \u7531\u5ea6\u69cb\u6210\u304c,\u5bfe \u8c61\u7269 \u306e3\u904b \u52d5 \u81ea\u7531\u5ea6 \u4e2d,2\n\u81ea\u7531\u5ea6 \u3092\u72ec\u7acb \u3057\u3066\u751f\u6210 \u3067\u304d\u308b \u3068\u3044 \u3046\u30ab \u30c3\u30d7 \u30ea\u30f3\u30b0\u6027 \u306e \u5c11\u306a \u3055\u306b\u8d77\u56e0 \u3057\u3066\u3044\u308b.\u3053 \u306e \u3088 \u3046\u306a\u89b3\u70b9 \u306b\u304a \u3044\u3066,\u63a1\nset of goal position and orientation" ] }, { "image_filename": "designv11_61_0003581_avss.2013.6636627-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003581_avss.2013.6636627-Figure4-1.png", "caption": "Figure 4: Comparison of energy consumption of three different approaches on Scenario 4: (a) Traditional approach-no energy consideration in task assignment; (b) proposed method employing only the critical energy level Ec; (c) proposed method employing both Ec and energy consumption rate Er .", "texts": [ " In this particular scenario, when only Ec is employed and camera 1 drops to that level, it was able to delegate three tracking jobs to the other cameras. Then, when one of the objects enters a region that is visible only by camera 1, it needs to take the tracking job, and dies soon after. Thus, it is very important to increase the time it takes to drop to a critical energy level, by considering the energy consumption rate. We performed energy comparison of the three different approaches described above, for Scenario 4 shown in Fig. 2d as well. The results are presented in Fig. 4. The network lifetimes for Fig. 4a, Fig. 4b and Fig. 4c are 2150, 2820 and 3268 simulation steps, respectively. The proposed approach provides 52% increase in the network lifetime as compared to not considering energy in the task assignment. We compared the communication cost of our proposed method with that of an active broadcast method. In our method, messages are exchanged in an event-driven manner when the nodes initiate the task reassignment. Whereas, in active broadcast method, communication happens at periodic, pre-determined intervals. As seen in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003616_med.2011.5983117-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003616_med.2011.5983117-Figure1-1.png", "caption": "Fig. 1. Sketch of the three-phase radial bearing", "texts": [ " In this article is presented an approach to estimate parameters of the radial bearing (the axial bearing case will be considered in future work). The parameters that are to be estimated here depend on the geometry and the material of the bearing. The estimation of these parameters is important because such parameters are difficult to calculate and may slightly vary over time. The rotor is levitating using a three-phase electromagnetic radial bearing, arranged like three coupled \u201dhorseshoe magnets\u201d around the rotor (Fig. 1). The three generated current provide three independent control inputs. The mathematical model of the bearing is based on the assumption of a rigid body and leads to decoupled equations for forces. The dynamics equations, under simplifying assumptions, are written as follows: mY\u0308 = Fy (1) mZ\u0308 = Fz (2) where Y and Z represent the coordinates of the center of mass of the rotor in a Cartesian frame (with axes y et z) which is fixed in the space, at a point being considered as the center of the device. The forces Fy and Fz represent the resulting forces applied in directions y et z, respectively. The rotor has a mass m. The resultant forces in the plan (y \u2212 z) are given by the superposition of the forces generated by the 978-1-4577-0123-8/11/$26.00 \u00a92011 IEEE 388 magnets: ( Fy Fz ) = ( sin \u03b11 sin \u03b12 sin \u03b13 cos\u03b11 cos\u03b12 cos\u03b13 ) F1 F2 F3 (3) The angles that appear in (3) are presented Fig. 1. Individually, the magnetic forces can be modeled by (k \u2208 {1, 2, 3}) Fk = \u03bbk i2k ( s \u2212 ( sin\u03b1k cos\u03b1k )T ( Yb Zb ) )2 (4) where Yb and Zb are the positions in the bearing plan. s is the nominal air gap and \u03bbk are parameters depending on the geometry and the materials of the bearing. They will be estimated on-line. Calculating control currents: The reference currents are obtained from the desired forces. In order to simplify the notations, we assume that the bearing is symmetric, i.e. \u03b11 = \u03c0, \u03b12 = \u2212\u03c0 3 and \u03b13 = \u03c0 3 " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001050_3-540-29461-9_99-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001050_3-540-29461-9_99-Figure3-1.png", "caption": "Fig. 3. The open loop system considered", "texts": [ " This leakage can cause the internal negative pressure to rise up and in this situation the robot could fall down. On the other side if the internal pressure is too low (high \u2206p), a very big normal force is applied to the system. As a consequence, the friction can increase in such a way to not allow robot movements. This problem can be solved by introducing a control loop to regulate the pressure inside the chamber to a suitable value to sustain the system. The considered open loop system and the most easily accessible system variables has been schematized in Fig. 3; in this scheme the first block includes the electrical and the mechanical subsystem and the second block includes the pneumatic subsystem. The used variables are the Motor voltage reference (the input signal that fixes the motor power) and the Vacuum level (the negative pressure inside the chamber). 1008 D. Longo and G. Muscato Since it is very difficult to have a reliable analytical model of that system, because of the big number of parameters involved, it has been decided to identify a black box dynamic model of the system by using input/output measurements" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002236_2007-01-3740-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002236_2007-01-3740-Figure6-1.png", "caption": "Figure 6: Single Cavity, Single Regime IVT", "texts": [ " Therefore, the proven two roller full toroidal Variator design from the OPE sector has been incorporated into a single regime IVT and the resulting transmission has been applied to a typical \u201centry level\u201d vehicle application. Use of the two roller, single cavity Variator in a single regime, geared neutral IVT with a planetary gearset results in an ultra low-cost transmission in a compact package, with no requirement for a launch clutch or torque converter nor a separate reversing device. As the system is an IVT, forward drive, neutral and reverse are achieved via the Variator ratio spread and the mixing planetary gearset. The arrangement is detailed in Figure 6 and comprises a 90mm roller diameter Variator in a package length of 340 mm. The transmission is designed for a typical \u201centry level\u201d vehicle with circa 30 hp / 22 kW and 45 Nm in a vehicle with a gross vehicle weight (GVW) of circa 850 kg. Performance comparison of this concept to a 4 speed manual transmission in the target vehicle with identical vehicle weights and over the fuel economy test cycle provides the following results:- Hence, due to the ability of the variable drive transmission to optimise engine operation, the IVT delivers improved vehicle performance versus a manual transmission without compromising fuel economy" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003752_physreve.81.061902-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003752_physreve.81.061902-Figure7-1.png", "caption": "FIG. 7. Color online A rod-shaped cell differentiates into a round myxospore.", "texts": [ " These cells form clockwise and counterclockwise cell streams in approximately equal proportions and they no longer reverse. In contrast, the inner domain consists of less ordered nonmotile cells at threefold lower cell density. C-signaling continues while cells in the high-density outer domain are moving in circular orbit within the annulus. The level of C-signal molecules on each cell surface increases until it reaches a final threshold of Ns=450 for sporulation. During sporulation, the cell differentiates into spore by shortening and rounding up its rodlike cell body. Figure 7 shows a schematic diagram of the cell differentiation. Myxospores are unable to move on their own. The cells in the outer domain that differentiate into spores can only be transported passively to the inner domain by undifferentiated motile cells 41 . In our simulation, we model this passive transport as a result of mechanical collisions between the spores and the motile cells, which occurs very likely due to the high density of the outer domain. The details on spore differentiation and spore transport are described in Sec", " If the difference in angles between these orientations is close to 180\u00b0 head-on collision , then cell k rotates its orientation by some small random angle away from cell j and take the rotated orientation as the new direction Vk. Otherwise, collision is resolved in a way defined in head-to-body collision. Spore transport occurs during the final stage of myxobacteria life cycle. Inside a fruiting body, rod-shaped cells that have accumulated enough C-signal molecules differentiate to form a round nonmotile spore. As illustrated in Fig. 7, the head node of the cell becomes the center of the spore and the radius r of the spore k is defined to be half of the length of the segment connecting two consecutive nodes in the initial cell, that is, r= 1 2 nk1 \u2212nk2 , where nk1 and nk2 denote the position of the first head and second nodes of cell k, respectively. Spore differentiation does not occur simultaneously during the fruiting body development, and thus mechanical collision between a spore and a nondifferentiated motile cell may occur inside the fruiting body" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000368_1-4020-3796-1_7-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000368_1-4020-3796-1_7-Figure1-1.png", "caption": "Figure 1. Schematic representation of a multibody system.", "texts": [ " The crash event can be preceded by warning signs, take a longer time than for road vehicles and consist on multiple impacts. For these reasons special attention is paid to the need for these models to include muscle actions. In order to identify the typical postures of railway passengers and the reflexive and controlled muscle actions and joints stiffening several numerical and experimental procedures, based on gait analysis methodologies, are also proposed here [12, 13]. A multibody system is a collection of rigid bodies joined together by kinematic joints and force elements as depicted in Figure 1. For the ith body in the system, qi denotes a vector of coordinates [14]. A vector of velocities for a rigid body i is defined as vi. The vector of accelerations for the body, denoted by iv , is the time derivative of vi. For a multibody system containing nb bodies, the vectors of coordinates, velocities, and accelerations are q, v and v that contain the elements of qi, vi and iv , for i=1, ..., nb. The kinematic joints between rigid bodies are described by mr independent constraints: q 0 (1) The time derivatives of the constraints yield the velocity and acceleration equations" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003960_14644193jmbd255-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003960_14644193jmbd255-Figure1-1.png", "caption": "Fig. 1 General tether model for \u2018N \u2019 number of rigid bodies", "texts": [ " The proposed recursive rigid-body tether formulation results in computations on the same order as the three DOFs lumped mass models with an additional state for a viscoelastic element. Furthermore, the elimination of high stiffness springs allows larger integration time steps, further improving computation speeds. The result is a computationally efficient model that can accurately represent a low-strain tether used in many engineering applications without the need to add stiff elastic elements. The tether is divided into a chain of N bodies connected by spherical joints with each link being a body of revolution. Figure 1 shows the tether attached to the ground with the jth body, bj , having two connections, joints cj\u22121 and cj , and an external load applied to the N th body. The N th body, bN , is the terminal link, body b1 is the root link, and b0 is a fixed body or ground where connection c0 is stationary. A body, bj , is attached to its parent, bjp, in the direction of the ground where the subscript jp represents the parent of j. Body b0 is attached to a fixed or inertial frame (I ) defined by three orthogonal unit vectors, iI , jI , and kI ", " In order to avoid a singularity in the rotation kinematics, the orientation can alternatively be defined by the four quaternion parameters q0j , q1j , q2j , and q3j [17] resulting in the transformation from the inertial frame, I , to the j frame given by T j I = \u23a1 \u23a3 2q2 0j \u2212 1 + 2q2 1j 2q1jq2j + 2q0jq3j 2q1jq2j \u2212 2q0jq3j 2q2 0j \u2212 1 + 2q2 2j 2q1jq3j + 2q0jq2j 2q2jq3j \u2212 2q0jq1j 2q1jq3j \u2212 2q0jq2j 2q2jq3j + 2q0jq1j 2q2 0j \u2212 1 + 2q2 3j \u23a4 \u23a6 (1) where q0j = cos ( \u03c8j 2 ) cos ( \u03b8j 2 ) cos ( \u03c6j 2 ) + sin ( \u03c8j 2 ) sin ( \u03b8j 2 ) sin ( \u03c6j 2 ) q1j = cos ( \u03c8j 2 ) cos ( \u03b8j 2 ) sin ( \u03c6j 2 ) \u2212 sin ( \u03c8j 2 ) sin ( \u03b8j 2 ) cos ( \u03c6j 2 ) q2j = cos ( \u03c8j 2 ) sin ( \u03b8j 2 ) cos ( \u03c6j 2 ) + sin ( \u03c8j 2 ) cos ( \u03b8j 2 ) sin ( \u03c6j 2 ) q3j = sin ( \u03c8j 2 ) cos ( \u03b8j 2 ) cos ( \u03c6j 2 ) \u2212 cos ( \u03c8j 2 ) sin ( \u03b8j 2 ) sin ( \u03c6j 2 ) (2) A transformation from the j \u2212 1 frame to the j frame can be formed using equation (1) and is given as T j j\u22121 = ( TI j )T TI j\u22121 (3) Position vectors from the j \u2212 1 connection to the jth body mass centre are conveniently expressed in the bj frame as rm j = xmjij . Similarly, the vector from connection j \u2212 1 to connection j, also expressed in the bj frame, is defined as rc j = xcjij . Both vectors rm j and rc j have only an ij component as a result of each body\u2019s symmetry. The tether configuration in Fig. 1 has spherical joints connecting the N bodies with no applied twisting torque at the ground or terminal link. In addition, the bodies are slender such that the moment of inertia Ixx will be small compared to the other moments of inertia. This combination results in spinning dynamics of each body having a minimal affect of the tether\u2019s overall motion. Elimination of tether spin will later aid in efficient computation of recursive dynamics. The angular velocity of jth body with respect to the inertial frame of reference is then defined as \u03c9j/I = qjjj + rjkj (4) where the spin rate, pj , is zero" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000278_1-84628-179-2_7-Figure7.5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000278_1-84628-179-2_7-Figure7.5-1.png", "caption": "Fig. 7.5. T-Wing vehicle configuration.", "texts": [ " Collective blade pitch control is still required to marry efficient high-speed horizontal flight performance with the production of adequate thrust on take-off, however, even this complication can be deleted, with little performance penalty, if high dash speeds are not required. \u2022 The current vehicle uses a canard to allow a more advantageous placement of the vehicle centre of gravity (CG). \u2022 Two separate engines are used in the current design though the possibility of using a single engine with appropriate drive trains could also be accommodated. The basic configuration of the T-Wing is presented in Figure 7.3 and Figure 7.4, with a diagram in Figure 7.5 showing some of the important gross geometric properties. The aircraft is essentially a tandem wing configuration with twin tractor propellers mounted on the aft main wing. The current T-Wing vehicle is a technology demonstrator and not a prototype production one. The aims of the T-Wing vehicle programme are to prove the critical technologies required of a tail-sitter vehicle before committing funds to full-scale development. The most important aspects of the T-Wing design that have to be demonstrated are reliable autonomous hover control and the ability to perform the transition manoeuvres between horizontal and vertical flight" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002665_09596518jsce550-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002665_09596518jsce550-Figure1-1.png", "caption": "Fig. 1 Simplified model of the toy helicopter", "texts": [ " 3 MODELLING OF THE HELICOPTER The main differences between toy helicopters and full-size helicopters include the following [10]: (a) a much higher ratio of the main rotor mass to the fuselage mass; (b) the rotation speed of the main rotor of toy helicopters is higher than that of most full-sized helicopters; JSCE550 F IMechE 2008 Proc. IMechE Vol. 222 Part I: J. Systems and Control Engineering at Harvard Libraries on June 28, 2015pii.sagepub.comDownloaded from (c) toy helicopters have very stiff main rotors without flapping hinges. The initial effects of the main rotor make a significant contribution to the rotational dynamics and can not be neglected. Not only the fuselage but also the main rotor should be modelled as rigid body. The simplified model of a toy helicopter is shown in Fig. 1, where the fuselage is labelled with an F, the main rotor with MR, the tail rotor with TR, the mass centre of the fuselage with Fo, the mass centre of the main rotor with MRo, the mass centre of the tail rotor with TRo and the variables are as listed in the Appendix. The kinematical equations of translation and rotation can be deduced as follows. The kinematical equations of translation are _px~u _py~v _pz~w \u00f011\u00de The kinematical equations of rotation _w~pztan h\u00f0 \u00de sin w\u00f0 \u00deqzcos w\u00f0 \u00der\u00f0 \u00de \u00f012a\u00de _h~cos w\u00f0 \u00deq{sin w\u00f0 \u00der \u00f012b\u00de _y~ sin w\u00f0 \u00deqzcos w\u00f0 \u00der\u00f0 \u00de=cos h\u00f0 \u00de \u00f012c\u00de As for the dynamical equations, the helicopter is regarded as two rigid bodies, the fuselage and the main rotor", " 222 Part I: J. Systems and Control Engineering at Harvard Libraries on June 28, 2015pii.sagepub.comDownloaded from FMR 3 lifting force generated by the main rotor Fr generalized active force F r generalized inertial force IF 11, IF 22, IF 33 principal moments of the inertia of the fuselage around the f1, f2, f3 axes respectively IMR 11 , IMR 22 , IMR 33 principal moments of inertia of main rotor around the f1, f2, f3 axes respectively LF offset of the point OF from the point Fo in the f3 direction in Fig. 1 LMR offset of the point OF from the point MRo in the f3 direction in Fig. 1 LT offset of the point OF from the point TRo in f1 direction in Fig. 1 mF mass centre of the fuselage mMR mass centre of the main rotor M mass centre of helicopter n1, n2, n3 unit vectors p, q, r rotation speeds with respect to body frame around the f1, f2, and f3 axes respectively px, py, pz coordinates of the helicopter in the inertial frame in f1, f2, and f3 directions respectively T MR 1 torques generated by main rotor in f1 axis T MR 2 torques generated by main rotor in f2 axis T MR 3 torques generated by main rotor in f3 axis T TR 2 drag torques generated by tail rotor u, v, w translational speeds of the helicopter in the inertial frame in f1, f2, and f3 directions respectively w, h, y roll angle, pitch angle, and yaw angle respectively vMR rotation speed of main rotor Proc" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003392_9780470876541.ch5-Figure5.21-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003392_9780470876541.ch5-Figure5.21-1.png", "caption": "Figure 5.21 Voltage constraint ellipse, current constraint circle, and constant torque locus.", "texts": [ " vesLsi e* qs 2 \u00fe veLsi e* ds 2 V2 smax \u00f05:84\u00de Under the assumption of the slow enough variation of the rotor flux linkage and the precise vector control, the current maximizing the torque of the induction machine, which is represented as (5.85), can be derived from the constraints given by (5.52) and (5.84). Like the case of PMSM drive system, the size of the ellipse by the voltage constraint decreases as the operating speed increases as shown in Fig. 5.19. Te \u00bc 3 2 P 2 Lm 2 Lr ie*ds i e* qs \u00f05:85\u00de The current constraint in (5.52) can be depicted as a circle in the current plane as shown in Fig. 5.20. And the possible operating region is the shaded area in Fig. 5.21, which is the cross section of the ellipse and the circle. And the torque is depicted as a reciprocal proportion curve in the current plane as shown in Fig. 5.21. 5.4.4.3 Constant Torque Region (ve vb) If d-axis current for the maximum torque, which is the current at the crossing point of the ellipse and the circle, is larger than the rated value of d-axis current of the induction machine, then the d-axis current reference should be set as the rated value to prevent the magnetic saturation of the induction machine. That is the case of point A in Fig. 5.22, where the torque, Te1 may be obtained by ieds1, but it is larger than i e ds rate and it would result in the severe saturation of the magnetic circuit of the machine" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000052_cdc.1999.832740-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000052_cdc.1999.832740-Figure4-1.png", "caption": "Figure 4: Contact line (fixed e).", "texts": [ " The corresponding contact surface in C is defined by: c c = i(Z,Y,8)1(5q(5,Y,e),Yq(z,Y,e)) E w) 4The case of type A paths does not require the application of PMP and has been solved geometrically 1221. We do not treat it in the paper. with (xq(x, y, e), yq(x, y, e)) the coordinates of q in the plane as functions of the robot configuration (z, y, 0 ) . Given the equation y = m x + n of w, w.r.t. a Cartesian reference frame, call 4 (-T 5 4 5 T) the angle between the line supporting the segment Pq and the orientation vector v of the car (see Fig. 4), 1 the length of Pq. For a given orientation 0 of the car, the line CO followed by P while q \u201c m ~ v e s \u201d ~ along w (see Fig. 4) has the equation Y = mz + n + Zmcos(8 + 4) - Isin(8 + 4) (7) which describes a ruled surface in C, as O varies in the interval [-T, TI. To have q in contact with w, the robot final state is constrained to belong to the set Sf = C, defined by Yf - mzf - n - Zm cos(Of + 4) + 1 sin(Of + 4) = 0. (8) Write (8) in the form x((f) = 0, with Cf = (zf, gf, Of), and define M = - = (-m, 1, zmsin(O,+4)+zcos(Of+4)). (9) $f = MTC. (10) Admissible solutions must satisfy the transversality condition: any) and all the inflection points of P must lie on a line perpendicular to w" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003357_978-1-4419-9792-0_37-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003357_978-1-4419-9792-0_37-Figure2-1.png", "caption": "Figure 2 \u2013 Geometry of SLM rectangular specimens (127 mm x 35 mm) with holes location (1,2,3) and disks position (diameter 35 mm)", "texts": [ " In order to reduce deformations that may lead to failures in the building process, specimens are manufactured onto a 15 mm thick substrate (building plate). Besides, parts are built using supports of 4 mm height in order to facilitate removal from the building platform. It has been observed that supports with a spare dimension of 2 mm cause the disengagement of the specimen from the building plate due to high thermal stress (Fig. 1). A spare dimension of 1 mm is recommended in order to avoid distortions of SLM parts. Specimen geometry is reported in figure 2. The rectangular specimen is oblique on the building plates in order to facilitate the powder deposition by means of a knife. Positions 1, 2 and 3 indicate the locations of residual stress measurements. These locations are chosen in order to have a correspondence between disks studied in [7] and rectangular specimens made with identical material and process parameters. This should point out the influence of geometry on residual stresses generated by SLM. The hole drilling method is utilized for the residual stresses measurements", " These facts should allow to maximize the adhesion between layers thus ensuring a nearly full density which limits the magnitude of residual stresses [13]. Specimens of circular geometry (disks of 35 mm diameter) studied in [7] exhibited very small warping effect which are instead very pronounced in specimens where one dimension predominates over the others (Fig. 1). The aim of the present work is to investigate on the influence of geometry (that is rectangular, with one dimension much larger than others, or circular) on residual stress values measured in the same location. As it can be observed in Fig. 2, the SLM rectangular specimens have been built so that SLM disks are comprised into its edge. Three SLM rectangular specimens have been studied. In all cases residual stresses measurements have been executed by HDM at location 1, 2 and 3. Fig. 3 shows the drilling device utilized. Strain values have been measured with System 5000 by Micro Measurements. (RS 200 Milling Guide by Micro Measurement) Released strains have been measured and strain versus hole depth have been obtained for each measurements" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003445_9781118361146.ch7-Figure7.35-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003445_9781118361146.ch7-Figure7.35-1.png", "caption": "Figure 7.35 Diagram of \u2018inside-out\u2019 electric motor", "texts": [ " Here the electric machine, which can work as either a motor or generator, is mounted directly in line with the engine crankcase. Such machines are in most cases a type of BLDC (or synchronous AC) motor as described in Section 7.3.2. They will be multiple-pole machines, since their location means their dimensions need to be short in length and wide in diameter. They nearly always are different from the machine of Figure 7.25 in one important respect: they are usually \u2018turned inside out\u2019, with the stationary coils being on the inside, and the rotor being a band of magnets moving outside the coil. The idea is shown in Figure 7.35. The larger diameter permits this construction, which has the advantage that the centrifugal force on the magnets tends to make them stay in place, rather than throw them out of their mounting. It is worth pointing out that this same type of \u2018inside-out\u2019 motor is used in motors that are integral with wheels, such as the machine of Figure 9.9. However, not all parallel hybrids use special multiple motors of this type. Some hybrid vehicles use a fairly conventional, single-pole, fairly high-speed machine, which is connected to the engine crankshaft much like the alternator in a conventional IC engine vehicle" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000975_have.2004.1391885-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000975_have.2004.1391885-Figure2-1.png", "caption": "Figure 2", "texts": [ "6 GhzP4 Pentium computers running on Windows 2000. The ' www.SensAble.com I-' 0-7803-88 17-8/04/$20.00 02004 JEEE computers at each site are connected together via Ethemet. The UDP protocol is used. Network delays between the two haptic systems were simulated to have a one way time delay of 65ms to 200m (random uniform distribution) with approx 30% packet loss. Both the remote and local site haptic devices interact with a virtual environment with haptic effects simulating a bifurcating artery. A screen shot of the experiment is shown in Figure 2. By manipulating the haptic devices, each user can feel the outside and inside of each tube. When a user hits the bifurcation, then there are two possible directions that the user can go. The goal is to have an instructor at the local site guide a student at a remote site through the bifurcation so that the decision as to which tube to proceed is done entirely by having the local instructor \"push on the hand of the remote student. In Section 2 of this paper, the time delay compensation scheme is briefly described" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003465_978-3-642-15621-2_43-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003465_978-3-642-15621-2_43-Figure1-1.png", "caption": "Fig. 1. The structure of two-leaf semi-rotary VAWT", "texts": [], "surrounding_texts": [ "For any control system, the mathematical model is the most fundamental basis of its control. For wind turbines, the uncertainty of aerodynamics and complexity of power electronics make the dynamic model difficult to establish [8]. The technology of HAWT is matured and its model establishing method is quite fixed. While VAWTs are diverse, so there is no uniform method for them. And the variety of VAWT makes modeling quite difficult. In addition, there are less references and theory, which produces more uncertainty in modeling. Therefore, the model about VAWT is quite difficult in establishing. For the purpose of establishing an accurate model, it is necessary to know the structure of two-leaf semi-rotary VAWT. The structural diagram is as shown in Fig .1: From Fig .1, we know that this VAWT belongs to lifting-dragging mixed type wind turbine. Due to the dominating dragging force, we still consider it as resistance-type wind turbine. In order to simplify the structure of wind turbine in this paper, we can just analyze one blade working situation. And the movement trace of single blade is as shown in Fig .2: Fig. 2 is the top view of single blade operation, where O is wind turbine spindle, OA = R, p is any point in blade. Axis-l is established along the blade direction which is H to A and A is the origin of axis-l. According to semi-rotary characteristic, the line where the blade leaves and the circle intersect at point H which is a fixed point and, the angular velocity of blade revolving around point O is twice as much as that revolving around point H [9]. If blade width is a and the angular velocity of blade revolving O is \u03c9, then the speed of point p at the normal direction of blade can be expressed as: vn = \u03c9/2 \u00d7 HP = (2R cos \u03b8 + l)\u03c9/2 . (1) The wind speed at the normal and tangential direction of blade can be respectively expressed as: v\u22a5 = v cos \u03b8, vr = v sin \u03b8 . Thus, the relative wind speed at the normal direction of blade is: vrel = v\u22a5 \u2212 vn = v cos \u03b8 \u2212 (2R cos \u03b8 + l)\u03c9/2 . (2) According to the aerodynamics, the aerodynamic acting on object surface can be expressed as: F = \u03c1Sv2 rel/2 . (3) Where, \u03c1 is air density (usually is 1.25kg/m3), S is the area that wind acts on object, vrel is relative wind speed. According to (3), the torque that the wind turbine gained at point p is: dTdr = \u03c1hv2 rel(R cos \u03b8 + l)dl/2 . (4) Where, the torque arm is equal to (R cos \u03b8 + l) and dl is the micro-element at axis-l division. vr leaves along the blade tangent direction so that there is no area that wind acts on object, thus it contributes nothing to blade. The movement of blade is symmetrical on axis-x, so according to (4), the average torque gained by a whole blade can be expressed as: Tdr = 2 \u03c0 \u222b \u03c0 2 0 \u222b a 2 \u2212 a 2 [ 1 2 \u03c1hv2 rel(R cos \u03b8 + l) dl ] d\u03b8 = 2 \u03c0 \u222b \u03c0 2 0 1 2 \u03c1h[Ra(v \u2212 R\u03c9)2 cos2 \u03b8 + a3 12 ( 5R\u03c92 4 \u2212 v\u03c9)] cos \u03b8d\u03b8 = 2\u03c1hRa 3\u03c0 (v \u2212 R\u03c9)2 + \u03c1ha3 12\u03c0 ( 5R\u03c92 4 \u2212 v\u03c9) . Because the value of the latter is very small, it can be ignored. The expression above can be expressed as: Tdr = 2\u03c1ha(v \u2212 R\u03c9)2/3\u03c0 . (5) In this paper, for the two-leaf semi-rotary VAWT, the total torque gained by wind turbine can be regarded as 2Tdr because of the little influence caused by two blades mutual covering. Assumed that the resisting torque of wind turbine is Tf , and then the wind turbine dynamic equation can be expressed as: 2Tdr \u2212 Tf = Jd\u03c9/dt . If v0 is rated wind speed and \u03c90 is rated angular velocity, the following equation will be gained: 2Tdr(v0, \u03c90) \u2212 Tf = 0 . (6) Only steady revolution, the wind turbine would have higher quality power output. So the constant revolution speed control must be guaranteed. When the wind speed is higher than rated wind speed, a step motor can be used to keep wind turbine working at rated revolution speed. The process of adjustment is as shown in Fig. 3 and the wind turbine control system block diagram is as shown in Fig. 4: In Fig. 3, \u03b1 is defined as blade power angle and it is adjusted through a step motor in Fig. 4. When the power angle is equal to \u03b1, the fixed point H will move to point H \u2032. At this time, the torque of wind turbine is no longer symmetrical about the horizontal axis. For convenient analysis and calculation, a new coordinate system will be gained through counterclockwise rotating original coordinate system by 2\u03b1 degrees. As shown in Fig. 5: In the coordinate system x\u2032oy\u2032, the wind speed can be divided into two parts: v1 and v2. Their value can be expressed as: v1 = v cos(2\u03b1), v2 = v sin(2\u03b1) . According to Fig. 4 and (5), under the effect of v1, the average driving torque gained by wind turbine can be expressed as: Tv1 = 2\u03c1ha(v1 \u2212 R\u03c9)2/3\u03c0 . When v2 acts on the wind turbine alone, the wind plays a negative role on wind turbine while the blade leaves above axis-x\u2032; and it plays a positive role on wind turbine while the blade leaves below axis-x\u2032. It is easy to know that v2 does not work on wind turbine in a period for the symmetry of action. Therefore, the average driving torque gained by wind turbine after power angle adjustment is: T\u03b1 = Tv1 = 2\u03c1ha[v cos (\u03b1) \u2212 R\u03c9]2/3\u03c0 . (7) From (6) and (7), the kinematic equation of wind turbine after power angle adjustment can be expressed as: q2 \u2212 P 2 = \u2212(3\u03c0J/4\u03c1hRa) \u00b7 dq/dt . (8) Where, q = v cos(2\u03b1), p = v0 \u2212 R\u03c9. From (8), we can gain a discrete equation as shown below: q(k) = 2pq[q(k \u2212 1) \u2212 p emTs [q(k \u2212 1) + p] \u2212 [q(k \u2212 1) \u2212 p] + p , \u03c9(k) = (v cos(2\u03b1) \u2212 q(k))/R . (9) Where, m = 3\u03c1hRap/3\u03c0J , Ts is sampling period and \u03c9(k) is the angular speed of wind turbine at k\u00b7Ts." ] }, { "image_filename": "designv11_61_0002163_j.mechmachtheory.2007.03.005-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002163_j.mechmachtheory.2007.03.005-Figure1-1.png", "caption": "Fig. 1. Helicoidal vector field and pencils of lines on nullplanes.", "texts": [ " (3) is also called the reciprocal condition. There are three known LLCs associated with kinematics: (1) the LLC associated with the instantaneous motion of a body, (2) the LLC associated with the homologous points of a finite displacement, and (3) the LLC associated with the homologous lines of a finite displacement. A pair of homologous points (lines) contains a point (line) and its corresponding point (line) after a finite displacement. The first two LLCs have been shown to be composed of pencils of lines on nullplanes. As shown in Fig. 1, given a screw with pitch p, one can draw12 equal-pitched helices, whose tangent vectors at all points in space form a helicoidal vector field [1,9]. The plane normal to the tangent vector at a point is called a nullplane [3]. The LLC corresponding to the screw consists of all the pencils of lines on nullplanes. In instantaneous kinematics, the motion of a rigid body can be modeled as the combination of a rotation about a screw axis with angular velocity x and a translation along the same screw axis with velocity v", " In finite kinematics, a displacement can be modeled as the rotation of a rigid body about a screw axis for an angle / followed by the translation of the body along the same screw axis with distance d. The vector connecting every pair of homologous points in space belongs to a helicoidal vector field of pitch \u00f0d=2\u00de= tan\u00f0/=2\u00de [4], and its corresponding LLC is usually referred to as the bisecting LLC. It has been shown that the bisecting LLC degenerates into the LLC of instantaneous kinematics; nevertheless, the LLC associated with homologous lines has been found to be different from the other two LLCs [3], and it does not take the form of Fig. 1. The LLC of homologous lines was introduced by Bottema and Roth [3] when investigating the two-position theory in kinematics. As shown in Fig. 2, given two-positions of a rigid body, one can determine the screw axis of the displacement and the corresponding rotation and translation parameters, / and d, respectively. For any point R in space, there exists a pair of homologous lines, L1 and L2, intersecting at R. The plane formed by L1 and L2 is denoted by a, whose normal at R is denoted by line N" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002685_rnc.1378-Figure10-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002685_rnc.1378-Figure10-1.png", "caption": "Figure 10. The CTOL aircraft.", "texts": [ " THE EXACT OUTPUT TRACKING PROBLEM FOR THE CTOL AIRCRAFT Consider the simplified CTOL dynamic equation introduced in [8]:( x\u0308 y\u0308 ) = R( ) (\u2212D L ) +R( ) ( u1 \u2212 u2 ) + ( 0 1 ) \u0308 = u2 where (x, y) is the aircraft center of mass, is the pitch angle, u1 and u2 are the controls, R( )= ( cos \u2212sin sin cos ) Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc is the rotation matrix, is the flight path angle, D and L are the drag and lift forces given by L=aLv2(1+c ), D=aDv2(1+b(1+c )2) where v= \u2225\u2225\u2225\u2225\u2225 ( x\u0307 y\u0307 )\u2225\u2225\u2225\u2225\u2225 Copyright q 2008 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2009; 19:1168\u20131196 DOI: 10.1002/rnc is the aircraft speed, = \u2212 is the angle of attack and aL , aD , c, b and are given constants (see Figure 10). Given a curve \u2208C2(R,R2), if ( x(0) y(0) ) = (0), the controls that allow to exactly track are given by ( u1 \u2212 u2 ) = R(\u2212 ) [ \u0308+ ( 0 1 )] +R(arg(\u0307)\u2212 ) (\u2212D L ) where L = aL\u2016\u0307\u20162(1+c( \u2212arg(\u0307))) D = aD\u2016\u0307\u20162(1+b(1+c( \u2212arg(\u0307))2)) and is the solution of the following equation: \u0308=\u2212 \u22121 [\u2329(\u2212sin cos ) , \u0308+ ( 0 1 )\u232a + \u2329(\u2212sin( \u2212arg(\u0307)) cos( \u2212arg(\u0307)) ) , ( D \u2212L )\u232a] (10) which represents the unstable internal dynamics. Our problem is to show that there exists an initial condition ( 0, \u03070) such that the solution of the internal dynamics (10) remains small" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002750_12.762111-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002750_12.762111-Figure1-1.png", "caption": "Fig. 1: Design of the flexible polymer based electro-enzymatic glucose sensor showing dimensions for three different designs.", "texts": [ " Furthermore, the process can be used for batch fabrication of many sensors at once, which is crucial for commercialization of any glucose sensor. The sensor design avoids any manual assembly or alignment steps and only uses a mask aligner, a metal deposition system, a spin coater, and hot plates. The presented sensors are intended to be very cost effective, simple to fabricate, and easy to release after complete fabrication. In order to test and characterize different properties of the flexible electro-enzymatic glucose sensor, we designed a flexible sensor with sandwiched gold electrodes between two SU-8 layers (Fig. 1). The bottom layer of the sensor is a 100\u00b5m thick rectangular SU-8 layer. This bottom layer is little larger than the area occupied by the electrodes. The top sensor layer is designed using a 100\u00b5m thick SU-8 layer with contact pad openings and electrode openings (Fig. 1). The electrode opening is provided only in the sensing area. Two separate openings are designed so that any of the gold electrodes can be used as an active electrode by immobilizing the enzyme on them. The area of the top SU-8 layer is a little smaller than the bottom layer area and the contact pad opening is not fully enclosed. And, as mentioned previously, both gold electrodes are sandwiched between the two SU-8 layers. Three designs of electrodes are used to fabricate different types of sensors (see Fig. 1). All three designs consist of three sections: a sensing area, a connecting conductor, and a contact pad. Each sensor die is 5 mm x 10mm in size and is fabricated with gold electrodes sandwiched between two layers of SU-8 (Fig. 1). In a single step metal patterning process, the active and the reference electrodes, the connecting conductors, and the contact pads were fabricated using only gold. It should also be noted that the sensor dimensions, including bond pad dimensions, were chosen for easy handling and testing; the batch fabrication process could be utilized for both larger and much smaller sensors, with scaling effects currently under investigation. The actual sensing area, in all three designs, is clearly labeled in Fig. 1. The sensing area in the first design is 2 mm x 3 mm (Fig. 1) and 2mm x 2.5mm in the other two designs (Fig. 1). However, the active and reference electrode areas in Proc. of SPIE Vol. 6886 68860G-3 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 07/02/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx all three designs are 2mm x 2mm in size. Large contact pads are designed for each electrode to make wire connections easier. Each 2mm x 2mm contact pad is connected with a 6mm long gold conductor. The openings in the top SU-8 layer for the contact pads and the sensing electrode (Fig. 1) are made for the physical connection with wires and for contact with the enzyme respectively. The gold conducting wire is covered with SU-8 to avoid direct external contact (Fig. 1). The hybrid polymer fabrication process of the flexible electro-enzymatic glucose sensor is carried out using PDMS as the process substrate. The PDMS substrate is used to facilitate release of the dual layer sandwiched SU-8 sensor. Moreover, the PDMS process substrate helps to achieve the desired surface profile from the SU-8 process without any process related issues such as cracking in the SU-8 layer. A handle glass wafer or glass slide is used as a support for the PDMS process substrate, and can be reused for many process cycles" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003551_etep.1613-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003551_etep.1613-Figure2-1.png", "caption": "Figure 2. Cross section of a salient pole machine for different axial eccentricities: (a)SAE, (b) DAE, and (c) MAE.", "texts": [ " As mentioned in the previous section, axial eccentricity fault appears when there is variation of the eccentricities along the rotor axis. Therefore, axial eccentricity can be treated as a variable radial eccentricity. There are three types of axial eccentricity: static axial eccentricity (SAE), dynamic axial eccentricity (DAE), and mixed axial eccentricity (MAE). SAE, DAE, and MAE in a salient pole machine are presented in Figure 1 to help obtain the model of the machine in these conditions. The thin line corresponds to the rotor axis, the thick line to the stator axis, and the dotted line to the rotor rotation axis. Figure 2 shows the cross section of an elementary salient pole machine for three types of axial eccentricities. In the case of SAE, the rotor axis and the rotation axis coincide with each other and are inclined compared with the stator one. In this case, the rotor rotates about its own axis, but this axis does not coincide with that of the stator. The air-gap length variation under SAE for salient pole synchronous machine can be described by air-gap function as follows: g \u2019; \u03b8; \u2019s; z\u00f0 \u00de \u00bc gh \u2019; \u03b8\u00f0 \u00de Asc z\u00f0 \u00de cos \u2019 \u2019s\u00f0 \u00de (1) where gh(\u2019, \u03b8) is effective air-gap function for the healthy machine, Asc(z) is the distance of stator and rotation axes, \u2019 is angle in stator reference frame, and \u2019s is angle at which rotation and stator axes are separated (Figure 2a). The static eccentricity coefficient is defined as follows: ds z\u00f0 \u00de \u00bc Asc z\u00f0 \u00de gp (2) where gp is the air-gap length above pole shoes for the healthy machine. Replacing Asc(z) from Equation (2) in Equation (1) yields g \u2019; \u03b8; \u2019s; z\u00f0 \u00de \u00bc gh \u2019; \u03b8\u00f0 \u00de gpds z\u00f0 \u00de cos \u2019 \u2019s\u00f0 \u00de (3) Therefore, the air-gap function for salient pole synchronous machine with SAE is described as follows: g \u2019; \u03b8; \u2019s; z\u00f0 \u00de \u00bc gh \u2019; \u03b8\u00f0 \u00de 1 gpds z\u00f0 \u00deg 1 h \u2019; \u03b8\u00f0 \u00de cos \u2019 \u2019s\u00f0 \u00de (4) The static eccentricity coefficient is the function of the position along the axial direction", "1002/etep ds z\u00f0 \u00de \u00bc tan as\u00f0 \u00de z ls\u00f0 \u00de=gp (5) where as is the inclined angle of the rotor and ls is the shaft misalignment level in the case of SAE. As shown in Figure 1b, in the case of DAE, the rotor symmetrical axis is inclined compared with the stator symmetrical axis, which is superimposed to the rotor rotation axis. The air-gap length variation under DAE can be described by air-gap function as follows: g \u2019; \u03b8; \u2019d; z\u00f0 \u00de \u00bc gh \u2019; \u03b8\u00f0 \u00de 1 gpdd z\u00f0 \u00deg 1 h \u2019; \u03b8\u00f0 \u00de cos \u2019 \u2019d\u00f0 \u00de (6) where \u2019d is angle at which rotation and rotor axes are separated (Figure 2b). The dynamic eccentricity coefficient, dd(z), is the function of the position along the axial direction. Through geometric analysis on Figure 1b, the dynamic eccentricity level at any point along the shaft is derived as follows: dd z\u00f0 \u00de \u00bc tan ad\u00f0 \u00de z ld\u00f0 \u00de=gp (7) where ad is the inclined angle of the rotor and ld is the shaft misalignment level in the case of DAE. In reality, static and dynamic eccentricities tend to coexist. In the case of the MAE, rotor axis and rotation axis are separated from each other and are inclined compared with stator axis (Figure 1c). The air-gap length variation under MAE can be described by air-gap function as follows: g \u2019; \u03b8; \u2019s; \u2019d; z\u00f0 \u00de \u00bc gh \u2019; \u03b8\u00f0 \u00de 1 gpds z\u00f0 \u00deg 1 h \u2019; \u03b8\u00f0 \u00de cos \u2019 \u2019s\u00f0 \u00de gpdd z\u00f0 \u00deg 1 h \u2019; \u03b8\u00f0 \u00de cos \u2019 \u2019d\u00f0 \u00de (8) Through geometric analysis on Figure 2c, it is easy to show that d z; \u2019s; \u2019d\u00f0 \u00de \u00bc ds z\u00f0 \u00de2 \u00fe dd z\u00f0 \u00de2 \u00fe 2ds z\u00f0 \u00dedd z\u00f0 \u00de cos \u2019d \u2019s\u00f0 \u00de h i1=2 (9) \u2019m z; \u2019s; \u2019d\u00f0 \u00de \u00bc \u2019s \u00fe sin 1 dd z\u00f0 \u00de sin \u2019d \u2019s\u00f0 \u00de d z; \u2019s; \u2019d\u00f0 \u00de (10) where d is the general eccentricity factor and \u2019m is the angle at which rotor and stator axes are separated. The air-gap function can then be defined for salient pole machine with a general eccentricity fault including static, dynamic, and mixed eccentricities as follows: g \u2019; \u03b8; \u2019s; \u2019d; z\u00f0 \u00de \u00bc gh \u2019; \u03b8\u00f0 \u00de 1 gpd z; \u2019s; \u2019d\u00f0 \u00deg 1 h \u2019; \u03b8\u00f0 \u00de cos \u2019 \u2019m z; \u2019s; \u2019d\u00f0 \u00de\u00f0 \u00de (11) It is clear that SAE and DAE are special cases of the MAE, where dd(z) = 0 in SAE and ds(z) = 0 in DAE" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002520_s00542-009-0849-7-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002520_s00542-009-0849-7-Figure5-1.png", "caption": "Fig. 5 Aerodynamic force calibration setup", "texts": [ "3 Force calibration Noting the study results given in Table 4, for the verification of the nozzle design, this work performs a series of applied aerodynamic force calibration tests with a dummy spindle motor that embeds a load cell as shown in Fig. 4. The hub of the dummy spindle motor is divided in two pieces and one of the pieces is separated from the structure so that it can be freely moved when it is under an external Fig. 3 Final nozzle design force input. The other piece of hub that embeds a load cell is attached to the motor base. The dummy spindle is then located into the designed air nozzle structure, as shown in Fig. 5, for the measurement of the aerodynamic force input. An air pressure gauge is also attached at the top of the air channel to monitor the flow conditions. And the nozzle is connected to an air compressor through a hose. A remote controllable solenoid valve is used for air flow adjustment. The measured signal from the load cell is amplified using a strain amplifier (Kyowa DPM-712A) and the data are collected by a computer using a DAQ board. The force calibration was repeated many times at various different air pressures" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002838_biorob.2008.4762869-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002838_biorob.2008.4762869-Figure3-1.png", "caption": "Fig. 3. (a) 3D bending view, (b) Orthogonal projection in the bending plane \u03a0 and (c) Bottom view", "texts": [ " The shape of the snake like unit is defined by three auxiliary backbones. In our application, two pull-wires running all the way through the endoscope shaft control the angle of the tip deflection. Theses pull-wires are controlled by rotary motors mounted on the endoscope handle. The link between the motors and the wires is ensured by pulleys of radius Rp. For modeling, we assume that the wires and the continuum section are inextensible and that the wires are equally spaced on the bending section circumference (see figure 3). We aim at controlling the pose r of the endoscopic camera. Hence we are interested in the relationship between the rotary motors positions q and the homogeneous transformation Tb\u2192c expressing position tb\u2192c and orientation Rb\u2192c of the camera frame Fc with respect to a reference frame Fb attached to the beginning of the bending section (cf Fig 2). The camera being embedded in a rigid section of length Lt at the extremity of a bending section of backbone length Lf , the transformation Tb\u2192c depend on \u03b1 the orientation of the bending plane \u03a0 with respect to the base frame and \u03b2 the bending angle in \u03a0. The rotation matrix Rb\u2192c is defined by a rotation \u03b2 around the unit vector u normal to the plan \u03a0 (cf Fig 3(c)) : Rb\u2192c = 0 @ s2\u03b1 + c\u03b2c2\u03b1 \u2212s\u03b1c\u03b1(1 \u2212 c\u03b2) c\u03b1s\u03b2 \u2212s\u03b1c\u03b1(1 \u2212 c\u03b2) c2\u03b1 + c\u03b2s2\u03b1 s\u03b1s\u03b2 \u2212c\u03b1s\u03b2 \u2212s\u03b1s\u03b2 c\u03b2 1 A (1) The translation vector tb\u2192c is obtained from the translation between Fb and Fc expressed in the bending plan frame (cf Fig 3(b)) : . tb\u2192c = 0 B@ Ltc\u03b1s\u03b2 + Lf \u03b2 (1 \u2212 c\u03b2)c\u03b1 Lts\u03b1s\u03b2 + Lf \u03b2 (1 \u2212 c\u03b2)s\u03b1 Ltc\u03b2 + Lf \u03b2 s\u03b2 1 CA (2) The forward kinematic model is then obtained by expressing the relation between angles \u03b2 and \u03b1 and the actuator positions q = [q1 q2] T . At rest (q = [q10 q20] T ), when no efforts are applied on the wires, the bending part of the endoscope is straight. The rotation \u2206qi = qi\u2212qi0 of the pulley i acts on the wire i and hence modifies the distribution of the length along the bending part, so that the length modification is given by : \u2206li = Rp\u2206qi, \u2200i \u2208 {1, 2} (3) The length distribution of the two wires \u2206l1 and \u2206l2 can be related to the angles \u03b1 and \u03b2 using \u2206r1 and \u2206r2 as shown on Fig 3(b) and 3(c) : Lf + \u2206li = Lf R (R + \u2206ri) \u21d2 \u2206li = Lf R \u2206ri, i \u2208 {1, 2} with \u2206r1 = \u2212D 2 cos(\u03b1), \u2206r2 = \u2212D 2 sin(\u03b1) and \u03b2 = Lf R We find \u2206l1 = \u2212D 2 \u03b2cos(\u03b1), \u2206l2 = \u2212D 2 \u03b2sin(\u03b1) which gives \u03b2 = 2Rp D q \u22062 q1 + \u22062 q1 (4) and \u03b1 = atan2(\u2212\u2206q2,\u2212\u2206q1) (5) Finally from equations 1, 2, 4, 5 we get the kinematic model Tb\u2192c(q1, q2). Let \u0398 = [\u03b1, \u03b2]T and \u2206l = [\u2206l1,\u2206l2] T . The camera screw V = [v, \u03c9]T expressed in Fc is then related to the joint velocity q\u0307 by the robot Jacobian Jq : V = Jq q\u0307 with Jq = \u2202r \u2202\u0398 \u2202\u0398 \u2202\u2206l \u2202\u2206l \u2202\u2206q where \u2202\u2206l \u2202\u2206q = \u201e Rp 0 0 Rp \u00ab , \u2202\u0398 \u2202\u2206l = 0 @ \u2212 \u2206l2 \u22062 l1 +\u22062 l2 \u2206l1 \u22062 l1 +\u22062 l2 2 D \u2206l1\u221a \u22062 l1 +\u22062 l2 2 D \u2206l2\u221a \u22062 l1 +\u22062 l2 1 A and \u2202r \u2202\u0398 = 0 BBBBBBB@ \u2212Lf \u03b2 s\u03b1(1 \u2212 c\u03b2) \u2212 Lts\u03b1s\u03b2 Lf \u03b22 c\u03b1(1 \u2212 c\u03b2) + Ltc\u03b1 Lf \u03b2 c\u03b1(1 \u2212 c\u03b2) + Ltc\u03b1s\u03b2 Lf \u03b22 s\u03b1(1 \u2212 c\u03b2) + Lts\u03b1 0 Lf \u03b2 (1 \u2212 s\u03b2 \u03b2 ) \u2212c\u03b1s\u03b2 \u2212s\u03b1 \u2212s\u03b1s\u03b2 c\u03b1 \u22121 + c\u03b2 0 1 CCCCCCCA The model of the visual loop is given in Fig 4" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003282_978-0-387-74244-1_11-Figure11.2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003282_978-0-387-74244-1_11-Figure11.2-1.png", "caption": "FIGURE 11.2. A vehicle with roll, and yaw rotations.", "texts": [ "5) The roll model vehicle dynamics can be expressed by four kinematic variables: the forward motion x, the lateral motion y, the roll angle \u03d5, and the roll angle \u03c8. In this model, we do not consider vertical movement z, and pitch motion \u03b8. 11.2 F Equations of Motion A rolling rigid vehicle has a motion with four degrees of freedom, which are translation in x and y directions, and rotation about the x and z axes. The Newton-Euler equations of motion for such a rolling rigid vehicle in the body coordinate frame B are: Fx = mv\u0307x \u2212mr vy (11.6) Fy = mv\u0307y +mr vx (11.7) Mz = Iz\u03c9\u0307z = Iz r\u0307 (11.8) Mx = Ix\u03c9\u0307x = Ixp\u0307. (11.9) Proof. Consider the vehicle shown in Figure 11.2. A global coordinate frame G is fixed on the ground, and a local coordinate frame B is attached to the vehicle at the mass center C. The orientation of the frame B can be expressed by the heading angle \u03c8 between the x and X axes, and the roll angle \u03d5 between the z and Z axes. The global position vector of the mass center is denoted by Gd. The rigid body equations of motion in the body coordinate frame are: BF = BRG GF = BRG \u00a1 mGaB \u00a2 = m B GaB = m Bv\u0307B +m B G\u03c9B \u00d7 BvB. (11.10) BM = Gd dt BL = B GL\u0307B = BL\u0307+ B G\u03c9B \u00d7 BL = BI B G\u03c9\u0307B + B G\u03c9B \u00d7 \u00a1 BI B G\u03c9B \u00a2 " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002371_0022-4898(66)90051-6-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002371_0022-4898(66)90051-6-Figure1-1.png", "caption": "FIG. 1. Free-body diagram of t ractor- trai ler vehicle negotiating a vertical obstacle.", "texts": [ "friction properties of a known coefficient of friction. Throughout the analysis, it is also assumed that the wheels and suspension of the vehicle are rigid. that is, the deformations of the pneumatic tires and of the suspension are neglected in the interest of simplification. All the wheels of the combined vehicle are supposed to have the same diameter. T R A C T O R A N D TOWED TRAILER (a) Perlormance of front wheels A free-body diagram of a tractor and trailer combination climbing a vertical obstacle is shown in Fig. 1. The forces N1, No and N:, are the normal components of the reactions at the points of contact of the wheels with the ground. When slipping impends, the tangential components at the driven wheels of the tractor are ~N1 and p.N~. The tangential component at the towed trailer wheels is simply fN3, where [ is the coefficient of rolling resistance. If the trailer is in equilibrium, then the conditions of equilibrium can be written as N3 + V = Wt H = fN3 N,4 +/N3 (1\" + he) - W~s = 0 Hence, the force acting between the tractor and trailer has the components [ S l w ~ and V = 1 t + f ( r + h c ) (I) /s H - W, t+I ( r+hc) Similarly, the three equations of equilibrium of the tractor for the vertical and horizontal forces and for the moments about the axis of the front wheel are : N~ sin a + #N~ cos a + No = W + V /~N~ sin a - N ~ cos a+ I", "= D~+E 2 (21) In the special case of a very small rolling resistance q = 0 ) , equation (19) is further modified to tan c~ = sW, -#he (W + W,) tz [tW + ( t - s) W,] As an example, the obstacle climbing performance of the trailer wheels for the reference tractor-trailer combination is presented in Fig. 9. As shown by the curves, TRACTOR AND DRIVEN CARRIER (a) Per[ormance of [ront wheels If a free-body diagram is drawn for the case of a tractor and powered two-wheel carrier, the only difference from Fig. 1 is that the tangential reaction on the driven OBSTACLE PERFORMANCE OF ARTICULATED WHEELED VEHICLES 51 wheels of the carrier is/~N3, and the direction of this force and of the horizontal tow point reaction is reversed. In consequence, the analysis developed in the preceding section is applicable to the present problem if, throughout the analysis, the coefficient I is replaced by ( -/~). In so doing, the force transmitted from the carrier to the tractor has the com- ponents s ] p.s W, (22) V= 1-.t_tz~r+hc) W, H= t-I~(r+hc) The performance equation again appears in the form where ( r cos a = 1 - # sin ot +/" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002897_tmag.2007.916494-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002897_tmag.2007.916494-Figure7-1.png", "caption": "Fig. 7. Compared region.", "texts": [ " In the finite-element analysis, we assumed that the magnetization voltage was 1500 V, the capacitance was 3000 F, and the conductivity of the permanent magnet was S/m, respectively. The easy axis of the permanent magnet was assumed to be parallel to the -axis. The anisotropy field was A/m. Fig. 5 shows the initial magnetization curve in the easy and hard axis. In verification of this method, we compared the calculated results and the measured ones. Fig. 6 shows the magnetizing coil. We can produce a strong magnetic field up to 6 T at the center with this magnetizing coil. We placed a permanent magnet in the magnetization coil and magnetized. Fig. 7 shows the evaluated lines used in the comparison. The distance between the evaluated lines and the magnet surface was 1 mm. In the measurement, we used a 3-D magnetic field distribution measurement device (manufactured by IMS Co. Ltd.). Figs. 8\u201310 show the comparison between calculated results and measured ones. The calculated results agreed well with the measured results at near the center of magnetic pole. However, the differences between the calculated results and the measured results are observed near the edge parts increase" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002930_icelmach.2008.4799943-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002930_icelmach.2008.4799943-Figure1-1.png", "caption": "Fig. 1. Schematic diagram of a three phase synchronous machine", "texts": [ " Hence, the variation of the air-gap length of such type of machines deviate far from the sinusoidal function if the reluctance component isn't good optimised or if the machine operates under high load condition (saturation effect). Since the theory of the salient pole synchronous machine until now is based only on the idealized machines (with sinusoidally variation of the air-gap length), the main motivation of this work was to investigate the influence of the effective air-gap length, especially the rotor permeance effect on the performances of this type of machines. II. SALIENT POLE SYNCHRONOUS PM MACHINE WITH SINUSOIDAL AIR-GAP LENGTH The following figure 1 shows the cross section and the winding distribution of a three phase synchronous machine. One side of the coil (coil side) is represented by a \"+\" indicating that the assumed positive direction of current is down the length of the stator (into the paper). The \".\" indicates that the assumed positive direction of current is out of paper. Depending on the reference frame used. self-, mutual- and phase inductances, or dq-inductances appear in the main mathematical relations of the electrical machines" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001424_jmes_jour_1973_015_066_02-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001424_jmes_jour_1973_015_066_02-Figure1-1.png", "caption": "Fig. 1. Geometry of vertical axis castor", "texts": [ " The theory presented below does offer some explanation of these cases and covers tyre deformation as a special case. A constant forward speed v will be assumed and the steering axis will be constrained to the vertical; therefore, gyroscopic couples can be ignored. Contact patch deformation is ignored in the first instance, i.e. \u2018hard\u2019 contact is assumed. Simple coulomb friction will be used, with frictional force proportional to normal reaction R, namely pR, with p the dynamical-friction coefficient. This force opposes directly the relative slip-velocity vector. 2 WHEEL LOCKED CASE Fig. 1 shows the basic geometry and Fig. 2 shows the resulting contact patch-force components, with V = dv2 + + 2vtd sin e. This gives the following equation of motion for a rotation l3 about the steering axis. t2&R vt sin epR V Id = --- V where Z is the moment of inertia about the steering axis. The MS. of this paper was received at the Institution on 6th November Simon Engineering Laboratories, University of Manchester, t References are given in the Appendix. I972 and accepted for publication on 19th June 1973" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000251_s1474-6670(17)63640-1-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000251_s1474-6670(17)63640-1-Figure3-1.png", "caption": "Fig. 3 Admissible gain region for flight condition 2", "texts": [ " (8) can be used to express the two free-gains by coefficients of a second order factor of the characteris tic polynomial, which is varied along the boundary in s-plane to produce the boundary in the plane of the two free gains, see Franklin (1980). Alternatively the genera lized D-decomposition technique described by Kaesbauer (1981) may be used. This tool will be applied to the aircraft example in the next section. ROBUSTNESS WITH RESPECT TO FLIGHT CONDITION The first design objective will be to design an output feedback controller, eq. (4), whi ch meets the nominal pole region requirements at all four flight conditions. The boundary for flight condition 2 is shown in Fig. 3. On a-b eigenvalues are on the lower natural frequency boundary w = 3.5, sp on b-c they are on the damping 0.35 lines. At c a real root boundary takes over: on c-d the actuator eigenvalue is at a = -70. On d-e a real short period eigenvalue is at the upper natural frequency limit a = -12.6 and for e-a the actuator eigenvalue is at a = -12.6. The condition for having no real root a = -3.5 is satisfied in the total re gion. This region R nom2 is bounded by two straight lines c-d and d-a resulting from real root conditions and by the two complex boundary curves a-b and b-c" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000266_robot.1988.12184-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000266_robot.1988.12184-Figure3-1.png", "caption": "Figure 3: Location of the positioning table and the weld part with respect to the world frame 0.", "texts": [ "00 0 1988 EEE (constraint (c) in the above), the motion of redundant system is generated through a nonlinear optimization process. This consists of the Cartesian motions of the robot (a Cybotech W 1 5 , is used in our study) and that of the track. Inverse kinematics is next used to compute the motions of the WV15 joints and that of the track. Again, the inverse kinematics is used to provide precise arm motions. (v) 0 R T ~ ~ ~ TorgRTbl TblRp,tPartRsur = II. Geometric Model of the Weld Contour and the Positioning Table Figure 3, indicates the relative location of the positioning table and the part with respect to a reference frame 0 (world origin frame). The transforms are defined in the figure. Geometric Description of Weld Part We adopt cylindrical coordinates to describe the position of the weld contour on the surface of the part, with respect to a part reference frame. We assume the shape of the part is arbitrary. Figure 4, defines the variables used in the geometric modelling of the part. If partEur is a position vector which is located on the weld contour it is defined as: where r, cy and z are subject to a surface equation: P a r t ~ u r = (r cosa, rsincy, z ) ~ surf(r, cy, z) = 0 (1) (2) Let be the direction of the surface normal, and sur be the direction of the surface tangent aligned along the weld contour" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001375_095440605x32020-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001375_095440605x32020-Figure5-1.png", "caption": "Fig. 5 Experimental bench", "texts": [ " The principle of time interval measurement of pulse signals makes it easy to acquire the instantaneous angular displacement and determine the amplitude and phase of the torsional vibration by eliminating the effect of uneven graduations among the ranked teeth. At the same time, it can measure the lateral vibration with two sensors installed on both sides of the rotor at one section in the horizontal direction. This system can distinguish an angle of torsional vibration of 0.00188 as well as accurate lateral vibration. As shown in Fig. 5, the direct current motor (ZYT11004) on the left is used to drive the shaft system. Its speed can be controlled. The shaft system is made up of four segments, which are 180, 310, and 150 mm long and the diameters are all 10 mm. Each segment is supported by two centreadjustable rolling bearings. Because the torsional rigidity of each segment is large, flexible couplings are used to link them and the motor. There are four discs on the first segment, one on the second, and one on the third. The diameters of these discs are 75 mm", " In other words, when the graduation is calibrated at a certain rotating frequency with no man-made misalignment, the real vibrations at one-, two-, or many-times rotating frequencies caused by inherent unbalance or inherent misalignment at these rotating frequencies are already added to the graduation of the measured discs\u2019 teeth, and the measured vibrations at these frequencies are zero. In this experiment, the graduation of the measured disks\u2019 teeth at rotating frequency 45.3 Hz is calibrated. Forty micrometer thick spacers are put under each bearing E3 and E4 (Fig. 5) to create a man-made misalignment, the measured result is shown in Fig. 6. From this figure, it is found that the torsional vibrations on the first segment are much smaller than those on the second and third. The reason is that the inertia of this segment is much larger than the second and the third ones, and there is nomisalignment in this segment. However, the torsional vibration at natural torsional frequency is obvious on this segment. On the second and third segments, where misalignment exists, the torsional vibrations at 1 rotating frequency are very obvious, accompanied by smaller torsional vibrations at 2 , 3 , 4 , 5 , 6 and natural torsional frequency" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000333_2004-01-3017-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000333_2004-01-3017-Figure1-1.png", "caption": "Figure 1. Construction of an alternator overrunning decoupler", "texts": [ " It has been widely used in the automotive industry most recently as a new technology to solve FEAD system problems induced by the large dynamic belt tension fluctuation. This paper will introduce the functional principles of the Overrunning Alternator Decoupler (OAD). Those functions will be validated through the performance tests of the FEAD system on the engine. Through analysis of the test data and numerical simulation, the benefits of the OAD on the FEAD system will be revealed. The OAD consists of a pulley, a one-way clutch assembly, a tuned isolator spring, a shaft and a ball bearing, see Figure 1. 2/7 The OAD is connected to the alternator through the OAD shaft. The ball bearing carries the hubload and allows the relative motion between the OAD pulley and shaft. The drive torque from the belt that applies on the OAD pulley will be transferred to the alternator rotor through the clutch assembly, isolator spring and the OAD shaft. The spring isolates the alternator rotor from the torsional vibration to reduce the influence of the rotor inertia on the system. The one-way clutch assembly allows the alternator rotor to run faster than its drive pulley during fast engine deceleration" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001476_s1064230706030099-Figure21-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001476_s1064230706030099-Figure21-1.png", "caption": "Fig. 21.", "texts": [ " Consider the problem 0 p, J u( ) u t( ) t ; y\u0307\u0307 u mind 0 12 \u222b \u2013y u w,+ += = Construct the synthesis of optimal control in the class of inertial relay controls for u(tk + 0) = 0.5, k = ; w(t) = w1sint, t \u2208 [0, t\u2217 = 5.8]; w(t) = 0, t \u2208 ]t\u2217, t* = 12]; and h0 = 0.1. For M = 1.25 and w1 = 0.13, the optimal vector of instants of switching \u03c4 = (t1, t2, t3, t4) = (0.6237699, 2.2816412, 7.186079, 8.4150114) with u0 = 0 and the value of the quality criterion J(u) = 3.42020 was constructed. In this case, the optimal pro- gram with the vector of switching \u03c40 = ( , , , ) = (0.6135205, 2.32874195, 6.896706, 8.6119273), u0 = 0 was acting on the interval [0, h0 = 0.1]. In Fig. 21, the phase trajectories generated by the optimal open-loop control and the realization of the optimal feedback are shown. The simplest dynamical regulator is described by a differential equation where \u03bd(t), t \u2265 0, is the controlling signal and u(t), t \u2265 0 is an inertial control action of order k. N.S. Pavlenok investigated optimal control problems in the class of inertial controls of order one [22, 23] and two taking into account geometrical constraints on \u03bd, u, and by means of the methods of Section 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001775_iecon.2006.348122-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001775_iecon.2006.348122-Figure2-1.png", "caption": "Fig. 2. Prototype of single-phase LSPM", "texts": [ " Though this makes difficult to analyze single-phase LSPM, recent studies show the analysis method of single-phase LSPM[1][2]. Iron loss calculation method is also reported[3]. Though these analytic tools of this motor are developed, the effect of considering squirrel cage windings into analysis hasn\u2019t been discussed. This paper aims at clarifying the effect of the cage at steady state. To explain this, analysis of neglecting the cage is compared with experimental values. And by using the model of considering the cage and experimental values, the effect of the cage is considered. Fig.2 shows the prototype of single-phase LSPM tested in this paper. The structure of this prototype is that the arc shaped PM is inserted in the rotor of multipurpose single-phase induction motor. So no optimization technique is done for this structure. Two types of rotors are constructed as shown in Fig.3. By comparing analysis of Rotor A with experimental value, the effect of squirrel cage in single-phase LSPM is examined. Furthermore, characteristics of Rotor A and Rotor B calculated by using FEM is shown to discuss the magnetic circuit" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000251_s1474-6670(17)63640-1-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000251_s1474-6670(17)63640-1-Figure2-1.png", "caption": "Fig. 2 Required pole region", "texts": [ "7 was chosen for minimum drag. Thus the short period mode stabilization is a single input problem. The required closed loop eigenvalue locations are given by military specifications for flying qualities of piloted airplanes (1969). For the short period mode described by s2 + 2s W s + w2 0 (2) sp sp sp the restricted range of damping s and natural frequency W is sp sp 0.35 ~ s sp ~ 1.3 (3) where wa and wb depend on the flight condi tion and are given in the appendix for the four conditions considered here. Fig. 2 shows the nominal region r., eq. (3) J together with the open loop eigenvalues for a subsonic flight condition j. Damping greater than one in eq.(3) corresponds to two real eigenvalues. Eq.(3) would admit some real pairs of poles with one of them outside the region r .. In the following no use is made J of this possibility. For all real pairs in side r . condition (3) is satisfied. We reJ quire, that the closed loop short period poles of each flight condition j = 1, 2, 3, 4 are loca ted in the respective region r ", "(3) and a natural frequency range wb ~ W ~ wd ' wd = 70 rad/sec is chosen in order Eo main- The assumed type of sensor failure is that the nominal gain v = 1 is reduced to some value 0 ~ v < 1. As far as eigenvalue loca tion is concerned, only this multiplicative error is important. There may be an additive bias or noise tenn, which should be removed by a failure detection system at a higher hierarchical level. The objective of this paper is to design the basic level control system such that the pole region requirements of Fig. 2 are robust with respect to changing flight conditions and sensor failures. This is an example for the application of a novel parameter space design technique and generalized D-decomposition, see Ackermann and Kaesbauer (1980, 1981). It will be reviewed briefly in the following paragraph. In application to the example it is then shown, how robustness with respect to changing flight conditions can be achieved by appropriate choice of kN and k in an z q output feedback control law u - [k Nz k q 0] ~ (4) For robustness with respect to sensor fai lures Franklin (1980, 1981) studied a con figuration with two gyros and one accelero meter and dynamic feedback" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003954_iros.2010.5650000-Figure14-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003954_iros.2010.5650000-Figure14-1.png", "caption": "Fig. 14. Two polyhedral objects for peg-in-hole insertion: (a) before triangulation and (b) after triangulation.", "texts": [ " It is the same as the methods proposed by Hirai [5] and Xiao [7]. For Hirai\u2019s method, the additional computation of l times is required to give the transitional connections between contact states. So its computational complexity is actually O(n2l) considering the additional l times computation. Xiao applies infinitesimal motions to find new contact states. If it needs k steps to reach a new contact state, the computational complexity is O(n2k). Therefore, the proposed method is more efficient than the other methods. A square peg-in-hole assembly of Fig. 14(a) is studied to verify the proposed method. First we define the topologies and geometries of the objects like the table of Fig. 2. The hierarchical structures of adjacent vertices and edges are constructed for peg and hole. Then all faces of each polyhedral object are triangulated according to the procedure described in Section II-B as shown in Fig. 14(b). Fig. 15 shows the structures of adjacent triangles. If CSinitial={(v1, t19)} is given, the adjacent elements of v1 are selected as {v2, v4, v5}, {e1, e4, e9}, and {t1, t2, t3} from Fig. 15(a). Also the adjacent elements of t19 are selected as {v7, v11, v15}, {d8, e15, e27}, and {t15, t20, t24} from Fig. 15(b). Since the type of adjacent triangles {t15, t20, t24} is {B, A, B}, we use Type-6 of Table I to obtain adjacent sub-contacts. Contact states are generated by combination of CSinitial and the adjacent sub-contacts" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002821_j.ijsolstr.2008.05.021-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002821_j.ijsolstr.2008.05.021-Figure4-1.png", "caption": "Fig. 4. Periodical loading path applied on the system.", "texts": [ " 3) are given by qel\u00f0h;b\u00de \u00bc F\u00f0b\u00de 2pR Sf1 Sf2 St1 St2 1\u00feSt1 St2 ! cos\u00f0b\u00de sin\u00f0h\u00de \u00fe b 3p 2 cos\u00f0h b\u00de if p=2 6 h < b F\u00f0b\u00de 2pR Sf1 Sf2 St1 St2 1\u00feSt1 St2 ! cos\u00f0b\u00de sin\u00f0h\u00de \u00fe b\u00fe p 2 cos\u00f0h b\u00de if b < h 6 3p=2 8>>><>>>: \u00f017\u00de pel\u00f0h; b\u00de \u00bc p0 \u00fe F\u00f0b\u00de 2pR Sf1 Sf2 St1 St2 1\u00feSt1 St2 ! cos\u00f0b\u00de cos\u00f0h\u00de b 3p 2 sin\u00f0h b\u00de if p=2 6 h 6 b p0 \u00fe F\u00f0b\u00de 2pR Sf1 Sf2 St1 St2 1\u00feSt1 St2 ! cos\u00f0b\u00de cos\u00f0h\u00de b\u00fe p 2 sin\u00f0h b\u00de if b 6 h 6 3p=2 8>>>><>>>: \u00f018\u00de For the sake of clarity, a rotating radial load with a constant intensity F > 0 (see Fig. 4) is considered in the whole following analysis: F\u00f0b\u00de \u00bc F 8 p=2 6 b 6 3p=2 \u00f019\u00de The statement of Melan\u2019s theorem adapted to the model is the following: If a self -tangential force q \u00f0h\u00de and a coefficient m > 1 exist such that the tangential force eq\u00f0h;b\u00de \u00bc m\u00f0qel\u00f0h;b\u00de \u00fe q \u00f0h\u00de\u00de satisfies for all p=2 6 h 6 3p=2 and for all p=2 6 b 6 3p=2 the condition jeq\u00f0h; b\u00dej 6 k\u00f0h; b\u00de; then there is necessarily a slip-shakedown whatever be the initial conditions: In the present case, the set RS is reduced to the relative rigid slips leading to an arbitrary rotation of solid (1), i" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000458_bf01922914-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000458_bf01922914-Figure1-1.png", "caption": "Fig. 1. Inhibition of A extrasynaptosomal and B synaptosomal GABAT by Lioresal. Reciprocal plot (according to LineweaverBurk) of 1/v against (GABA)-1 concentration at 2 concentrations", "texts": [], "surrounding_texts": [ "1136 Specialia Experientia 34/9\nExperiments on the mechani sm of the inhibition of mitochondrial Ca 2+ transport by La 3+ and ruthenium red l\nVerena Niggli, P. Gazzott i and E. ca ra fo l i\nSummary. The effects o f La 3+ and ruthenium red on the energy-linked uptake of Ca 2+ media ted by a synthetic neutral 2 + \" Ca lonophore have been investigated in rat liver mitochondria. The results indicate that unspecific surface charge effects do not play a major role in the mechanism of inhibition of mitochondrial Ca 2+ transport by La 3+ and ruthenium red.\nLa 3+ and ruthenium red are powerful inhibitors of the active transport of Ca 2+ in mitochondria 2,3. The mechanism of the inhibit ion has not yet been investigated in detail. It could either be based on a competi t ive interaction with specific Ca 2+ transport sites on the mitochondrial membrane, or on non-specific charge effects at the inner membrane surface. Indeed, La 3+ is known to change the ion permselectivity of artificial lipid membranes from a cation to an anion exchanger 4. Effects of this type have been described also for di- and trivalent cations (Mg 2+, A13+, etc.), which reduce markedly the rate of val inomycin dependent Rb+-uptake by yeast cells s. A similar effect on mitochondrial Ca ~+ transport has been reported for the positively charged p o l y a m i n e spermine 6. Even if the amounts of polyvalent cations required for a substantial reduction of the rate of uptake is in general rather high, the decrease of the negative surface charge density due to La 3+ or ruthenium red might still conceivably play a role in the inhibition of Ca + transport. This possibilit~r has indeed already been suggested by Scarpa and Azzone ' . To investigate possible surface charge effects of La 3+ and ruthenium red, use was made of a neutral synthetic Ca 2+ ionophore, which selectively transports Ca 2+ in bulk phases in electrodialysis experiments 8. Recently, it has been shown that this Ca 2+ ionophore can transport Ca 2+ across mitochondrial and other biological membranes 9. The experiments have provided conclusive evidence that surface charge effects do not play a major role in the mechanism of inhibit ion of the energy-linked Ca 2+ transport by La 3+ and ruthenium red. Materials and methods: Liver mitochondria were prepared by a conventional method from rats fasted for 12 h 1~ The isolation medium contained 220 m M mannitol, 70 m M sucrose, 10 m M HEPES-Tris , pH 7.4 and 0.5 mg BSA/ml . The Ca + accumulat ion in the presence of increasing amounts o f ruthenium red or La 3+ was measured by using 45Ca2+ and mill ipore filtration. The initial rates of Ca 2+ uptake in the presence o f increasing amounts of ru thenium red or La 3+ were determined using 45Ca2+ and an inhibitor stop-method. The amount of endogenous Ca 2+ released by mitochondria in the presence of rotenone was determined by using the Ca2+-indicator Arsenazo III ll. Ruthenium red was obtained from Fluka and used without further purification. The neutral synthetic ionophore was a gift f rom Prof. W. Simon, ETH Zurich. Results and discussion. Table 1 shows the accumulation o f Ca 2+ by energized rat l iver mitochondria, in the presence of the synthetic neutral Ca 2+ ionophore, and of increasing amounts of La 3+ or ruthenium red. The Ca 2+ bound to mitochondria in the absence o f ionophore was between 2 and 4 nmoles Ca 2+/mg protein. This amount was subtracted from the values reported in tables 1 and 2. The accumulation was not influenced by concentrations of La 3+ in the range of 1-50 gM, or o f ruthenium red in the range o f 0.8- 10 ~tM. The average amount of Ca 2+ accumulated via the ionophore was the same with both inhibitors. In table 2 the initial rates o f Ca 2+ uptake in the presence of the Ca2+-ionophore and of increasing amounts of La 3+ or ruthenium red are shown. Clearly, the initial rate of Ca + uptake remained the same in the presence of concentrations o f La 3+ up to 30 gM, or of ruthenium red up to 10 ~tM. The rates of uptake were comparable for the 2 inhibitors. It must be stressed that these concentrations o f inhibitors are 10-50 times higher than those required to block almost completely the energy-linked uptake. Only when La 3+ was added in concentrations in excess Of 50 gM,\nTable 1. Maximum levels of Ca 2+ accumulated by rat liver mitochondria in the presence of the neutral synthetic ionophore, and of varying amounts of LaCI3 or ruthenium red\nInhibitor Ca 2+ uptake after 2 min (nmoles/mg protein)\nTable 2. Initial rates of Ca 2+ uptake by energized rat liver mitochondria, in the presence of the neutral synthetic ionophore, and of varying amounts of LaC13 or ruthenium red\nInhibitor Initial rate of uptake (nmoles Ca2+/mg protein min)\n2 gM La 3+ 10 5gM La 3+ 12 5 gM La 3+ 12 10gM La 3+ 8.8 10 gM La 3+ 12 20gM La 3+ 10 40 gM La 3+ 7 301aM La 3+ 10 50 I.tM La 3+ 8 35 gM La 3+ 3 60 gM La 3+ 6 70 gM La 3+ 1 2gM RR 8.6\n3 gM RR 8.3 0.8 gM RR 7 5 gM RR 8.1 1 gM RR 10 6gM RR 7.4 4 IxM RR 8 7 gM RR 5.3 5 g M R R 7.8 8 g M R R 11.2\n10 gM RR 6.6 10gM RR 9\n1 mg of rat liver mitochondrial protein were incubated for 8 min 45 2 + at 25 ~ with 2 gM rotenone (final vol. 1 ml). Then, 20 gM Ca and different amounts of La 3+ or ruthenium red were added. After these additions, mitochondria were incubated for 2 min with 40 gM of the synthetic neutral Ca 2+ ionophore. Ca 2+ uptake was then started by the addition of 2 mM Tris-succinate. The Ca 2+ bound by mitochondria in the absence of ionophore was determined for each La 3+ or ruthenium red concentration, and subtracted from the value measured in the presence of ionophore. Ca 2+ uptake was Started with Tris-succinate under the same conditions as described for the Ca 2+ accumulation measurements (table 1), and terminated 30 sec later by the addition of 20 mM KC1. The latter binds to the synthetic ionophore, and completely inhibits the further uptake of Ca 2+. The Ca 2+ bound by mitochondria in the absence of ionophore was determined for each La 3+ or ruthenium red concentration and subtracted from the value measured in the presence of ionophore.", "15.9.78 Specialia 1137\nthe initial rate of Ca 2\u00a7 uptake decreased markedly. In the absence of ionophore and inhibitors of the natural Ca 2\u00a7 cartier, the mitochondria accumulated about 40 nmoles Ca 2\u00a7 protein/min with an initial rate of approximately 100 nmoles Ca 2+/mg protein min.\nFrom these data, it is clear that La 3+, up to a concentration of 40 I.tM, or ruthenium red, up to a concentration of 10 gM, do not inhibit the energy-linked uptake of Ca 2+ by mitochondria mediated by a neutral Ca ~+ ionophore. If non-specific surface charge effects, induced by the interaction of the inhibitors with the mitochondtial membrane, were important, one would have expected the ionophoremediated uptake to be inhibited at the concentrations of La 3+ and ruthenium red that block the natural uptake process. The 2 inhibitors, therefore, most likely interact with a specific site on the natural Ca 2+ transport system of the membrane. At higher concentrations, La 3+ inhibits also the ionophore-mediated Ca 2+ transport. This effect is probably due to unspecific charge effects at the membrane surface (or, possibly, also to competitive binding of La 3+ to the ionophore), but it is most likely unrelated to the 'normal' inhibitory effect of the cation.\nAbbreviations: BSA, bovine serum albumine; RR, ruthenium red; HEPES, N-(2-hydroxyethyl)piperazine-N'-2-ethanesulfonic acid.\n1 Acknowledgments. The authors are indebted to Prof. W. Simon, ETH Zurich, for having provided samples of the synthetic neutral Ca 2+ ligand, and to M. Mattenberger for the valuable technical assistence. The work was supported by a grant of the Swiss Nationalfonds (grant No. 3.1720.75). 2 L. Mela, Archs Biochem. Biophys. 123, 286 (1968). 3 K.C. Reed and F. L. Bygrave, Biochem. J. 140, 143 (1974). 4 C. van Breemen, Biochem. biophys. Res. Commun. 32, 977\n(1968). 5 A.P.R. Theuvenet and G.W.F.H. Borst-Pauwels, Biochim.\nbiophys. Acta 426, 745 (1976). 6 K. Akerman, J. ofBioenerg. Biomemb. 9, 65 (1977). 7 A. Scarpa and G.F. Azzone, Eur. J. Biochem. 12, 328 (1970). 8 W.E. Morf, P. Wuhrmann and W. Simon, Analyt. Chem. 48,\n103I (1976). 9 P. Caroni, P. Gazzotti, P. Vuilleumier, W. Simon and E. Cara-\nfoil, Biochim. biophys. Acta 470, 437 (1977). 10 W.C. Schneider, in: Manometric Techniques, p. 188. Ed.\nW.W. Umbreit, R. Burris and J.F. Stauffer. Burgess, Minneapolis, Minnesota, 1957. l l R. DiPoto, J. Requena, J. Brinley, L.J. Mullins, A. Scarpa and T. Tifferet, J. gen. Physiol. 67, 433 (1976).\nAction of fl-(4-chlorophenyl-GABA) on uptake and metabolism of GABA in different subcellular fractions of rat brain\nM. Tardy, B. Rolland, J. Bardakdjian and P. Gonnard\nSummary. /%(4-chlorophenyl)-GABA, a GABA mimetic compound, acts as an inhibitor of GABA metabolism in both synaptosomal and extrasynaptosomal compartments. It has no significant action on GABA or Glu uptake by synaptosomes.\nfl-p-chlorophenyl-7-aminobutyric acid (fl-p-CPG, Lioresal) has shown considerable potential in the control of spasticity, and has been extensively studied from a pharmacological point of view 1-7. The drug is structurally related to the central inhibitory transmitter 7-aminobutyric acid (GABA) and is apparently able to penetrate the blood-brain barrier on systemic administration. We undertook a study of the biochemical aspect of fl-pCPG activity, that is to say, its possible action on the 2 enzymes responsible for the synthesis and degradation of GABA: Glutamate decarboxylase (GAD) and GABA transaminase (GABA-T). fl-p-CPG appears to have some GABA-ergic properties and could possibly inhibit GABAT like amino-oxyacetic acid (AOAA) which enhances brain GABA levels in this way. The action of fl-p-CPG on the enzymes was studied in the 2 sites implicated in GABA metabolism: synaptosomal GAD, a cytosolic enzyme, and extrasynaptosomal and synaptosomal GABA-T, a mitochondrial enzyme. Material and methods. Preparation of synaptosomes. Male Sprague-Dawley rats (180-200 g) were decapitated and synaptosomes were prepared from the mesencephalon by the method of Gray and Whittaker s as modified by Israel and Frachon Mastour 9. Preparation of extrasynaptosomal mitochondria. Extrasynaptosomal mitochondria were prepared by the method of Gray and Whittaker 8 as modified\nI=0 5\ni o ' : ,o ' 3o ' 4'o\nr=lZS// = /\nof Lioresal and a fixed concentration of a-ketoglutarate (25 mM). Velocity is expressed in gM of succinyl semialdehyde formed in 1 h, by 1 ml of enzyme solution." ] }, { "image_filename": "designv11_61_0003110_iemdc.2009.5075253-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003110_iemdc.2009.5075253-Figure1-1.png", "caption": "Fig. 1 The leading side of stator coils in one pole-arc.", "texts": [ " Strategies in [1] that involve different current levels in the remaining phases were not considered due to limitations on the capacity of the inverters. The machine modelled for all the simulations was a 2.2MW, 50Hz, 2200V, 10 pole motor with a full load current rating of 745A and rated torque of 37 kilo-Newton meters. The parameters were taken from the manufacturer\u2019s data sheet and found to model full load values remarkably well. II PARAMETER VARIATION The equivalent circuit parameters supplied by the manufacturer are generally the values for a 2-phase motor with a coil structure as shown in Fig. 1a. The slots in each pole-arc are divided into two equal regions and the coils in those slots connected in series so there are 90 electrical degrees between the 2 phases. For simplicity a single layer winding is assumed with sinusoidal distribution so that slot harmonics can be ignored. In Fig. 1b the same pole-arc has been divided into 4 regions to create a 4-phase motor and the angle between the phases is now 45 deg. In this case the coils are deemed to have the same current rating as before but only half the length so that all the 2-phase equivalent circuit parameters have to be divided by 2 and the applied voltage per phase 501978-1-4244-4252-2/09/$25.00 \u00a92009 IEEE must also be reduced to 50%. An equally valid model would split the coils lengthwise with half the current rating and maintain the same voltage but that is not used in this paper" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003521_iceee.2010.5661361-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003521_iceee.2010.5661361-Figure1-1.png", "caption": "Figure 1. Single spur gear reducer and gear structure", "texts": [ " By using genetic algorithm and genetic toolbox of MATLAB to get optimum solution quickly and accurately, the efficiency and quality of gear design is greatly improved and the production cycle is shorten. II. MATHEMATICAL MODEL FOR OPTIMIZATION DESIGN Optimization design of gear reducer is usually given power P, gear ratio i, input speed n1 and other technical conditions and requirements, to seek a group of design parameters that achieve technical and economic indexes of a reducer optimum. A single spur gears reducer transmission diagram and gear structure is shown in Fig 1. Assume that the gear ratio is known as i, the input power P(kW), active gear speed n1 (r/min), it\u2019s required to figure out various design parameters to make the reducer as light as possible under the condition that the strength and stiffness are guaranteed. Since the size of the gear and axis is to determine the size and quality of reducer assembly, the objective function is the sum of their value, without considering cabinet and bearing volume or quality. According to the fig 1, the sum of the 978-1-4244-7161-4/10/$26.00 \u00a92010 IEEE volume of gears and shafts can be approximately expressed as follows: ( ) )1( 4 2.1 4 )(3.0 4 )( 4 2 1 22 1 222222 j j j j jnjnjwjwjoj l d bddbddbddV \u00d7+\u00d7\u2212+\u00d7\u2212+\u2212= == \u03c0\u03c0\u03c0\u03c0 Where d01=mz1;d02=mz1i;dw1=mz1-8m;dw2=mz1i-8m;dn1=1.6d1; dn2=1.6d2. From above equation, it can be seen that if i have been given beforehand, the volume of gear and shaft is determined only by face width b, the number of pinion teeth z1, modulus m, shaft diameter d1 and d2 and distance l between two bearings, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000460_1.1835358-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000460_1.1835358-Figure2-1.png", "caption": "Fig. 2 Diagram and anthropometry of walking model", "texts": [ " forward velocity of level bipedal walking can be explicitly and simply controlled by the virtual downhill gradient. A simulation of bipedal walking was developed including two components: ~1! forward-dynamic implementation of a pendulum walker @4# capable of either passive downhill walking or active driven locomotion, and ~2! a nonlinear feedback controller to impose active joint torques. The simulation represents a twodimensional knee-less walker including two legs of length L and mass M L , joined by a revolute joint at the point mass of the head-arms-trunk ~HAT!, M H ~Fig. 2!. Leg masses were located at a distance ucm from the hip along a line joining the hip to the point-foot. Anthropomorphic data ~Fig. 2! including segment lengths and mass were based upon a 77 kg adult male @14#. The walker moved along a plane of slope g with respect to horizontal. The walker configuration was represented as a time-dependent vector u5@uS ,uN#T, where uS is the angle of the stance-leg with respect to ground and uN is the angle of the nonstance-leg with respect to the stance-leg. Ground clearance of the swing-leg is ignored in this treatment because simple mechanisms such as prismatic joints are readily established to facilitate swing foot clearance that does not influence walker dynamics @4#" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000443_s00170-005-0009-x-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000443_s00170-005-0009-x-Figure2-1.png", "caption": "Fig. 2 The local coordinates of spatial parallel mechanism with 4-PUU", "texts": [ "1 The forward and inverse displacement of a kind of 4-PUU parallel manipulator A spatial parallel manipulator, shown in Fig. 1, is made up of 4-PUU (1 prismatic joint and 2 universal joints) kinematic chains. The absolute coordinate system oxyz are created as Fig. 1 shows, where z-axis is perpendicular to the guide plane P1P2P3P4, the origin is on the midline of the two guides, x is superposed with the midline of the two guides and y-axis is perpendicular to the two guides. According to the method presented above, the local coordinate system oc xc yc zc is shown in Fig. 2, where zc-axis is perpendicular to the plane of the manipulator M1M2M3M4, the origin is superposed with the geometric center of M1M2M3M4, xc and yc axes are parallel to the two orthogonal sides of the manipulator. Firstly, we will analyze the DoF of the manipulator as Fig. 1 shows. According to [9], we can find the dimension of the constraints spaces that all of the reciprocal screws, shown in Fig. 3, can be spanned is: d \u00bc dim span $rB1P1M1 $rB2P2M2 $rB3P3M3 $rB4P4M4 8>< >: 9>= >; \u00bc 2: (11) Therefore, F \u00bc 6 d \u00bc 6 2 \u00bc 4: (12) So, the manipulator shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002926_pime_proc_1970_185_113_02-Figure23-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002926_pime_proc_1970_185_113_02-Figure23-1.png", "caption": "Fig. 23. Relation between C and F (cf. Figs18 and 19). Not to scale", "texts": [ " During the design of the circulating power machine my colleagues and I used the actual curve reproduced in Fig. 18 as our basis of slip versus torque and it seems to have been satisfactory in the sense that the drum steps are in the right range. Could the author say whether the curve is linear with r or linear with 111, or linear with log r ; or is it non-linear, that is, a curve? Again the discussion above leading to equations (1) and (2) leads one to ask whether the driving and braking coefficients of Fig. 19 are in terms of C or in terms of F, as the zeros are not the same (see Fig. 23) nor are C and F Vol 185 74/71 at UNIV OF VIRGINIA on June 5, 2016pme.sagepub.comDownloaded from ENGINEERING ASPECTS OF TYRE TESTING D371 Proc lnstn Mech Engrs 1970-71 Vol 185 74/71 at UNIV OF VIRGINIA on June 5, 2016pme.sagepub.comDownloaded from D372 DISCUSSION ON J. A. TURLEY proportional one to the other; CIF depends on all five quantities, h, r, a, D and W. This can be seen from C (Dh- Wa) (DW\u2019-a/h) F - (Or- Wa)\u2019 = iD/W-a/P which can be deduced from equations (1) and (2), and like those equations, being deduced from the energy and moment equations without further assumption, it must apply in all cases, as it is basic mechanics", " 24 is based on fundamental theory, and so the values derived from it contain no empirical content other than that contained in the values used in entering the coefficients rlh, alh, and Dl W to read off the value of CIFh, which is, of course, \u2018effective moment arm\u2019/axle height. In any given case, the D/W, a/h point on the rectilinear chart lies on the same radial line through 0, as the intersection of the curve of ClFh and the appropriate circular arc \u20acor the relevant r/h value (for example, the heavy line ~t, y). Note that r for driving is less than r,, for roIIing and r for braking is greater than ro : also r,/h is empirically found to be greater than unity. Fig. 24 supports the comments on Fig. 23; it also explains the region between C = 0 and F = 0; ground force and axle torque have opposite signs, and share the work of overcoming tyre losses. (14) GOUGH, V. E. and WHITEHALL, S. G. \u2018Universal tyre test machine\u2019, Proc. 9th FISZTA Conf. 1962, 117 (Instn Mech. Engrs, London). V. Novopolsky Moscow The paper gives a detailed description of a machine with circulating power, enabling various tests of passenger car tyres in respect of the rolling motion of driven, driving or braking wheel to be carried out" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003046_1.3650515-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003046_1.3650515-Figure4-1.png", "caption": "Fig. 4 Sector stepped pad configuration", "texts": [ " To illustrate how the load and load-to-friction ratio vary as a function of the sector ratio, the load, friction, and load-to-friction ratios are shown in Fig. 3.8 It can be seen that the optimum step ratio is 0.29 if load is the only consideration, and 0.43 if minimum friction coefficient is required. B. Bearing Characteristics Based on Ausman's Solution As stated previously, a considerable part of this report is based on Ausman's analysis [2], The configuration analyzed and nomenclature used are shown in Fig. 4. As in all gas bearings, the dimensionless load can be expressed as a function of the dimensionless bearing number and several parameters describing the bearing geometry. For a sector stepped pad bearing, these parameters are: Load parameter L = W vA 8 This and all subsequent figures apply to gas lubricants. 2 1 4 / M A R C H 1 9 6 5 Transactions of the A S M E Downloaded From: http://fluidsengineering.asmedigitalcollection.asme.org/ on 01/19/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use Bearing number A = 6/uo)R Pah2 Radius ratio b = r/R Step ratio 77 = r / 7 Number of sectors n Compression ratio q = A + h 1 + ", "420 40 M 80 100 Mod i f i ed Bearing Number A - \u2014 P a h 2 The pocket bearing always has higher load capacity than the load determined by the Ausman analysis, which does not include the pocket. The agreement between the computer solution and the Ausman analysis for the plain bearings is fairly good, particularly since the computer solutions do not apply to bearings of optimum geometry. The Design of Self-Acting Stepped Thrust Pads thrust bearings is developed. This section of the report can be used independently as a design handbook. A stepped thrust pad and the nomenclature used to define the geometry are shown in Fig. 4. In addition to the geometry, the gas viscosity p, the ambient pressure p\u201e, and the bearing angular velocity to are needed to define the bearing. For the design of a thrust pad, it will be assumed that the bearing inner and outer radii, speed, ambient pressure, gas viscosity, and minimum gap are known, and that it is desired to obtain a load of minimum gap curve and optimum geometry (step 1 Use the bearing power dissipation formula, equation (6), and assume that t] = 0 and ijy = 2ir, as a first approximation" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002650_14644193jmbd202-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002650_14644193jmbd202-Figure8-1.png", "caption": "Fig. 8 Mappings from stick switching planes: (a) absolute motion and (b) relative motion", "texts": [ " From the stick switching planes, the mappings are defined as P1 : (i) 21 \u2212\u2192 (i) 12 , P2 : (i) 12 \u2212\u2192 (i) 21 , P3 : (i) 12 \u2212\u2192 (i) 23 P4 : (i) 23 \u2212\u2192 (i) 32 , P5 : (i) 23 \u2212\u2192 (i) 32 , P6 : (i) 32 \u2212\u2192 (i) 21 } (69) JMBD202 \u00a9 IMechE 2009 Proc. IMechE Vol. 223 Part K: J. Multi-body Dynamics at UNIVERSITE LAVAL on June 29, 2015pik.sagepub.comDownloaded from The corresponding switching planes for the mappings relative to impacting chatter with and without stick can be treated as the same. Accordingly, the mappings in equation (69) are the same as in equations (65) and (66) except for the addition of P1 and P4. The mappings related to the stick switching planes are sketched in Fig. 8. In Fig. 8(a), the two stick mappings P1 and P4 are new, and the other four mappings are the same as in Fig. 7. Similarly, mappings based on the relative switching planes can be defined, and within the relative frame, the switching planes are points as described in Fig. 8(b). With mixed switching planes for chatter with and without stick, six mappings are defined by P2 : (i) 12 \u2212\u2192 R (i) 2\u221e, P2 : R (i) 2\u221e \u2212\u2192 (i) 21 P3 : (i) 12 \u2212\u2192 L (i) 2\u221e, P3 : R (i) 2\u221e \u2212\u2192 (i) 23 P5 : (i) 23 \u2212\u2192 L (i) 2\u221e, P5 : L (i) 2\u221e \u2212\u2192 (i) 32 P6 : (i) 32 \u2212\u2192 R (i) 2\u221e, P6 : L (i) 2\u221e \u2212\u2192 (i) 21 \u23ab\u23aa\u23aa\u23aa\u23aa\u23aa\u23ac \u23aa\u23aa\u23aa\u23aa\u23aa\u23ad (70) The stick mapping is difficult to illustrate, but the possible mappings based on the stick and impact switching planes are presented in Figs 9(a) and (b). For mappings in the absolute and relative frames, set the vectors of the switching points as yk \u2261 (tk , x(i) k , x\u0307(i) k , x\u0307(i\u0304) k )T wk \u2261 (tk , z\u0307(i) k , x(i\u0304) k , x\u0307(i\u0304) k )T \u23ab\u23ac \u23ad (71) For impacting maps P\u03c3 (\u03c3 = 2, 3, 5, 6) in the absolute coordinate, yk+1 = P\u03c3 yk can be expressed by P\u03c3 : (tk , x(i) k , x\u0307(i) k , x\u0307(i\u0304) k ) \u2212\u2192 (tk+1, x(i) k+1, x\u0307(i) k+1, x\u0307(i\u0304) k+1) (72) From Appendix 2, the absolute displacement and velocity for two gear oscillators can be obtained with initial conditions (tk , x(i) k , x\u0307(i) k ) and (tk , x(i\u0304) k , x\u0307(i\u0304) k )" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001970_j.colsurfb.2007.06.021-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001970_j.colsurfb.2007.06.021-Figure6-1.png", "caption": "Fig. 6. The cyclic voltammograms on the GC electrode modified with 150 mol/g film of PE-C5-NQ/DPPC membrane as a function of pH at 42 \u25e6C. Scan rate: 0.1 V s\u22121.", "texts": [], "surrounding_texts": [ "110 Y. Suemori et al. / Colloids and Surfaces\nFig. 4. The cyclic voltammograms on the GC electrode modified with 1 a a\nt fl m b w D o c r l D s C t\nF b\nq t N b c o t o t f T c c t s t b a r e s t e g [\n3\ns F r\n50 mol/g film of PE-C5-NQ/DPPC membrane as a function of the temperture: background (dashed line), 40 \u25e6C (broken line), 42 \u25e6C (solid line) in an queous B.R. buffer, pH 7.0. Scan rate: 0.1 V s\u22121.\nhat a set of waves is clearly visible at 42 \u25e6C (above the geluid phase transition temperature, Tm, of DPPC; 41.5 \u25e6C) as entioned above. At 40 \u25e6C (below the Tm), the redox response ecame small. Fig. 5 shows cathodic peak currents of PE-C5-NQ ith various phospholipid membranes, DMPC (Tm: 23.5 \u25e6C), PPC (Tm: 41.5 \u25e6C) and DSPC (Tm: 55.5 \u25e6C) [34], as a function f temperatures. The peak currents of the PE-C5-NQ indiate a sharp break point in the current\u2013temperature behavior, eflecting the gel-to-fluid phase transition of the multilamelar membrane. The sharp break point for DMPC, DPPC and SPC were observed at 23, 42 and 52 \u25e6C, respectively. This uggests that the extent of free movement and diffusion of PE5-NQ in the membrane may be a crucial factor for the electron ransfer from the electrode to PE-C5-NQ molecules and subse-\nig. 5. Plots of the cathodic peak current of PE-C5-NQ in phospolipid memranes of DMPC ( ), DPPC ( ) and DSPC ( ) as a function of temperature.\nT w q [ c p e p Q 4 t P p t q v i m t\n3\nc e\nB: Biointerfaces 61 (2008) 106\u2013112\nuent self-exchange reaction via quinone/hydroquinone. Similar emperature-dependent CV responses were observed for PE-C0Q and PE-C11-NQ (data not shown). The temperatures at the reak point are near the Tms of the phospholipid used. It is onsidered that below the Tm the lipids are arranged in tiled ne-dimensional lattices and then at pretransition temperature wo-dimensional arrangements of the lipid are formed with peridic undulations. Above the main phase transitions lipids revert o one-dimensional lattice arrangements, separated somewhat rom each other, and assume mobile liquid-like conformations. hus, near the main phase transition quinones become to assoiate with this structural transformation of lipid bilayers and an catalyze electron transfer via self-exchange reaction. Below he phase transition of the lipid bilayer membrane all quinones howed little or no electron transfer activity, where electron ransfer from this quinone to a hydroquinone cannot occur ecause the lipids are arranged in tiled one-dimensional lattice nd then the quinones are frozen in the lipid bilayers. These esults again demonstrate that the phospholipid-linked quinones xhibit electron transfer in the lipid membranes, depending on tructural change of the lipid bilayer such as the main phase ransition properties of the lipid bilayer. Such phase-dependent lectron transfer is consistent with that observed for mananese porphyrin derivative as described in the previous papers 1\u20134].\n.5. Effect of the pH on the redox response of PE-Cn-NQ\nThe redox behavior of quinone derivatives on the electrode trongly depends on the pH condition of the buffer solution. ig. 6 shows the cyclic voltammograms measured in the pH ange of 3\u201312 for the PE-C5-NQ/DPPC-covered GC electrode. he shape of the cyclic voltammograms significantly changed ith the pH value. This indicates that reaction sequences of uinone/hydroquinone were different with the pH condition 35]. Fig. 7 shows the pH dependence of the anodic (Epa), athodic (Epc) peak potentials and quinone-hydroquinone redox otentials (E1/2) of the PE-C5-NQ/DPPC membrane on the GC lectrode at 42 \u25e6C. The slope of pH-E1/2 was \u221263 mV between H 3 and pH 11, corresponding to the literature value for the /QH2 reaction [32]. Above pH 11, the slope was \u221231 mV at 2 \u25e6C, corresponding to Q/QH\u2212 of two-electron and one-proton ransfer [36]. Below pH 6, the first step of the reduction of E-C5-NQ was a monoprotonation of quinone, whereas above H 10 the first step of the oxidation was a monodeprotonaion of hydroquinone. This indicates that reaction sequences of uinone/hydroquinone are different with the change of the pH alues. These results show that the quinone derivatives exhibted electron transfer associated with proton transfer in the lipid\nembranes, depending on the pH and the structural change on he multilayer assembly of the lipid bilayers.\n.6. Effect of the scan-rate dependence\nFig. 8 shows the effect of the scan-rate dependence on the athodic peak current for the reduction of PE-C5-NQ on the GC lectrode modified with DPPC membrane at 42 \u25e6C. The cathodic", "Y. Suemori et al. / Colloids and Surfaces B: Biointerfaces 61 (2008) 106\u2013112 111\nF ( p\np w d w t t o\n3 ( w\nC i p o [ t t m a t c t b t a a t f p d t F\nig. 8. Plots of the cathodic peak current of PE-Cn-NQ/DPPC membranes 150 mol/g film) as a function of the scan rate in an aqueous B.R. buffer, H 7.0 at 42 \u25e6C.\neak currents for the reduction of PE-C5-NQ increased linearly ith the square root of the scan rate. Similar scan-rate depenency was observed for PE-C0-NQ and PE-C11-NQ and under ith difference pH conditions. These facts indicate that the elec-\nrochemistry of phospholipid-linked quinone derivatives inside he lipid membrane is under conditions under which diffusion f charge occurs during the voltammetric scan.\n.7. Effect of various spacer methylene groups Cn):comparison of quinone-mediated electron transfer ith porphyrin-mediated one in the lipid membrane\nAs shown in Fig. 8, the effect of the methylene spacer lengths, n (n = 0, 5, 11), of PE-Cn-NQ on the cathodic peak current s negligible. In contrast, phospholipid-linked manganese porhyrins reported previously exhibited the marked dependence f the Cn on the peak currents, in the order of C5 > C11 > C0 4]. The observed Cn dependence indicated that the length of he spacer methylene group is an important factor for conrolling the electron transfer between Mn(III) and Mn(II) of anganese porphyrin moiety. For the manganese porphyrin as n electron mediator, n = 5 is probably suitable for the elecron transfer on the electrode because the manganese moieties an freely move and approach each other enough for elecron transfer both between electrode and porphyrin and also etween porphyrins (Fig. 9a). Electron transfer mediated by he PE-Cn-NQ, that is diffusion-controlled process, is associted with proton transfer in the lipid membrane as mentioned bove. This may be the reason for the Cn-independent elecron transfer of PE-Cn-NQ. In fact, the observed peak currents or the PE-C -NQ were 10\u201320 times larger than that for the\nn\nhospholipid-linked manganese porphyrins. Taken together, the iffusion-controlled electron transfer process for PE-Cn-NQ, hat is independent from the spacer length, can be depicted in ig. 9b.", "112 Y. Suemori et al. / Colloids and Surfaces\nFig. 9. Schematic representation of electron transfer mediated by phospholipidlinked manganese porphyrins (a) and NQs (b) incorporated in DPPC lipid membranes cast on the GC electrode. The single lipid bilayers on the elect a\n4\nC q b m o c p a l m f m l t p o t c c i t b o s s\nA\nt S\nR\n[\n[ [ [\n[\n[ [\n[\n[ [ [ [ [\n[ [ [\n[ [\n[\n[ [\n[\n[\nrode were depicted as an interface layer between the multilamellar membrane nd the electrode.\n. Conclusions\nThe phospholipid-linked naphthoquinone derivatives, PEn-NQ (n = 0, 5, 11), were synthesized to investigate the uinone-mediated electron transfer in the phospholipid memranes. Cyclic voltammetry on the ITO and the GC electrodes odified with the PE-Cn-NQ/DPPC membrane showed that\nne set of waves was clearly observed under the neutral pH ondition, indicating the consecutive two-electron and tworoton transfer of the quinone unit (quinone/hydroquinone) only bove the phase temperature of the membrane. Phospholipidinked quinones in liposomal membrane and on the electrode\nodified with lipid bilayer membrane caused the electron transer, depending on the pH and the structural change on the ultilayer assembly of the lipid bilayers. Interestingly, the arge peak currents and no or less dependency of the Cn on he behavior of the electron transfer was observed for the hospholipid-linked quinone in comparison to the behavior f phospholipid (PE)-linked porphyrin derivatives, suggesting hat electron transfer associated with proton transfer plays an rucial role in the quinone-mediated electron transfer. These haracteristics in liposomal membrane and on electrode modfied with lipid bilayers are of considerable interest to mimic he vectorial electron transfer controlled in biological memranes and to construct an efficient electron transfer systems n electrode. Thus, these quinone complexes can be utilized to ystematically examine electron transfer in these lipid bilayer ystems.\n[ [ [ [\nB: Biointerfaces 61 (2008) 106\u2013112\ncknowledgement\nThis study was partially supported by Grant-in-Aid for Scienific Research from the Ministry of Education, Culture, Sports, cience and Technology of Japanese Government.\neferences\n[1] K. Iida, M. Nango, M. Matsuura, M. Yamaguchi, K. Sato, K. Tanaka, K. Akimoto, K. Yamashita, K. Tsuda, Y. Kurono, Langmuir 12 (1996) 450. [2] M. Nango, K. Iida, M. Yamaguchi, K. Yamashita, K. Tsuda, A. Mizusawa, T. Miyake, A. Masuda, J. Yoshinaga, Langmuir 12 (1996) 1981. [3] T. Ohtsuka, T. Hikita, M. Nango, J. Electroanal. Chem. 438 (1997) 105. [4] M. Nango, T. Hikita, T. Nakano, T. Yamada, M. Nagata, Y. Kurono, T.\nOhtsuka, Langmuir 14 (1998) 407. [5] T. Ohthuka, M. Nagata, H. Komori, M. Nango, Electrochemistry 67 (1999)\n1184. [6] T. Yamada, T. Hashimoto, S. Kikushima, T. Ohtsuka, M. Nango, Langmuir\n17 (2001) 4634. [7] T. Yamada, M. Nango, T. Ohtsuka, J. Electroanal. Chem. 528 (2002) 93. [8] T. Yamada, S. Kikushima, T. Hikita, S. Yabuki, M. Nagata, R. Umemura,\nM. Kondo, T. Ohtsuka, M. Nango, Thin Solid Film 474 (2005) 310. [9] R.K. Clayton, W.R. Sistrom, The Photosynthetic Bacteria, Elsevier/North-\nHolland Biochemical Press, New York, 1978. 10] J. Barber, Topics in Photosynthesis, Elsevier/North-Holland Biochemical\nPress, New York, 1979. 11] Govinjee, Photosynthesis, Academic Press, New York, 1982. 12] H. Sheer, Chlorophylls, CRC Press, Florida, 1991. 13] M. Nango, T. Dannhauser, D. Huang, K. Spears, L. Morrison, P.A. Loach,\nMacromolecules 17 (1984) 1898. 14] T. Dannhauser, M. Nango, N. Oku, K. Anzai, P. Loach, J. Am. Chem. Soc.\n108 (1986) 5865. 15] M.R. Wasielewski, Chem. Rev. 92 (1992) 435. 16] M.A. Fox, M. Chanon, Photoinduced Electron Transfer, Elsevier, Amster-\ndam, 1988. 17] Y. Sakata, S. Nakashima, Y. Goto, H. Tatemitsu, S. Misumi, T. Asahi, M.\nHagihara, S. Nishikawa, T. Okada, N. Mataga, J. Am. Chem. Soc. 111 (1989) 8979. 18] M.A. Cusanovich, Photochem. Photobiol. 53 (1991) 845. 19] J.A. Runquist, P.A. Loach, Biochim. Biophys. Acta 637 (1981) 231. 20] J.T. Groves, G.D. Fate, S. Lahiri, J. Am. Chem. Soc. 116 (1994) 5477. 21] M.T. Rojas, M. Han, A.E. Kaifer, Langmuir 8 (1992) 1627. 22] M. Tominaga, J. Yanagimoto, A.-E.F. Nassar, J.F. Rusling, N. Nakashima,\nChem. Lett. (1996) 523. 23] J. Zak, H. Yuan, M. Ho, L.K. Woo, M.D. Porter, Langmuir 9 (1993) 2772. 24] V. Heleg-Shabtai, E. Katz, I. Willner, J. Am. Chem. Soc. 119 (1997) 8121. 25] Y. Xiao, F. Patolsky, E. Katz, J.F. Hainfeld, I. Willner, Science 299 (2003)\n1877. 26] E. Katz, N. Itzhak, I. Willner, Langmuir 9 (1993) 1392. 27] J.F. Smalley, S.W. Feldberg, C.E.D. Chidsey, M.R. Linford, M.D. Newton,\nY.P. Liu, J. Phys. Chem. 99 (1995) 13141. 28] F. Mukae, H. Takemura, K. Takehara, Bull. Chem. Soc. Jpn. 69 (1996)\n2461. 29] L. Zhang, T. Lu, W.G. George, E.K. Angel, Langmuir 9 (1993) 786. 30] R.V. Duevel, R.M. Corn, M.D. Liu, C.R. Leidner, J. Phys. Chem. 96 (1992)\n468. 31] L. Zhu, R.F. Khairutdinov, J.L. Cape, J.K. Hurst, J. Am. Chem. Soc. 825\n(2006) 825. 32] D. Marchal, W. Boireau, J.M. Laval, J. Moiroux, C. Bourdillon, Biophys.\nJ. 72 (1997) 2679. 33] M. Nagata, Ph.D. Thesis, Hokkaido University, 2003. 34] G. Cevc, Phospholipids Handbook, Marcel Dekker Inc., New York, 1993. 35] E. Laviron, J. Electroanal. Chem. 164 (1984) 213. 36] E. Laviron, J. Electroanal. Chem. 146 (1983) 15." ] }, { "image_filename": "designv11_61_0002026_icma.2007.4303538-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002026_icma.2007.4303538-Figure4-1.png", "caption": "Figure 4. Position of potentiometer", "texts": [ " 4100 children are using all over the world and more than 180 ones in Japan. It consists of knee-ankle-foot orthosis and carriage. The greatest asset of the hart walker is that user\u2019s hands become free and user can keep right posture. Load to the leg is controllable by modifying the position of connecting point between the stem and knee-ankle-foot orthosis. Moreover it is very easy to adjust the length of frames to the body as shown in Fig.3. In order to measure angles for hip and knee joint, we utilize potentiometers described in Fig.4. The McKibben-type actuator consists of an internal bladder surrounded by a braided mesh shell (with flexible yet non-extensible threads) that is attached at either end to fittings. As shown in Fig. 5, when the internal bladder is pressurized, the highly pressurized air pushes against its inner surface and against the external shell, tending to increase its volume. Due to the non-extensibility of the threads in the braided mesh shell, the actuator shortens according to its volume increase and/or produces a load if it is coupled to a mechanical load" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000173_gt2003-38718-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000173_gt2003-38718-Figure5-1.png", "caption": "Fig. 5 Geometric relations for a bristle at an angle.", "texts": [ " [18-20] provide more details for the 3-D model. As focus of the present work is bristles stresses, further discussion about the model itself will be avoided here. Before investigating bristle stresses under pressure-friction coupling, model can be verified through comparison with the nloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?u beam theory calculations for both interference and pressure loading cases in the absence of friction. Using the bristle configuration illustrated in Fig. 5 bristle tip normal deflection, yd, can be expressed in terms of the radial rotor interference, \u2206R, and cant angle, \u03b8, as follows. d R y sin\u03b8 \u2206= , (1) where yd, \u03b8 and \u2206R are as shown in Fig. 5. Beam theory [21] suggests that for a cantilever beam in bending, 3 3d WL y EI = , (2) where W is the normal force acting at the end, L is the length (a+b), E is the elastic modulus, and I is the moment of inertia. Maximum bending stress is given by 2 WLd I \u03c3 = , (3) where d is the beam (or bristle) diameter. Substitution of Eqs. (1) and (2) in Eq. (3) provides the maximum bristle stress as a function of radial rotor interference, \u2206R: 2 3 2int Ed R L Sin \u03c3 \u03b8 \u2206= . (4) When a sample rotor interference of 2.032 mm (0.080 in.) is tested with the parameters listed in Table 1, results from the model show a good match with Eq. (4). When a uniform pressure load (per unit length), w, is applied at the overhanging section b in Fig. 5, beam theory suggests that the moment, M, at the clamped top point is calculated by 2 4 wb M = , (5) from which bending stresses due to the pressure loading can be calculated around the retaining plate as 4 Copyright \u00a9 2003 by ASME rl=/data/conferences/gt2003/71962/ on 03/20/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Down 2 2 8press Md wb d I I \u03c3 = = , (6) where I is the moment of inertia defined by 4 64 d I \u03c0= . (7) A frictionless sample case is modeled using the seal parameters listed in Table 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003949_j.ast.2011.09.010-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003949_j.ast.2011.09.010-Figure4-1.png", "caption": "Fig. 4. Cut-away plot of the fuselage model exposing the RBE3 constructs (black spider webs) used for concentrated mass attachment to the frame/skin.", "texts": [], "surrounding_texts": [ "The fuselage model mesh has been refined at the crown of frame 482 thus providing enough detail to capture the actual loca- tion of strain gauges in this region. Stress results are derived from the whole fuselage model with refined mesh at FS 482. Fig. 6 illustrates the mesh refinements in the region of strain gauge 2. No separate component model is used in the present study. The dynamic fuselage properties used in the aeromechanics model are derived from the fuselage model including this refinement. Frequency results for the refined fuselage model at various aircraft weights are presented in Table 3. Note that frequencies have changed significantly by weight changes and the reduction of hub mass (compare with Table 2). The mode shapes remain very similar." ] }, { "image_filename": "designv11_61_0001875_6.2007-6195-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001875_6.2007-6195-Figure5-1.png", "caption": "Figure 5. Two Sided Qualification Part", "texts": [], "surrounding_texts": [ "David P. Heck1 and Dr. Kevin Slattery2 Boeing Phantom Works, St. Louis, Missouri 63166 Robert Salo3 and Scott Stecker4 Sciaky, Inc, Chicago, Illinois 60638 [Abstract] Boeing Phantom Works has initiated a studies of Electron Beam Deposition of Titanium 6Al-4V for the fabrication of aircraft and spacecraft structure and mechanisms. This technique involves the melting of Ti 6-4 material in a vacuum and solidifying it to form engineered shapes. Boeing has evaluated two deposition methods and materials, namely low power, powder based and higher power, wire based. Each of these systems has applications for aircraft and spacecraft. The purpose of this paper is to report the results of mechanical testing of the higher power, wire based method. This testing was undertaken to qualify this process for the production of flight worthy, production parts. In general, for a given set of welding parameters, this process were found to yield mechanical properties similar to the parent material. This paper will detail these results. I. Background In the mid to late 1990s, Boeing\u2019s Phantom Works Division developed a mechanical property database for the deposition of Titanium 6Al-4V using a large, 18 kilowatt CO2 laser. This process was called Laser Added Manufacturing (LAM). This database was used to create a material specification for this process with the American Material Society, namely AMS 4999. This database was very extensive and was sufficient to create \u201cA\u201d basis allowables for Mil-Handbook-5. Boeing also created three internal specifications for each of its three primary heritage companies (Boeing, McDonnell, and Douglas). This process showed excellent mechanical properties at a substantial cost reduction over forgings and thick plate raw material (in certain cases). Boeing used this process to create \u201cpreforms\u201d for several production parts for several aircraft using this process. It was aimed at depositing large amounts of material in a short period of time, thereby creating a perform at very low cost. This laser system, owned by the US Army Research Lab, was housed at the Aeromet Corporation, in Minneapolis, Minnesota. In September, 2005, the parent company of Aermet, MTS Systems Corporation, closed the Aeromet company. In order to continue production of these low cost, high quality preforms, Boeing turned to the Sciaky Company of Chicago, Illinois and began an in-depth investigation and qualification of Sciaky\u2019s electron beam (e-beam), wire fed deposition process. The results of this investigation and qualification process are included in this report. II. Process Overview The Sciaky Company of Chicago developed the process of adapting their electron beam welding systems into a metal deposition system called Electron Beam Free Form Fabrication (EBF3) in 2004. In late 2005, Boeing initiated a joint project with Sciaky to further develop and qualify the process into one capable of producing flight worthy hardware. The EBF3 process is depicted in Figure 1. The process takes place in a vacuum welding chamber, where the ebeam gun is mechanically moved over the surface of a substrate, such as wrought plate, forging or casting. 1 Associate Technical Fellow, Advanced Manufacturing Research & Development, S245-1003, AIAA Senior Member 2 Senior Manager, Advanced Manufacturing Research & Development, S245-1003 3 Sales Manager, Sciaky, Inc, 4915 West 57th Street 4 Project Engineer, Sciaky, Inc, 4915 West 57th Street AIAA SPACE 2007 Conference & Exposition 18 - 20 September 2007, Long Beach, California AIAA 2007-6195 Copyright \u00a9 2007 by David Heck. Published by the American Institute of Aeronautics and Astronautics, Inc., with permission. American Institute of Aeronautics and Astronautics The e-beam is directly and steered electromagnetically to strike a titanium 6-4 wire, thereby heating it into a molten pool. This molten pool is deposited onto the substrate in successive layers to build an engineered shape. This shape, or preform, will be machined to a finished shape to remove any rough edges left from the deposition process. Figure 2 shows an example of a section of a deposited part. III. Process and Supplier Qualification With the change from laser to e-beam melting methods, Boeing has undertaken a revision to AMS 4999 and Boeing\u2019s internal specification to allow suppliers and OEMs to create qualified, flight worthy performs and finished parts using a variety of metal deposition systems. This will allow the qualification of systems powered by lasers, electron beams, TIG and other sources, along with powder based, round wire and flat wire metal sources. This revision will allow deposition system operators to prove their individual systems as being equivalent to the original LAM database without creating an extensive and expensive database. For any particular system, the specification requires that the supplier to create a qualification matrix similar to that shown in Figure 3 below. American Institute of Aeronautics and Astronautics The results of these tests must meet several requirements. These requirements are a set of minimum tensile strength values, a set of mean tensile strength values, a maximum coefficient of variation of the entire data set of tensile data, a minimum fatigue life for a given load level, a minimum fracture toughness value and chemistry tests to verify the composition of the final article. These requirements will be shown in the following section. To assist in this qualification process, Boeing has developed 3 demonstration part geometries. These parts contain geometric features typical of aerospace parts, such as curved flanges, multiple intersecting stiffeners and varying heights. Figures 3 and 4 show these three part geometries. The part shown in Figure 4 is a two sided part, with depositions on both sides of the substrate plate. American Institute of Aeronautics and Astronautics These part geometries also contain sufficient volume to extract numerous coupons for mechanical testing. CAD models of these parts are supplied to the companies, which also allows them to compare the geometric accuracy of their deposited surfaces to the engineering model. IV. Process Overview At the onset of this qualification effort, Sciaky undertook the Stage 1 testing as a means of prequalifying their system. After verifying Sciaky\u2019s results, Boeing then contracted Sciaky to create the 3 demonstration parts shown in Figures 4 and 5. During the testing of the Sciaky EBF3 system, Boeing tested 162 tensile, 24 fatigue, and 6 fracture toughness coupons from the three large qualification parts. The tensile and fatigue coupons were aligned in specific directions relative to the deposition, such as parallel to the direction of deposition (\u201cX\u201d direction), perpendicular to the direction of deposition (\u201cY\u201d direction) and in the direction of the build height (substrate thickness direction). These directions can also be seen from the illustration in Figure 1. In the following sets of figures, summaries of each orientation of static test specimens will be presented. These summaries will show the criteria for each of three sets of data, namely the mean ultimate and yield tensile strength, the coefficient of variation, and the minimum ultimate, yield tensile strength plus minimum elongation. Summaries are being shown in lieu of the raw data due to the large number of tensile specimens. These figures also show the acceptance criteria as defined in the pending revision to AMS 4999.. American Institute of Aeronautics and Astronautics In reviewing this data, it is clear that the mean ultimate and yield strength for both the \u201cX\u201d and \u201cY\u201d direction specimens were below the required mean values. The mean \u201cX\u201d direction values were significantly below the required mean, while the \u201cY\u201d direction values were only slightly below the criteria. The next set of data to be reviewed is the fatigue results. Figure 9 shows the results from the testing as well as the criteria. As can be seen from the figure, the test coupons had results well above the minimum requirement. American Institute of Aeronautics and Astronautics The final set of mechanical tests is the fracture toughness specimens. Again, the test results were well above the minimum criteria, as can be seen from Figure 10. The final test performed was a test of the chemistry of the deposited material. The results of this test are shown in Figure 11. The figures shown are percentages of aluminum and vanadium content of the total material tested. The specification minimum is also listed as a percentage of the whole. American Institute of Aeronautics and Astronautics As can be seen, the aluminum values of two of the test parts were below specification minimums. The implications of these values and their relationship to the tensile results will be discussed in the following section. V. Conclusions and Summary In reviewing all the data, the results of the tests of the tensile strengths of the \u201cZ\u201d direction specimens, the fatigue specimens and the fracture toughness specimens were all above the specification minimums. The tensile strength in the \u201cX\u201d and \u201cY\u201d direction specimens, however, were below the minimum values required for qualifying this process. Also, there were insufficient amounts of aluminum in the chemistry tests of the deposited material. Several in depth discussions have taken place between Sciaky and Boeing personnel regarding the implications of these results, specifically regarding the loss of aluminum. This loss of aluminum appears to be a result of the reduced vapor pressure of aluminum in a vacuum and has been experienced previously in electron beam welding sessions. The loss of aluminum is theorized to be the reason for the reduced tensile strength in the \u201cX\u201d and \u201cY\u201d directions, and is thought to have occurred during times when the temperature of deposition melt pool was at an elevated level for an extended period of time. Boeing and Sciaky, along with other companies and organizations, have undertaken a study of modifying the parameters of the process, such as electron beam strength, beam dwell time, etc. to reduce the amount of aluminum being burned off while still maintaining a high deposition rate. It is thought that by carefully controlling the energy input into the weld pool and thereby controlling the pool temperature, the amount of aluminum lost can be minimized. The results of this study, while slightly below expectations in some regards, highlights the capabilities of the Electron Beam Free Form Fabrication Process. The beam power can be easily modified by controlling the beam voltage and by rastering and \u201cdefocusing\u201d the beam to spread the power density at the weld pool at any point during the deposition process (at corners, layer by layer, at start and stops, etc.). The fact that there is excess power available also leads to the possibility of decreasing the wire feed rate, thereby achieving even higher deposition rates. Since it is still very early in the development of this process, there is significant room for experimentation to find optimum \u201cfeeds and speeds\u201d for a given set of deposition geometries. This again serves to highlight the broad potential of this process." ] }, { "image_filename": "designv11_61_0000898_cira.2003.1222115-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000898_cira.2003.1222115-Figure8-1.png", "caption": "Figure 8: Geometric meaning for optimal forward kinematics ,map", "texts": [ "To and T*J*), we can associate a covector in T&Ja with a vector in TO,.^. Let dh.; E TtaQOL cT&Jo f o r i = l , . . , , m be TOW components of 2 which can be interpreted as basis vectors (columns) fo r normal vector space Te,Q:. Let f : .7- - E be a forward kinematics map satisfying the constmint floe = fQ . and dfj E TtaJ0 be the row components for $. Then minimizing for all 8. E Q. is equivalent to the condition that the inner products < dh,,,df, >= 0 for all i , j . I t is pictorially shown in Figure 8. 11 can be seen that the constraint f I Q . = f Q a restricts how the forward hnematics map f varies along the tangential directions (Tee.Q.) on Q., i.e., the projection of df on the tangent plane To, Qa is determined and is just dfQa. While the optimal forward kinematics map will have no component on the normal direction T8.Q: ,i.e., i t is insensitive to the unriations in noma1 directions. Remark 3 Uniqueness of the solution It can also be seen that both the constraint (flQ, = f Q a ) and the optimal condition ((dh,,,df,) = 0 f O T all i , j ) only limit and determine the behavior o f f locally on Q. C X. Therefore, even with the additional optimal condition, f is still not uniquely determined over Ja. However, the local behavior of f f o r some neighborhood of any point on Q. is completely and uniquely determined. As it can be seenfrom Figure 8 that for a %dimensional Q. imbedded into a Pdimensional space Ja, let f be the optimal forward kinematics map which has no component in the normal direction Te,Q:; Then, df is given by the linear combination of two bases only, say, df iz and dfz3, i f they are supposed to be linear independent. That is to say as long as only local behavior is concerned, we do not have to use an infinite number of function bases to form the mapping. Since the optimal forward kinematics map problem is very difficult to solve, here we propose, for practical implementation, a method to obtain a better forward kinematics map which utilizes the sensor redundancy to minimize the effect due to noise and other uncertainties" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000477_iros.2005.1545236-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000477_iros.2005.1545236-Figure4-1.png", "caption": "Fig. 4 The relationship between FZMP and support polygon", "texts": [ " Otherwise ijE is not VCE and should be ignored. Here, sP , tP is respectively the one of the eight corner points. >\u2212 >\u2212 0 0 ' ' tt ss yy yy or <\u2212 <\u2212 0 0 ' ' tt ss yy yy ).8,...1,( tsts \u2260= (6) Where ' sP and ' tP are the projection points of sP and tP on the ijL , ijsijs bxay +=' , ijtijt bxay +=' . B. The Relationship Between FZMP and Support Polygon We have to determine which is the rotation edge in all VCEs when robot lose stability. and in this case the distance from FZMP to the rotation edge can be calculated. In figure 4 the distance from FZMP to ijE is expressed as follows: ijFZMPnpji pp \u2212= =, mins (7) Where ijp is the position vector of vertical point from FZMP to ijE . 2 2 2 1 1 ij fzmpijfzmpijij P ij fzmpfzmpijijij P a xayab y a xyaba x ij ij \u2212 +\u2212 = \u2212 +\u2212 = (8) We denote the point op as the position vector of the vertical point that satisfies (7). That is ofzmp pp \u2212=s (9) If we know the position of FZMP the rotation edge can be determined according to (7), the distance from FZMP to the rotation edge and the direction of losing stability can be calculated by (9)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000111_00022660310457257-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000111_00022660310457257-Figure2-1.png", "caption": "Figure 2 Force diagram", "texts": [], "surrounding_texts": [ "The ability of an aircraft to fulfil a design requirement needs a number of factors to be satisfied simultaneously. The discussion could take many forms but an example of aircraft combat requirements is presented. A combat aircraft will often need to manoeuvre in flight in order to attack a foe. This is often termed \u201cdog fighting\u201d. One of the most important performance aspects of such an aircraft is its turn rate. This indicates Engineering design education \u2013 the integration of disciplines Simon Newman, David Whatley and Ian Anderson Having produced this lift force, it is imperative that the structure of the aircraft is capable of withstanding these loads. Figure 5 shows a similar plot to Figure 4, only the structural limitation is defined. The limit of structural failure is shown together with the normal never exceed speed limit. The aircraft design must adhere to both of these requirements and so the composite figure is shown in Figure 6. The character of the graphs shows that there is an optimum speed at which the turn rate is maximised. This is known as the corner speed. The final consideration is the requirements of the powerplant. In generating the considerable amount of lift in the turn a high value of drag occurs. In order to drive the aircraft through the turn the engines must be capable of providing the high thrust required. Figure 7 shows the available thrust and is therefore a third limit to be scrutinised. Figure 8 shows the complete dog house plot where all three limitations are included. This shows graphically how the aircraft turn performance varies considerably over the entire flight speed range. It is the interaction of the various diverse engineering disciplines, which influences the design of the aircraft and also its performance and flying characteristics. If an aircraft is flying level at a speed of V, then in order to execute the most effective turn the pilot will bank the aircraft up to the structural limit and then bleed off speed until the sustained turn rate point is achieved. If the speed is bled off even further, then the instantaneous turn rate can be invoked. It is apparent that this example shows how an aircraft design process must address many different factors. It is often the case that these factors are in conflict and decisions have to be made such that the overall vehicle performance matches, or exceeds, the stated requirement." ] }, { "image_filename": "designv11_61_0002336_1.2999933-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002336_1.2999933-Figure2-1.png", "caption": "FIGURE 2. Sketch of the \"single-wave\" regime. Typical view of melt pool (scale: 1mm) and cross-section (scale:0.5 mm)", "texts": [ " Therefore, this regime can be described by a cylindrical keyhole surrounded by a large melt pool and corresponds to what could be described to the first order by a kind of \"Rosenthal\" heat flow regime. from the keyhole surface heated at evaporation temperature. b) Welding Speeds Between 6 to 8 m/min: \"Single Wave\" Regime. This regime is observed for welding speeds ranging from 6 to 8 m/min. It is characterized by the presence of a rather large single swelling generated near the top of the rear keyhole wall (see Fig. 2). It is only from this region that melt droplets are emitted. This large wave is ejected rearwards quite periodically and generates back and forth oscillations of the melt pool therefore leading to closures of the keyhole. The vapor plume that is emitted rather deeply inside the keyhole and that collides with the melt pool triggers these oscillations. Because of the high level of welding speed, the inclination of the keyhole front begins to be important and this vapor plume, emitted perpendicularly from the keyhole front surface, is then rather directed rearwards [4]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000390_05698190500313478-Figure14-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000390_05698190500313478-Figure14-1.png", "caption": "Fig. 14\u2014Project area DYi and DYa.", "texts": [ " The frictional moment about the Y axis due to the slip of an element of area ds between the ball and the polymer lubricant dmbY is given by dmbY = \u00b5p (d/2) \u221a (d/2)2 \u2212 Y2\u221a (d/2)2 \u2212 Y2 \u2212 Z2 dYdZ [10] From Eq. [10], the frictional moment about the Y axis due to the slip between one ball and the polymer lubricant mbY is given by mbY = 2 \u222b \u222b DY dmbY \u2212 (\u222b \u222b DYi dmbY + \u222b \u222b DYa dmbY ) [11] where DY is the projection of the ball surface on the YZ plane and is given by DY = {(Y, Z)| \u2212 \u221a (d/2)2 \u2212 Z2 \u2264 Y \u2264 \u221a (d/2)2 \u2212 Z2, \u2212 d/2 \u2264 Z \u2264 d/2} [12] In Eq. (11), DYi and DYa are the orthographic projections on the YZ plane of the exposure areas at the inner and outer race sides, respectively (as shown in Fig. 14). DYi and DYa are given by DYi = { (Y, Z) \u2223\u2223\u2223 \u2212 ( li cos \u03b2 \u2212 \u221a d2 \u2212 l2 i sin \u03b2 )/ 2 \u2264 Y, ( li cos \u03b2 + \u221a d2 \u2212 l2 i sin \u03b2 )/ 2 \u2265 Y, \u2212si / 2 \u2264 Z\u2264 si / 2 } [13] Within the conditions of this experiment, the raceway control in each test bearing is the \u201cinner raceway control\u201d (Yamamoto (14)). The angle \u03b2 defining the ball rotation axis is given by \u03b2 = \u03b1 + tan\u22121 ( d sin \u03b1 2R ) [15] where R is the distance from the bearing axis to the ball-inner raceway contact point and is given by R = RLi + RRi (1 \u2212 cos \u03b1) [16] In Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001054_bfb0015074-Figure4.2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001054_bfb0015074-Figure4.2-1.png", "caption": "Fig. 4.2. Tasks of robust holding", "texts": [ " One of f~, ff~ and ai constitute the independent parameters, that are redundant parameters to be optimized for load sharing. This is more generally stated in [27, 28]. We have proposed to tune the internal forces/moments f~ for simplicity of equations and also for consistency with control [23, 21]. One interesting problem regarding the load sharing is that of robust holding: a problem to determine the forces/moments ~, which the two robots apply to the object, in order not to drop it even when disturbing external forces/moments are applied. Tasks to illustrate the problem are shown in Fig. 4.2. This problem can be solved by tuning the internal forces/moments (or the load-sharing coefficients, of course). This problem is addressed in i[26], where conditions to keep holding are expressed by the forces/moments at the end-effectors, and Eq. (4.7) being substituted into the conditions, a set of linear inequalities for both f~ and a~ are obtained as: Af'~ + Bc~ < c (4.11) where A and B are 6 \u00d7 6 matrices, c a 6-dimensional vector, and Og : [ O~1, Of 2, \" ' ' , OL 6 ] T (4.12) In [26], a solution of c~ for the inequality is obtained, heuristically" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000073_s0094-114x(03)00005-3-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000073_s0094-114x(03)00005-3-Figure6-1.png", "caption": "Fig. 6. The PUMA robot analyzed in this paper. The screw $5, which is not explicitly designated on the Figure is on the joint axis 5 in the middle of the spherical wrist.", "texts": [ " Once this coordinate system is chosen, we obtain better results using graph theory techniques. This section illustrates the concept of hierarchical analysis to a widely spread industrial robot configuration. PUMA robots are also one of the most studied configurations found in research papers on robotics. To obtain the hierarchical canonical form of the Jacobian of this robot, the graph- and matrix-orientated techniques described as hierarchical analysis in the prior sections were used. The PUMA robot, Fig. 6, is a serial manipulator with six degrees of freedom. All joints are rotative kinematic pairs. The last three joint axes intersect one another in a single point forming a so-called spherical wrist. Position angles hi at the rotative joints are shown in Fig. 6. Screws $i are aligned with the joint axes and drawn as conical arrows. The chosen coordinate system is represented by the triad x, y, z. The Jacobian of a PUMA robot is a rather complex matrix-valued function, especially when obtained from the Denavit\u2013Hartenberg method. Using such Jacobian with huge elements the alternative is to select some auxiliary variables to compose these elements recursively [22]. However, using screw theory with a good choice of the coordinate system a significantly sparser Jacobian can be obtained [17]. This Jacobian, with the coordinate system sketched in Fig. 6, is J \u00bc s23 0 0 1 0 c5 0 1 1 0 s4 c4s5 c23 0 0 0 c4 s4s5 fc23 gs3 0 0 0 0 x14 0 0 0 0 0 fs23 x041 h 0 0 0 2 66666664 3 77777775 \u00f022\u00de where si \u00bc sin hi; sij \u00bc sin\u00f0hi \u00fe hj\u00de; ci \u00bc cos hi; cij \u00bc cos\u00f0hi \u00fe hj\u00de, etc. . . and x14 \u00bc gc2 \u00fe hc23, x041 \u00bc \u00f0gc3 \u00fe h\u00de, letters f , h, g are the distances shown in Fig. 6. The structure of the inverse Jacobian is [17] J 1 \u00bc 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 2 6666664 3 7777775 \u00f023\u00de where represents the non-null terms of the matrix, which are omitted here and in the following equation due to space restrictions. We use the Eq. (19) to rearrange the system, where the end-effector screw _x \u00bc $e \u00bc \u00bd _x1; _x2; _x3; _x4; _x5; _x6 T is represented in the coordinate system shown in Fig. 6, see details in [17]. Using Algorithm 1, the hierarchically rearranged linear system corresponding to Eq. (19) is \u00f024\u00de Extracting only the terms relative to the forward kinematics, i.e. dropping rows 2, 5, 6, 8, 10 and 12, Eq. (24) becomes \u00f025\u00de Matrix J is the hierarchical canonical form of the Jacobian given by Eq. (22). Now, we describe how this hierarchical canonical form may be explored to solve the rearranged system, Eq. (25), i.e. to invert the Jacobian by decomposition. In the sequence, we take a more detailed view of the singularities of the robot configuration" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000025_ecc.2003.7085301-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000025_ecc.2003.7085301-Figure1-1.png", "caption": "Figure 1: Vario X-Treme helicopter", "texts": [ " The helicopter dynamic model presented hereafter was the basis for the development of a simulator, named SimModHeli [2], implemented in Matlab, using Simulink and C MEX-file S-functions, that will be made freely available for the scientific community. This simulator is completely parameterizable and describes the dynamics of helicopters with any number of blades, with or without a Hiller or Bell-Hiller stabilizing bar. The simulation model is specially tailored for model-scale helicopters, such as the one depicted in Figure 1, and includes the rigid body, main rotor flapping, and stabilizing bar dynamics. The dynamics of the helicopter can be described using a six degree of freedom rigid body model driven by forces and moments that explicitly include the effects of the main rotor, BellHiller stabilizing bar, tail rotor, fuselage, horizontal tailplane, and vertical fin. To derive the equations of motion, the following notation is required: {U} - universal coordinate frame; {B} - body-fixed coordinate frame, with origin at the vehicle\u2019s centre of mass; p = [ x y z ]T - position of the vehicle\u2019s center of mass, expressed in {U}; \u03bb = [ \u03c6 \u03b8 \u03c8 ]T - Z-Y-X Euler angles that parameterize locally the orientation of the vehicle relative to {U}; v = [ u v w ]T - body-fixed linear velocity vector; \u03c9 = [ p q r ]T - body-fixed angular velocity vector" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000368_1-4020-3796-1_7-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000368_1-4020-3796-1_7-Figure4-1.png", "caption": "Figure 4. Joint resistance torque modeled with a non-linear torsion spring and damper.", "texts": [ " The set of data for the models is described in reference [15]. In the biomechanical model, no active muscle force is considered but the muscle passive behavior is represented. Applying a set of penalty torques when adjacent segments of the biomechanical model reach the limit of their relative range of motion prevents physically unacceptable positions of the body segments. A viscous torsion damper and a non-linear torsion spring, located in each kinematic joint, describe the joint torques. Take the elbow of the model, represented in Figure 4 for instance. The total damping torque for the relative rotation of the lower and upper arm is d i i ijm (5) where the torsion damper has a small constant coefficient ji and i is the relative angular velocity vector between the two bodies interconnected by joint i. 66 A constant torque mri that acts resisting the motion of the joint is applied in the whole range of motion in the dummy model [15]. For the human joint this torque has an initial value, which drops to zero after a small angular displacement" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001580_iemdc.2005.195714-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001580_iemdc.2005.195714-Figure2-1.png", "caption": "Fig. 2. Symbolic representation of studied two-phase machine", "texts": [ " 1, shows that the real machine, with phases star coupled, can be considered as the association of two two-phase fictitious machines: - The Main Machine (MM) whose quantities are the greatest - The Secondary Machine (SM) It can be proved that each quantity of a fictitious machine is composed of harmonic set of the same quantity of the real machine. In the case of five-phase machines, the Main Machine is composed of harmonics 1,9,... and Secondary Machine of harmonics 3,7,... [2], [16]. The Multi-machine Multi-converter System concept can be extended to systems with an arbitrary number of phases and shows that the control of a multi-phase drive can be in fact broken down into many simpler controls of two-phase machines. A symbolic representation of the two-phase machine is shown in Fig. 2. Let us define a stationary basis Bs = {\u2212\u2192x ,\u2212\u2192y } associated with the stator and a rotating basis Br = {\u2212\u2192d ,\u2212\u2192q } associated with the rotor vector flux. \u2212\u2192 t is a unity vector common to the basis Bs and Br and orthogonal to the four previous vectors. Quantities are: \u2022 \u2212\u2192 \u03c6s stator flux vector \u2022 \u2212\u2192 \u03c6sf flux vector exclusively produced by rotor magnets \u2022 \u2212\u2192v , \u2212\u2192 i stator voltage and current vectors \u2022 R, \u039b resistance and inductance of a phase \u2022 p number of pairs of poles \u2022 \u2126 = 1 p d\u03b8 dt mechanical speed \u2022 \u03b4 = \u03b8f \u2212 \u03b8s load angle between vectors \u2212\u2192 \u03c6sf and \u2212\u2192 \u03c6s Voltage equation is: \u2212\u2192v = R \u2212\u2192 i + ( d \u2212\u2192 \u03c6s dt ) /Bs (8) with: ( d \u2212\u2192 \u03c6s dt ) /Bs = \u039b ( d \u2212\u2192 i dt ) /Bs + ( d \u2212\u2192 \u03c6sf dt ) /Bs (9) and: ( d \u2212\u2192 \u03c6sf dt ) /Bs = ( d \u2212\u2192 \u03c6sf dt ) /Br + d\u03b8f dt ( \u2212\u2192 t \u00d7\u2212\u2192\u03c6sf ) (10) As Br is linked to the rotor flux: ( d \u2212\u2192 \u03c6sf dt ) /Br = \u2212\u2192 0 (11) leading to: ( d \u2212\u2192 \u03c6sf dt ) /Bs = d\u03b8f dt ( \u2212\u2192 t \u00d7\u2212\u2192\u03c6sf ) (12) Electromechanical torque is then: T = ( d \u2212\u2212\u2192 \u03c6sf dt ) /Bs " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002739_s1052618808040092-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002739_s1052618808040092-Figure1-1.png", "caption": "Fig. 1 Bevel gear and its equivalent.", "texts": [ " The mean normal gear module m n plays the role of a scale factor. In calculation of contact durability in the first approximation, the bevel gear with circular teeth can be replaced by an equivalent helical gear with tooth numbers z \u03c5 1 = z 1 /cos \u03b4 1 and z \u03c5 2 = z 2 /cos \u03b4 2 , where \u03b4 1 and \u03b4 2 are the angles of the pitch cone of the pinion and the wheel and are determined according to GOST (State Standard) 19326\u201373 . The interaxis distance is found as the sum of the radii of the initial circumferences r \u03c5 1 and r \u03c5 2 (Fig. 1). The aim of the optimization, i.e., of choosing the optimal values of \u03b1 n , , and x n for the given combination of tooth numbers z 1 and z 2 , is to achieve maximal contact durability of the gear. For cylindrical gears, ways of improving the geometry have been widely investigated [1\u20133]. The optimization procedure suggested in this work is based on the analysis of how design contact stresses at the pitch point depend on gear geometry. According to GOST (State Standard) 21354\u201387: Involute Cylindrical Gears of External Engagement: Strength Analysis , the relationship for determining design ha* ha* Vol", " 3, it can be seen that, as a result of optimization, the increase in rotating torque allowed by contact durability can be 46%. Technological synthesis of the optimized bevel gear. We will design a bevel gear with the profile angle \u03b1w at the initial cones with coefficients of normal displacement xn1 = \u2013xn2 = xw and with a coefficient of tooth head altitude = . Recall that the sum of initial cone angles should be equal to the interaxis angle of the gear; in this case, the angles of cone sockets \u03b4fi and the cone angles of the tooth tops \u03b4ai (Fig. 1) of the pinion and the wheel are determined according to GOST (State Standard) 19326-73 depending on the axial form of the tooth. The idea of the technological synthesis of the gear is to choose setups of the gear-cutting machines intended for pinion and wheel tooth processing such that they ensure the desired form and position of the contact spot. Under conditions of individual manufacturing, both sides of the wheel tooth are processed, as a rule, from one setting. In this case, the rolling ratio ir and tool radial setting U are calculated" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003792_053041-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003792_053041-Figure4-1.png", "caption": "Figure 4. (a) Side view of a walker composed of two spherical beads on a flat horizontal substrate, aligned with the shaking direction (the x-axis). The circled numbers define the indices corresponding to the three points of contact. The angular velocities \u03d5\u0307i are considered to be positive in the counterclockwise direction, and the velocity of the walker on the substrate vx is considered to be positive from left to right, as indicated by the arrow. (b) Sketch representing, with arrows, the friction forces F\u2016 i and the net normal forces1Fi that act on the three contact points, and the inertial forces that act on the centre of mass of the beads. All force vectors lie in the plane of symmetry of the system. The solid dark grey arrows correspond to forces that act on bead 1, while the dashed light grey arrows correspond to forces that act on bead 2. The inertial forces are considered positive when they point from bead 2 towards bead 1. The friction forces are considered to be positive in the direction opposite to the positive sliding velocities, and so in this sketch F\u2016 1 and F\u2016 2 are positive, whereas F\u2016 3 is negative. The resulting normal forces 1Fi are positive when they point towards the centre of the bead to which they apply.", "texts": [ " Finally, the friction laws provide the link between the friction and normal forces and the sliding velocities, and thus the locomotion velocity, as will be discussed in more detail below. We consider a walker, composed of two spherical beads with radii R1 and R2 and masses m1 and m2, which are held together and attached to a flat horizontal substrate by adhesive forces provided by liquid bridges. The walker is subjected to an external driving force that is parallel to the substrate. When the walker is aligned with the direction of the excitation, the plane defined by the three contact points is the plane of symmetry of the system (see figure 4), and all of the forces lie within that plane. The beads can rotate and slide on the substrate and on each other. The angular velocity of sphere i about an axis perpendicular to the symmetry plane is denoted by \u03d5\u0307i , and its sliding velocity on the substrate is denoted by vi . The sliding New Journal of Physics 13 (2011) 053041 (http://www.njp.org/) velocity of bead 1 on bead 2 is thus v3 = R1\u03d5\u03071 + R2\u03d5\u03072. Assuming that the spheres remain in hard mechanical contact with each other and the substrate, such that the geometry of the walker does not change over time, the velocity of the centre of mass of both beads and of the walker New Journal of Physics 13 (2011) 053041 (http://www", "org/) 9 Under driving (a 6= 0), when the beads are subjected to an inertial driving force, the friction forces are nonzero and the normal contact forces change from their value at rest to become F = F0 +1F. (3) As long as contact is maintained at each contact point, the variation in the normal contact forces 1F, which are the resulting forces acting on the contact points in the normal direction after summing the adhesion forces with the normal contact forces, can be derived from the balance of forces in the rest frame of the walker, as illustrated in figure 4(b). In the presence of an inertial driving force in the x-direction, the balance of forces acting on bead 1 gives F\u2016 3 cos \u03b4 +1F1 +1F3 sin \u03b4 = 0, 1F3 cos \u03b4\u2212 m1(a + v\u0307x)\u2212 F\u2016 1 \u2212 F\u2016 3 sin \u03b4 = 0, (4) where v\u0307x is the acceleration of the structure with respect to the substrate. The balance of forces acting on the walker gives 1F1 +1F2 = 0 (5) and F\u2016 1 + F\u2016 2 + M(a + v\u0307x)= 0, (6) since there is no net force acting on contact point 3 according to Newton\u2019s third law. Using equations (4)\u2013(6), the normal contact forces can then be written as F = F0 + 1 cos \u03b4 \u2212 1\u2212M 2 sin \u03b4 1+M 2 sin \u03b4 \u22121 1\u2212M 2 sin \u03b4 \u2212 1+M 2 sin \u03b4 1 1\u2212M 2 \u2212 1+M 2 sin \u03b4 F\u2016", " In the quasi-stationary limit, an approximation whose application is validated by the fact that no frequency dependence has been observed in the walkers\u2019 velocities, the rate of change of the rolling velocities of the beads on the substrate, Ri \u03d5\u0308i , and the acceleration of the walker on the substrate v\u0307x are negligible with respect to the acceleration of the substrate in the reference frame of the lab, a. An expression for the rate of change of the sliding velocities of the spheres comprising a walker in the general case is given in the theory supplement, section 1.1 (available from stacks.iop.org/NJP/13/053041/mmedia). In the quasi-static limit, a constant angular velocity \u03d5\u0307i for both of the spheres means that the overall torque due to the friction forces applied on bead i has to be equal to zero (as represented in figure 4(b)). As a result, and given our sign convention, F\u2016 1 = \u2212F\u2016 3 = F\u2016 2 . (8) Similarly, a constant locomotion velocity means that equation (6) becomes F\u2016 1 + F\u2016 2 = \u2212Ma. (9) After normalizing all of the forces by the force scale k0 R\u0304 given by the adhesive forces, the combination of equations (8) and (9) gives F\u2016 k0 R\u0304 = 0 2 \u22121 \u22121 1 . (10) New Journal of Physics 13 (2011) 053041 (http://www.njp.org/) Hence, the friction force on each contact point, normalized by the force scale k0 R\u0304, is proportional to the dimensionless driving force 0(t)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000768_j.snb.2005.02.018-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000768_j.snb.2005.02.018-Figure1-1.png", "caption": "Fig. 1. Electrochemical flow cell for the incorporation of the CuSPE.", "texts": [ " Potential links between changes in the seected urinary metabolites and the occurrence of both oxalate 28\u201332] and cystine [33] renal stone formation have been reorted, where creatinine may be used as an internal standard 30,31]. We hope that due to its simplicity and reproducibilty, the present method can be of clinical value in diagnosis of isorders to provide a prompt and simple technique in routine nalysis. rocessor pump drive, a Rheodyne model 7125-sample inector (20 l sample loop) with an interconnecting Teflon ube and a specifically designed electrochemical cell for ousing screen-printed electrodes. A picture of the assemly for the incorporation of the CuSPE into the wall-jet elecrochemical cell is depicted in Fig. 1. Screen-printed elecrodes are easily inserted and snapped into the leak tight ow cell. Connections for the separation column eluent (d), eference electrode (b), and counter electrode (c) are dislayed. Eluent from the flow cell was collected from the ounter electrode, which consists of a stainless steel HPLC ube. The disposable SPEs used to prepare the CuSPE were repared as reported earlier [23,24]. In brief, a Cu layer as electrochemically plated on a bare SPE in 200 mg/l u(NO3)2 plus 0.1 M HNO3 at \u22120" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000523_1.1850942-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000523_1.1850942-Figure6-1.png", "caption": "Fig. 6 Experimental RBFS rig for identifying the configuration state", "texts": [ " While there results a significant decrease in identification accuracy, the identification is still superior to that obtained using the Reynolds equation when there is no measurement error but a 1% error in estimating the running bearing clearance 7 . The current approach is likely to give superior identification to that using the Reynolds equation, and provided errors in measuring the slope can be kept sufficiently low, the current approach promises to provide believable configuration state identification in practice. Figure 6 is a three-dimensional view of a four-bearing test rig designed to run up to 3000 rpm. Load cells are attached to the flexible bearing supports to enable the transmitted forces to be measured. Changes in the configuration state can be introduced at the two inboard bearings. The bearing shells of the four circumferentially grooved journal bearings are free to pivot in the bearing housing, minimizing angular misalignment effects, and the flexible bearing supports are fixed to a rigid concrete block that is isolated from floor vibrations by air springs" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001311_bf00882589-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001311_bf00882589-Figure2-1.png", "caption": "Fig . 2", "texts": [], "surrounding_texts": [ "w 2. As an example we cons ide r a v i s c o e l a s t i c med ium whose h e r e d i t a r y p r o p e r t i e s a re de t e rmined by RabotnovVs weakly s ingu la r ke rne l s (see [5, 6])\nf n\nR (0 = : - ~ ~ ( - - 1) ~ ,~(~+')V [~ (n + I)1 \" n-=O\n(2.1)\nHere ~ is the c h a r a c t e r i s t i c r e l axa t ion t ime; F is the g a m m a - f u n c t i o n ; 3~ is a s ingu la r i t y (fractionali ty) p a r a m e t e r ; f o r ~/= 1 we obtain the usua l exponent ia l model , which c o r r e s p o n d s to the mode l of a s t anda rd l i nea r body.\nThe Lap lace t r a n s f o r m of the ke rne l (2.1) has the f o r m\nR. (p) = [l + 0~,)v] - l . (2.2)\nSubst i tut ing the e x p r e s s i o n (2.2) into Eq. (1.7), we obtain co\n]} u , = exp - oxp[-p . )2(k + : ~ J, I. \" coo\n, , (2 .3)\na v ----- ~-VA; A ---- EoE~'; b v ---- T~-v; oa v (p) = (pV + av)T (pv + by ) - T.\nThe so lu t ion in the space of o r ig ina l s is obtained f r o m the M e l l i n - F o u r i e r i nve r s ion f o r m u l a\nct+t'oo\nu (x, t) = ~ u, (x, p) exp (P0 dp. (2.4)\nTo evalua te the in teg ra l (2.4) i t i s n e c e s s a r y to d e t e r m i n e all the s ingu la r points of the complex func - t ion u . . I t is evident f r o m the e x p r e s s i o n s (2.3) that the in t eg rand funct ion has a s e c o n d - o r d e r pole at the point p = 0 and b r anch points at p = 0 and p = - , o .\nSince the i n v e r s i o n t h e o r e m is appl icable to mul t ivalued funct ions only on the f i r s t shee t of a R i e - mann su r f ace (0 < a r g p ~ 2~r), we m u s t choose the c losed con tour of in tegra t ion as shown in Fig. 1. By J o r d a n , s l e m m a the in t eg ra l s along the cu rves c R tend to z e r o fo r R --- ~o provid ing that t > (2kl + x )c~ 1, t > [2(k + 1 ) / - x ] c ~ . Using the fundamenta l r e s idue t heo rem, we can wr i te the e x p r e s s i o n fo r u(x, t) in the f o r m\nHere\nF k = exp [s (2tet + x) c~tR v cos 0 v - - st] sin [s (2kl + x) c~.'R v sin 0vl;\nG k = exp is 12 (k + 1) l - - xl c~ l i l~ , COS 0~. - - St]'\n\u2022 sin {s [2 (k + 1) 1 - - x] c~'R v sin 0v};\nyk = t - - ( 2 k l + x ) ~ t ; z, = t - - [ 2 ( k + 1) / - - x ] ~ ;\n1 l 1 R, = r ?~,r2v ~ ; o , = -~ (%~ - %.) ;\n2 ~ s ~ + 2bvs.~ cos n? + b2v; 2 s ~ + 2avsV cos z~, + a~; r v r l y\n--1 t g ~ , v = s V s i n ~ y ( a ~ + s cos~7) ; tg~2v=sVsina?(bv+sVcosaT)-l;\nH is the Heavis ide unit function.\nI i i 0", "F o r the s t a n d a r d l i n e a r body m o d e l (y = 1) we ob ta in\nr\nu, p2 + ~ ( - -1) k xp - -po) -2M + x + e x p - - p o ( k + l ' ) l - - x\nk = o\nf r o m Eq. (2.3), w h e r e ~ = [ i m w ,. 7->1 Y\nThis func t ion has a s e c o n d - o r d e r pole a t the poin t p = 0 and b r a n c h poin ts a t p = - a and p = - b .\nChoos ing the i n t e g r a t i o n con tour shown in F ig . 2, we ob ta in the e x p r e s s i o n\n(2.6)\n(2.7)\nwhere\nF~(x,t,s, 1) = limF~(x,t,s, 7); a = lim av; ~,,_) [ 'y,-~! O k (x, t, s, 1) = lim G~ (x, t, s, 7); b = ltmb v. 'y,,,~,l 'Y\"> |\nThe so lu t i ons (2.5) and (2.7) a r e s u i t a b l e fo r use only whi le the rod m a i n t a i n s con tac t wi th the t a r g e t , i . e . , p r o v i d i n g the s t r e s s o-(0, t) < 0.\nThe i n s t a n t a t which r e c o i l o c c u r s i s ob ta ined f r o m the cond i t ion ~ (0, t) = 0; i t d e t e r m i n e s the n u m - b e r of t e r m s in the s u m m a t i o n s of the s e r i e s (2.5) and (2.7). Th is quan t i t y depends on the r e l a x a t i o n t i m e ~'e, which i s r e l a t e d to the t e m p e r a t u r e T by t h e A r r h e n i u s f o r m u l a ~'e = Te0exp (UR-tT-1) . It i s t h e r e f o r e n e c e s s a r y to d e t e r m i n e how the r e c o i l t i m e depends on the t e m p e r a t u r e .\nw 3. The e x p r e s s i o n fo r the s t r e s s in the L a p l a c e t r a n s f o r m s p a c e has the fo l lowing f o r m :\na , (0, p) = - - pc. (p) Vop-i th [plc71 (p)]. (3.1)\nUs ing the t h e o r e m s c o n c e r n i n g the e x p a n s i o n of m e r o m o r p h i e func t ions [3], we ob ta in the fo l lowing e x p a n s i o n f o r or,(0, p)\nr\n2% X-~ P~ + ~x; -v G , (0, P) pc~\n-7- 7~ p2+~ + p2xVv + c2p~ + Co,~ 2 _~ 9 (3.2)\nMoreover\n1 (2n ~ 1) ~ l - l c o . 1 (2n - - 1) ~l--lc~; COn = E o ~ ; c=~ = ~- =\nTo i n v e r t the t r a n s f o r m (3.2) and r e t u r n to the s p a c e of o r i g i n a l s i t i s n e c e s s a r y to d e t e r m i n e a l l the s i n g u l a r po in t s of the e x p r e s s i o n (3.2). Th i s e x p r e s s i o n has b r a n c h poin ts a t p = 0 and p = _0o and a l so s i m p l e po le s at those po in ts p w h e r e the d e n o m i n a t o r v a n i s h e s , i . e . , f o r the r o o t s of the equa t ion\np~+V + p2.~-v + c=np + co,~'~ = O. (3.3)\niiii", "We can show tha t Eq. (3.3) has two c o m p l e x - c o n j u g a t e roo t s Pi,2 = r n e x p (~:iCn) whose dependence on -ra was inves t iga t ed in [1, 2].\nTak ing these fac t s into account , we can wr i t e an e x p r e s s i o n f o r ~r(0, t) in the f o r m\nco 20o o(0, t) = - - c~,P-l-- [A~ + An exp (-- %t) sin (co~t - - ~)] , (3.4)\nn~-I\nH e r e\nA n = 2 V-(h 2 + q2)-i (rn2V ._~ ~2T~-2v .~ 2~,~-VrV cos 7%);\nrnV sin (7~n + Zn) + ~v/~v sin Zn . hn tg r = - - r~ cos (y% + Z.) + ~%-v cos Z. ' tg Z n = --qn ;\nh, = (2 + 7) r~ +v cos (I + 7) @, + 2g~-Vr. cos %\n+ v c ~ r v - ' cos(,] - - 1) r\nqn = (2 + ~) r~+V sin (I + 7) % + 2v[vr n sin ,~ + wL~rX -~ sin (y - - 1) , . ;\nr\nA ~ (0 = i s - ' B (s) exp ( - - st) cls. 0\nThe quant i ty A~ can be r e g a r d e d as the L a p l a c e t r a n s f o r m of the s p e c t r a l funct ion\nsinz~7 [--s2+~(1 --~) + sV ( e L ~ - ~.)l~2~-I (s)\n(3.5)\n(3.6)\nw h e r e\nC 2 \"~--1 Qn (S) = (S 2+~ + cLn8 ~)(S 2 + On] 9\nThis funct ion f u r n i s h e s a d i s t r i bu t ion of the r e l a x a t i o n a l p a r a m e t e r s of the d y n a m i c a l s y s t e m .\nIn the ca se of the s t anda rd l i nea r body mode l the e x p r e s s i o n f o r er(0, t) a s s u m e s the f o r m\no (0, t) = y ~ (0, 0 + ~ (0, 0. n = I n-~-k+,\nH e r e\n3 e~ = - - 2voc~p1-1 ~, H ~ exp ( - - [~i~t);\no~ = - - 2VocLpl-' IQI~ exp ( - - ~ ld ) + Q~ exp (-- u~nt) Sin (~nt - - X~)]; 1\n0,,n = de-~; Q~ = (d 2 + f~) ~- e-'; tg ~,~ = df~q;\nd = ~L n ( ~ l n \"-~ S~); e = ~g n [ ( ~ I n ~ $r )2 -97 ~L2I;\n(3.7)\nThe quant i t ies Bin, ~ l n a r e the r e a l roo t s of the cubic equa t ion obtained f r o m Eq. (3.3) fo r y = 1; ~2n :~ ip n a r e the c o m p l e x - c o n j u g a t e roo t s of Eq. (3.3). The p r o p e r t i e s of the r o o t s of this cubic equat ion as a funct ion of r e w e r e s tudied in [4]." ] }, { "image_filename": "designv11_61_0000927_tmag.2005.846256-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000927_tmag.2005.846256-Figure1-1.png", "caption": "Fig. 1. Arrangement of the Roebel bars inside the slot.", "texts": [ " To accurately include both the volume forces inside conducting regions and the surface forces at material boundaries in the coupled analyses, a weak coupled cascade algorithm of electromagnetic and mechanical finite-element calculations can be applied. Moreover, this algorithm allows to describe the transitional state between both fixed and loose bars in dependence of the electrical bar currents in an efficient way. Index Terms\u2014Coupled field problems, electromagnetic coupling, finite-element methods, synchronous generators. I. INTRODUCTION THE cross section of a typical Roebel bar structure with filling materials inside the slot is shown in Fig. 1. The upper and lower bar itself consists of several subconductors, insulation, middle filler piece, and enclosing envelope [1]. A slot wedge fixes all parts inside the slot and should avoid any free space in radial direction. The mechanical stress inside the bars is caused by forces of mechanical and electromagnetic origin. In case of normal operational states, the mechanical preload applied by the slot wedge mainly determines the mechanical behavior. This is due to the fact that stress contributions caused by rated currents are not observable in the total mechanical stress distribution" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000898_cira.2003.1222115-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000898_cira.2003.1222115-Figure6-1.png", "caption": "Figure 6: Forward kinematics from links 2 and 3", "texts": [ " Therefore, the problem we are facing becomes: How to find a forward kinematics map which minimizes the error d u e to these uncertainties? 5 Different Forward Kinematics Maps Taking the 2-dof redundant actuated parallel manipulator We will present several forward kinematics as an example. maps. 5.1 Forward Kinematics Map From Links 2 and 3 Only f23(aZ,a3) Consider the serial l i n k 2 and 3. Th? is a 5-bar linhge. Given the corresponding actuators or sensors angles (a2,a3). one can compute the end-effector coordinates (x, y). A s we can see from Figure 6, (x,y) is just the coordinates of the point of intersection of two circles with radii T Z y d TJ centered a t rbz and rb3 ,respectively, where Therefore, the end-effector coordinates (x, y) can he computed by solving the following two equations: (2) (Z - Zb2)\u2019 + (Y - Yb2)\u2019 = TZ2 (5 - Zb3) \u2019 + ( Y - Yb3)\u2019 = 73\u2019. To solve the above set of equations, we f i s t find the equation of the straight line passing through the intersection points by subtracting the first equation from the second equation: a X + b y + c = o (3) a = -2(Zb2 - 5 b 3 ) (4) b = -2(Yb2 - Y b 3 ) ( 5 ) where 2 2 2 2 2 C = ( X b z \u2019 t Ysz - T2 ) - ( 2 b 2 + Ybz - T 2 )" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003257_1.2952448-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003257_1.2952448-Figure1-1.png", "caption": "Fig. 1. Example 1. The cat model with chosen as the angle between the x axis and the b rod.", "texts": [ " The maximal rotation due to the largest possible stroke must be the same as a consequence of a topological invariant see Sec. VII . The model is composed of four spheres each with mass m connected by four massless rods to form a parallelogram. 1040 Am. J. Phys. 76 11 , November 2008 http://aapt.org/ajp Downloaded 01 Oct 2012 to 152.3.102.242. Redistribution subject to AAPT lic The cat can control the base angle and the length a of one of the pairs of parallel rods. The other pair of rods has a fixed length b see Fig. 1 . Because the model is made of four masses, each with two degrees of freedom, it has a total of eight degrees of freedom. However, there are five constraints\u2014four distances between the masses due to the rods and one angle. Thus, there are only three physical degrees of freedom, which can be expressed as the location of the center of mass and the total rotation . Because the velocity of the center of mass will not change without external forces, we will work in the reference frame in which the center of mass is at rest, and will not consider its two degrees of freedom. This choice leaves only one physical degree of freedom\u2014the total rotation of the body. There are many ways to choose the angle as we will discuss in Sec. III. We will choose as the angle between the x axis of an arbitrarily oriented reference frame whose origin is at the center of mass and one of the b rods see Fig. 1 . We now show that the cat can rotate itself by changing a and . The equation of motion of the cat is based on the conservation of angular momentum which, in a twodimensional system, is a scalar , dL /dt=0. The initial angular momentum is zero, so L=0 for all times. We can express L in terms of the rate of change a\u0307, \u0307, and \u0307. A direct calculation shows that the equation of motion is L = 4\u0307 a2 + b2 + 4\u0307a2 = 0. 1 From Eq. 1 we can derive an equation for the change in . Because \u0307=\u2212\u0307a2 / a2+b2 , we write =\u2212 \u0307a2 / a2+b2 dt, or = \u2212 a2 a2 + b2d ", " Thus must be equal for all choices. For a sequence of changes that does not end with the same a and as the initial configuration that is, one that is not a \u201cstroke\u201d , the total rotation is not well defined and depends on the way we have characterized the angle our choice of gauge, which is, as we will see, a local choice . To demonstrate why the total rotation is not well defined for a sequence of shape changes that does not end with the same shape, we consider two cases see Figs. 1 and 2 . In Fig. 1 is chosen as in Sec. II as the angle between the x axis of the reference frame and the b rods. In Fig. 2 the angle is chosen as the angle between the x axis and the line connecting two opposite masses. The two choices for are equally valid: In both cases a , and completely describe the system, and neither choice is superior to the other or any other way to choose . However, it is easy to see that for a path in a , -space that is not closed, is different for the two choices. For example, consider a change in a alone while keeping = /2", " Redistribution subject to AAPT lic A\u0303 = A + f . 9 B will not change under the gauge transformation, because B\u0303= A\u0303 / a\u2212 A\u0303a / =B+ 2f / a\u2212 2f / a =B. As for the magnetic vector potential, A can be changed by f for any scalar function f . For magnetism we usually have no simple interpretation for f , but we do have a simple interpretation for f for the cat\u2014it is the difference in angle between the two ways of measuring the system\u2019s orientation. Problem 2. Calculate A for Fig. 2 and verify that B is the same as in Fig. 1. We have seen that our simple mechanical model can be used as a concrete example of gauge freedom. Can the gauge theory formalism teach us anything about the mechanical system? We will give two examples of properties of the system that are naturally understood in the gauge theory formalism: The first one in this section is the answer to the question, \u201cWhat deformations lead to a net rotation?,\u201d and the second in Sec. VII is the rotation due to the maximal possible stroke. Consider the same system, but fix = /2 and let b be the control instead" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000522_progress.1.2005.1.2-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000522_progress.1.2005.1.2-Figure1-1.png", "caption": "Figure 1. Special idle running rear wheel hub constructed for measurement in only front-wheel drive work mode", "texts": [ " The surface of the field used for the experiments was horizontal and plane, without considerable surface irregularities. The soil was loamy and it was covered with wheat stubble. The experiments were performed with New Holland tractor type TM165 which was offered for the tests by the New Holland Office in Hungary. The front-wheel drive work mode meant a problem: how to switch off the rear-wheel drive, or rather how to operate the tractor only with frontwheel drive. For this reason a special idle running rear wheel hub was constructed (Figure 1). The front-wheel drive was fitted with torque transducer and rotation sensor (Figure 2). The flowchart of the measurement and data processing is shown in Figure 3. The experiments, including the measurements of the dynamic rolling radius were carried out with the TM165 tractor. The rolling radius was calculated from the distance run and from the number of wheel revolutions. During these measurements the tractor was either driven or pulled. The average values were calculated from the results. The tractor was driven Slip Calculation and Analysis for Four-wheel Drive Tractors 11 with rear wheel drive for determining the dynamic rolling radius of the rear tires in \u201cdriven mode\u201d" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002236_2007-01-3740-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002236_2007-01-3740-Figure5-1.png", "caption": "Figure 5: Two Roller Control Mechanism", "texts": [ " Figure 3: Principle of Torque Control Castor Angle Tangential Force EndloadEndload Reaction Force Figure 4: Variator Force Balance To compete in the cost competitive OPE market, the full toroidal traction drive Variator required significant simplification in both design and operation. Recognising the low power and torque requirements of this market, the simplification was primarily achieved by utilizing a single cavity design, reducing the number of rollers in the cavity from three to two and introducing a simple \u2018yoke\u2019 style roller control mechanism (figure 5). This two roller design has been validated, is in \u2018series production\u2019 and has now been applied to higher power and torque applications in both IVT format for entry level vehicles for emerging markets and CVT format for subA, A & B class vehicles. VEHICLES IN EMERGING MARKETS Of the multiple automotive markets and sectors throughout the world, the area of greatest growth is the requirement for personal mobility for the emerging mass markets such as India, China, etc. Automotive solutions for these markets need to be simple, low-cost and efficient whilst being easy to use, technically innovative and, critically, of high quality" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001308_s11249-006-9080-1-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001308_s11249-006-9080-1-Figure1-1.png", "caption": "Figure 1. Contact between a rough deformable surface and a rigid flat plane.", "texts": [ " The present work considers a new elastic plastic model for adhesive contact analysis of JKR contacts following the plastic asperity concept by Abdo and Farhang [26] that provides the more realistic picture of plastic deformation of rough surfaces through the definition of fictitious plastic asperities that are assumed to be embedded at a critical depth within the actual surface asperities. The best-known stochastic model for elastic contact of real rough surfaces is the model given by Greenwood and Williamson [16]. In such a model a contact is represented by contact between a rigid smooth surface and a rough deformable surface with small-scale asperities as shown in figure 1. The rough surface is isotropic, asperities are spherical near their summits with a uniform radius R and with a Gaussian distribution of heights with standard deviation r. The model assumes that the asperities are far apart, and there is no interaction between them and no bulk deformation occurs during contact. In figure 1, z and d denote the asperity height and mean separation of the surfaces, respectively. In a recent pioneering work [26], Abdo and Farhang presented the plastic asperity concept for modelling the elastic\u2013plastic contact of rough surfaces. The salient feature of their approach is outlined here in brief in order to set a scene for the present analysis. Considering the contact between one single asperity on rough surface and a rigid plane, the behaviour of the asperity is initially elastic. As the load is increased the elastic behaviour continues to describe the deformation until a critical interference is reached", " This can simply be written as Pe pa2 0:6H: \u00f05\u00de Replacing the radius of apparent Hertzian contact of an asperity by \u00f0Rd1\u00de1=2 and combining equations (3) and (5), results d3=41 0:6pH K R1=2d1=41 \u00f06pc\u00de1=2R1=4 K1=2 0: \u00f06\u00de The equation gives the condition for inception of yielding and may be solved to give the critical value of asperity displacement dc1, which marks the transition from elastic to elastic\u2013plastic deformation. In equation (6) d1 represents apparent displacement due to an apparent Hertz load given by Pe \u00fe \u00f06pcKa3\u00de1=2. Following Johnson [13], dc1 may be related to real critical displacement dc by dc1 \u00bc dc \u00fe 2 3 6pca K 1=2 : \u00f07\u00de Considering now the contact between a rigid smooth surface and a rough deformable surface (figure 1) with a Gaussian distribution /(z) of asperity heights z such that /\u00f0z\u00de \u00bc 1 r ffiffiffiffiffi 2p p e z 2=2r2 \u00f08\u00de and d=z)d. If N is the number of asperities per unit area of the rough surface, the total applied load on all the asperities in contact per unit area may be written following equation (2) and is given by P1 \u00bc N Z 1 d Pe/\u00f0z\u00dedz N Z 1 d\u00fedcl Pe/\u00f0z\u00dedz \u00feN Z 1 d\u00fe\u00fedc Pp/\u00f0z\u00dedz, \u00f09\u00de where the first two integrals represent the net elastic contribution to applied load and the third integral represents the plastic contribution to applied load" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001913_s11668-007-9016-6-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001913_s11668-007-9016-6-Figure2-1.png", "caption": "Fig. 2 Eccentricity defect", "texts": [ " \u00f0m1 \u00fe m2\u00de \u20acx1 \u00fe C1 _x1 \u00fe kx1x1 \u00bc FTx1=2 \u00f0m1 \u00fe m2\u00de \u20acy1 \u00fe C1 _y1 \u00fe ky1y1 \u00bc FTy1=2 \u00f0m1 \u00fe m2\u00de\u20acz1 \u00fe C1 _z1 \u00fe kz1z1 \u00bc P \u00f0m3 \u00fe m4 \u00fe m5\u00de _x2 \u00fe C1 _x2 \u00fe kx2x2 \u00bc FTx1=2 \u00f0m3 \u00fe m4 \u00fe m5\u00de \u20acy2 \u00fe C1 _y2 \u00fe ky2y2 \u00bc FTy1=2 \u00f0m3 \u00fe m4 \u00fe m5\u00de\u20acz2 \u00fe C1 _z2 \u00fe kz2z2 \u00bc 0 \u00f0I1 \u00fe I2\u00de \u20acd1 \u00fe 2kdd1 Tf \u00f0 _d2\u00de \u00fe Tf \u00f0 _d3\u00de \u00bc Te\u00f0t\u00de \u00f0I2 \u00fe I3\u00de \u20acd2 kdd1 \u00fe 2Tf \u00f0 _d2\u00de \u00fe Tf \u00f0 _d3\u00de K1f1\u00f0d4\u00de C34 _d4 \u00bc 0 \u00f0I1 \u00fe I3\u00de \u20acd3 \u00fe kdd1 \u00fe Tf \u00f0 _d2\u00de \u00fe 2Tf \u00f0 _d3\u00de K1f1\u00f0d4\u00de C34 _d4 \u00bc Te\u00f0t\u00de \u00f0I3 \u00fe I4\u00de \u20acd4 Tf \u00f0 _d2\u00de Tf \u00f0 _d3\u00de \u00fe 2K1f1\u00f0d4\u00de K2f2\u00f0d5\u00de \u00fe 2C34 _d4 C45 _d5 \u00bc 0 \u00f0I4 \u00fe I5\u00de \u20acd5 K1f1\u00f0d4\u00de \u00fe 2K2f2\u00f0d5\u00de C34 _d4 \u00fe 2C45 _d5 \u00bc TD 8 >>>>>>>>>< >>>>>>>>>: \u00f0Eq 6\u00de An eccentricity of a disc is theoretically the distance between the geometric and rotating axes of the disc. We create an eccentricity on the flywheel and the pressure plate of the clutch system as shown in Fig. 2. This eccentricity is defined by the parameter e1, which represents the distance between the axes, and by a phase k1 to specify the initial position. The terms O1 and G1 represent, respectively, the rotational and geometric centers of the flywheel and the pressure plate. An eccentricity induces a kinetic energy variation, since the position of the flywheel and the pressure plate will be changed. The kinetic energy coming from the eccentricity defect is given by: Ecex \u00bc 1 2 \u00f0m1 \u00fe m2\u00de 2 _x1 _Ux1 \u00fe 2 _y1 _Uy1 \u00fe _U2 x1 \u00fe _U2 y1 n o \u00f0Eq 11\u00de The terms Ux1 and Uy1 designate the supplementary linear displacements expressed in the stationary reference axis defined by: Ux1 \u00bc e1 cos x1t k1\u00f0 \u00dehboxand Uy1 \u00bc e1 sin x1t k1\u00f0 \u00de \u00f0Eq 12\u00de We apply the formalism of Lagrange to determine the kinetic energy coming from the eccentricity defect: d dt @Ecex @ _qi @Ecex @qi \u00bc Fcexf g \u00f0Eq 13\u00de Consequently, {Fcex (t)} can be written as: Fcex\u00f0t\u00def g \u00bc f \u00f0m1 \u00fe m2\u00dee1X 2 1 cos x1t k1\u00f0 \u00de; \u00f0m1 \u00fe m2\u00de e1x 2 1 sin x1t k1\u00f0 \u00de; 0; 0; 0; 0; 0; 0; 0; 0; 0gT \u00f0Eq 14\u00de An eccentricity defect also causes additional tangential forces FTxex1/2 and FTyex1/2 applied by the first block to the second" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001628_icar.2005.1507479-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001628_icar.2005.1507479-Figure3-1.png", "caption": "Fig. 3. Coordination: in-line formation", "texts": [ " The speeds at which the robots are required to travel can be imposed in a number of ways; for example, by nominating one of the robots as a formation leader, assigning it a desired speed, and having the other robots adjust their speeds accordingly. Figures 2 and 3 show the simple cases where 3 vehicles are required to follow the straight paths and circumferences \u0393i; i = 1, 2, 3 while keeping a desired \u201dtriangle\u201d or \u201din-line\u201d formation pattern. In the simplest case, the paths \u0393i may be obtained as simple parallel translations of a \u201dtemplate\u201d path \u0393t (Fig. 2). A set of paths can also be obtained by considering the case of scaled circumferences with a common center and different radii Ri (Fig. 3). In this paper, for simplicity of presentation, we restrict ourselves to \u201din-line\u201d formation patterns and to the types of paths described above. However, the methodology proposed for coordinated path following can be extended to deal with arbitrary paths and even with time-varying formation patterns. See (Ghabcheloo et al., 2004c) for details. Assuming that separate path following controllers have been implemented for each robot, it now remains to coordinate (that is, synchronize) them in time so as to achieve a desired formation pattern" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003217_wmso.2008.61-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003217_wmso.2008.61-Figure1-1.png", "caption": "Figure 1. The mechanical structure", "texts": [ " Because all the nations regard space technology as the secret, no unabridged data can be referred. Nevertheless, under the cooperation between Dalian University of Technology and Beijing Institute of Space Test Technology, a dynamic thrust measurement system still has been researched for attitude and orbit rocket, which has independent intellectual property. II. MODEL OF THE SYSTEM Considering the measurement principle and method, static calibration, rocket installation etc, the dynamic thrust measurement system for attitude/orbit rocket is as shown as in fig.1. The key equipment is 4 as shown, in which there are piezoelectric sensors assembled. The main aim is to analyze theoretically the effects the system takes to the rocket pulse thrust measurement. Due to complexity of structure, it is difficult to separate every part to solve. And the Finite Element Method is affected greatly by boundary conditions. Finally, select the experimental modal analysis to model the system. For the mechanical system, viscous damping and structural damping act together. This two damping determine the dynamic model expression different" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002340_978-3-540-87732-5_96-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002340_978-3-540-87732-5_96-Figure1-1.png", "caption": "Fig. 1. Configuration and mechanical properties of hybrid five-bar mechanism", "texts": [ " In recent years, there have been a lot of reported works focused on the PSO which has been applied widely in the function optimization, artificial neural network training, pattern recognition, fuzzy control and some other fields [8] in where GA can be applied. Dimensional synthesis is an important subject in designing a hybrid mechanism. The main purpose of the paper is to present modeling and analysis of hybrid mechanism system, and to investigate the optimal dimensional synthesis of the mechanism by its mathematical model. By means of dynamics objective functions, optimum dimensional synthesis for hybrid five bar mechanism is taken by using a PSO algorithm. The results and analysis of an example are obtained in this study herein. Fig.1 represents five link mechanism structure having all revolute joints except one slider on output link, and shows the positional, the geometrical and the dynamic relationships. Notations shown in Fig. 1 are applied throughout the study. The hybrid mechanism has an adjustable link designed to include a power screw mechanism for converting rotary motion to linear motion by means of a small slider. The slider is assumed to move on a frictionless plane. The crank is driven by a main motor (DC motor) through a reduction gearbox; the slider is driven by a lead screw coupled an assist motor (servomotor). Here the main motor is applied as a constant speed motor, and the constant speed motor profile is applied", " (3) From (3), we may found the dynamic equation of hybrid mechanism in the form 0241321 =+++ UAUAqAqA KIKI . (4) Where 321 ,, AAA and 4A are coefficient matrixes of dynamic equation, 1U represent input torques (forces) vector, and 2U represent other torques (forces) on hybrid mechanism. Thus, the 2-vector 1U can be found from equation (4), )()1( 52421 1 31 AUAqAqAAU KIKI +++\u2212= \u2212 . (5) Kinematic analysis of five bar linkage is needed while carrying out dynamic analysis. The mechanism is shown with its position vectors in Fig. 1. The output of system is dependent on two separate motor inputs and the geometry of five bar mechanism. By referring to Fig. 1, the loop closure equation is written as: EDAECDBCAB +=++ . (6) By solving vector loop equation (6), we can obtain angular position of the each link. Having found the angular displacements of each linkage in the five bar linkage, time derivatives can be taken to find angular velocity and accelerations. They are also definitely needed during the analysis of dynamic model. Hybrid mechanism can be determined by selecting a design vector as follows Txxxxxx ],,,,[ 54321= , (7) where 0504321 xxdexdbxdax \u03c8\u03c6 ===== ,,,, ", " When the number of maximum iterations or minimum error criteria is attained, the particle with the best fitness value in X is the approximate solution andSTOP; otherwise let iter=iter+1 and turn to Step 2. Where iter denotes the current iteration number. Particles\u2019 velocities on each dimension are clamped to a maximum velocity v Bmax B, If the sum of accelerations would cause the velocity on that dimension to exceed v Bmax, which is a parameter specified by the user. Then the velocity on that dimension is limited to v Bmax B. In order to test the validity of the proposed procedure and its ability to provide better performance, an example problem was solved. Experimental model in Fig. 1 have been designed that link dimensions can be adjusted statically relative to crank, the masses of links, inertias and positions to the masses center of links in the local coordinates are independent of the static adjustment. Mechanical properties of five bar mechanism, link lengths, positions to the center of gravity of each link, link masses, and link inertias on the masses center are shown in Table 1. These link lengths and angle values for hybrid five bar mechanism in the studies of optimal kinematics design were obtained by Wang [5]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002241_978-1-84628-469-4_15-Figure15.2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002241_978-1-84628-469-4_15-Figure15.2-1.png", "caption": "Figure 15.2. (a) Sensor range of UAV. (b) Example of obstacle avoidance", "texts": [ " Once an obstacle has been detected, the previous flight path is corrected to ensure safe flight while minimizing the deviation of UAV from the pre-planned path. Simulation studies on multiple obstacles with various shapes are conducted and the effectiveness of the proposed method is verified. We consider the problem of UAV trajectory tracking without detail information about the flight environment \u2013 only the position of the UAV is available from a GPS (Global Position System) receiver. We assume that the obstacles can be detected by a UAV obstacle sensor. The sensor range is shown in Figure 15.2. We are interested in maintaining close tracking of heading speed and heading angle of UAV, which are governed via (.) (.) (.) (.) v vv f g vc f g c (15.1) where and vc c are the commanded velocity and heading angle to the autopilots. and (.)fv (.)f are the nonlinear functions of the system, and(.)gv (.)g represent the control gains. With the lateral and longitudinal coordinates as shown Figure 15.3, we can establish cos sin x l y l , (15.2) where l , in which is the distance between the UAV and the departure point, and v l x and are the lateral and longitudinal positions of UAV with respect to the departure point" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002675_j.jmatprotec.2009.01.005-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002675_j.jmatprotec.2009.01.005-Figure4-1.png", "caption": "Fig. 4. Vector representation between the neighboring", "texts": [ " .2. Laser scan path generation through coordinate ransformation The method discussed above for calculating the intersection of a ontour point is prone to error and is shown in a situation depicted y Fig. 3. This situation occurs when the calculated scan point is between he contour points and scan rays. A wrong scan sequence of the jump\u201d and \u201cmark\u201d commands is produced. To address this probem, contour points can only be selected as scan points if the ondition reflected in Eq. (3) is satisfied. Fig. 4 shows the vector epresentation among three points. Eqs. (1) and (2) represent the ngle between each vector and the X-axis or Y-axis, respectively Kim et al., 2007a,b). 1 = cos\u22121 ( x1 \u2212 x0\u221a (x1 \u2212 x0)2 + (y1 \u2212 y0)2 ) ; (1) 2 = cos\u22121 ( x2 \u2212 x1\u221a (x2 \u2212 x1)2 + (y2 \u2212 y1)2 ) ; and (2) 0\u25e6 < 1 \u2264 180\u25e6 and 0\u25e6 \u2264 2 < 180\u25e6, 180\u25e6 < 1 \u2264 360\u25e6 and 180\u25e6 \u2264 2 < 360, (3) scan points for sintering parallel to X and Y-axis. a h i E s A scanned path is selected parallel to each of the X and Yxis, alternately. Due to alternate scanning and sintering, 1 and 2 ave to be defined separately for each axis" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003612_s11044-010-9190-2-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003612_s11044-010-9190-2-Figure8-1.png", "caption": "Fig. 8 Vibration model split up into subsystems", "texts": [ " The stiffness values cfy and cfz and the damping values dfy and dfz are the values for each machine side. The coordinate systems for V (zv;yv) and B(zb;yb) have the same point of origin, as well as the coordinate systems for the stator mass ms(zs;ys) and for the rotor mass mw(zw;yw). They are only shown with an offset to show the connections through the various spring and damping elements. As excitation of this vibration system all three different kinds of dynamic eccentricity\u2014eccentricity of rotor mass \u00eau, magnetic eccentricity \u00eam and bent rotor deflection \u00e2\u2014are superposed in the model (Fig. 8). Static eccentricity is not considered in the vibration model. First of all, it is necessary to derive the differential equation of the vibration system of the soft mounted electrical machine. 4.1 Differential equation of the vibration system To derive the differential equation, it is necessary to split up the vibration system into individual systems\u2014rotor mass system (a) journal system, (b) bearing house system, (c) and stator mass system, (d) (Fig. 8). The fictitious point A in Fig. 8(a) describes the orbit of the shaft centre point W if the rotor shaft were to be rigid (c \u2192 \u221e). Therefore, the displacement between points A and W describes the dynamic elastic bending of the shaft w\u0302elast. Because of small displacements of the stator mass zs, ys, \u03d5s compared to the dimensions of the motor h,b,\u03a8 , linearization is possible. Therefore, the displacements of the motor feet on the left side zfL, yfL and on the right side zfR, yfR can be described by the displacements of the stator mass zs, ys, \u03d5s [9]: zfL = zs \u2212 \u03d5s \u00b7 b; zfR = zs + \u03d5s \u00b7 b (1) yfL = ys \u2212 \u03d5s \u00b7 h; yfR = ys \u2212 \u03d5s \u00b7 h (2) With the geometrical coherences of the rotor mass system, za = zv + a\u0302 \u00b7 cos(\u03a9 \u00b7 t + \u03d5a); ya = yv + a\u0302 \u00b7 sin(\u03a9 \u00b7 t + \u03d5a) (3) zm = zw + e\u0302m \u00b7 cos(\u03a9 \u00b7 t + \u03d5m); ym = yw + e\u0302m \u00b7 sin(\u03a9 \u00b7 t + \u03d5m) (4) zu = zw + e\u0302u \u00b7 cos(\u03a9 \u00b7 t + \u03d5u); yu = yw + e\u0302u \u00b7 sin(\u03a9 \u00b7 t + \u03d5u) (5) the linear inhomogeneous differential equation can be derived, described by mass matrix M, damping matrix D, stiffness matrix C, coordinate vector q, and excitation vector f: M \u00b7 q\u0308 + D \u00b7 q\u0307 + C \u00b7 q = f (6) Mass matrix M and coordinate vector q are described by M = \u239b \u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239d ms 0 0 0 0 0 0 0 0 0 mw 0 0 0 0 0 0 0 0 0 ms 0 0 0 0 0 0 0 0 0 mw 0 0 0 0 0 0 0 0 0 \u0398sx 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 \u239e \u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u23a0 ; q = \u239b \u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239d zs zw ys yw \u03d5s zv zb yv yb \u239e \u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u23a0 (7) Damping matrix D is described by D = \u239b \u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239d 2dfz 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 2dfy 0 \u22122dfyh 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 \u22122dfyh 0 2(dfyh 2 + dfzb 2) 0 0 0 0 0 0 0 0 0 2dzz \u22122dzz 2dzy \u22122dzy 0 0 0 0 0 \u22122dzz 2dzz \u22122dzy 2dzy 0 0 0 0 0 2dyz \u22122dyz 2dyy \u22122dyy 0 0 0 0 0 \u22122dyz 2dyz \u22122dyy 2dyy \u239e \u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u23a0 (8) Stiffness matrix C is described by C = \u239b \u239c\u239c\u239c\u239c\u239c\u239c\u239d 2(cfz + cbz) \u2212 cm cm 0 0 0 0 \u22122cbz 0 0 cm c \u2212 cm 0 0 0 \u2212c 0 0 0 0 0 2(cfy + cby) \u2212 cm cm \u22122cfyh 0 0 0 \u22122cby 0 0 cm c \u2212 cm 0 0 0 \u2212c 0 0 0 \u22122cfyh 0 2(cfyh2 + cfzb 2) 0 0 0 0 0 \u2212c 0 0 0 2czz + c \u22122czz 2czy \u22122czy \u22122cbz 0 0 0 0 \u22122czz 2(czz + cbz) \u22122czy 2czy 0 0 0 \u2212c 0 2cyz \u22122cyz 2cyy + c \u22122cyy 0 0 \u22122cby 0 0 \u22122cyz 2cyz \u22122cyy 2(cyy + cby) \u239e \u239f\u239f\u239f\u239f\u239f\u239f\u23a0 (9) Vector f, describing the excitation vector, is split into its components for each single source of excitation\u2014index u for eccentricity of rotor mass, index m for magnetic eccentricity and index a for bent rotor deflection" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003117_ichr.2009.5379588-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003117_ichr.2009.5379588-Figure3-1.png", "caption": "Fig. 3. Model of the door.", "texts": [ " Assuming the humanoid robot to work in humans' office environments, we chose to consider the manipulation of a swing door. Differently from regular doors used in houses, a swing door is heavier and loaded with a spring set in its hinge. The main aspects to be considered while performing the task are: 978-1-4244-4588-2/09/$25.00 \u00a92009 IEEE 134 Finally we report the set of differential equations describing the whole system, whose solution is implemented by the path planner using 4t h order Runge-Kutta method: (I) (2) F apfx + F intx - D Xr u, Fapfy + F inty - D Yr u ,tir angle is denoted as ed. As shown in Fig.3, we chose to model the door with the same shape as the manipulable area of the robot to simplify the computation of the interaction force, P int, and to guarantee continuous contact at the edge. The dynamic behavior of the spring-loaded door is defined by the position gain K p and by the velocity gain K v ; the mass and momentum of inertia are M d and Id . The interaction force Pint is computed according to the spring-dumper model in (5) where Cp and C; are the position and velocity gain respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000471_j.wear.2003.11.001-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000471_j.wear.2003.11.001-Figure1-1.png", "caption": "Fig. 1. Cross-sectional scheme of the cladding head.", "texts": [ " E-mail addresses: takacs@kgtt.bme.hu (J. Taka\u0301cs), toth@kgtt.bme.hu (L. To\u0301th), friedrich.franek@tuwien.ac.at (F. Franek), pauschitz@user.ifwt.tuwien.ac.at (A. Pauschitz), sebi@bzaka.hu (T. Sebestye\u0301n). As base material we used selective-laser-sintered phosphor bronze specimens for the first series of tests. These specimens were turned, grinded and degreased before the tests. The laser-coated specimens\u2014after grinding and degreasing in alcohol\u2014were coated by a TRUMPF TLC 5000 CO2 laser using a special coating head (Fig. 1). We used 1000 W laser power and 500 mm/min feed. Ar as shield and carrying gas was used during spraying. Both types of specimens were sprayed with a Co-based (C 0.76/Mn 0.31/Fe 3.13/Ni 13.13/Cr 19.23/W 7.75/B 1.79/Co balance) and an Fe-based (FeB) (B 19/Si 2.11/C 0.44/Fe balance) powder. Particle diameter of the powders was 45\u201375 m. We chose the one-step technique when the powder is blown through the laser beam into the laser-melted pool on the surface. The particles do not melt completely flying through the laser beam, the almost complete melting takes place in the melt pool" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003789_ijtc2011-61146-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003789_ijtc2011-61146-Figure5-1.png", "caption": "Figure 5: Photo of micro motor drive and PM-integrated rotor", "texts": [ " Figure 4 shows optical microscope images of Design II elastic foundation made through X-ray lithography and electroplating. A test rig was designed with a brushless permanent magnet (PM) DC motor drive [7] developed by Power Electronic Systems Laboratory, ETH Zurich, Switzerland. The original PM rotor with two ball bearings had the first critical speed of 198,840 rpm according to [8] (including an axial turbine). To implement both radial foil bearings and thrust air bearings, rotor was redesigned (Figure 5(b)) with a thrust disc with 12 mm diameter with higher bending critical speed above 1,000,000rpm. Monolithic stator housing was designed to house both the motor stator and front radial foil bearing. Rear part of the stator housing was designed to accommodate front air thrust bearing (Figure 6) and thrust spacer. Separate rear housing contains both rear air thrust bearing and rear radial foil bearing (Figure 7). The air thrust bearings have Rayleigh steps of 10 \u00b5m high, formed through selective electroless Ni coating" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002607_978-0-387-77747-4_6-Figure6.4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002607_978-0-387-77747-4_6-Figure6.4-1.png", "caption": "Fig. 6.4 Schematic cross-section of MEMS air turbine with axial flow hydrostatic gas journal bearings", "texts": [ " Assuming a plain annular seal with an off-centered shaft and imposing an axial pressure difference across the seal, the pressure drop due to the inlet loss into the seal will be higher at the large clearance than at the small clearance due to the increased local flow rate. Thus, the axial pressure gradient inside the seal will be lower at the large clearance compared to the axial pressure gradient at the small clearance. The difference in the pressure gradients provides a restoring force which opposes the shaft displacement, yielding a potentially large direct stiffness. In summary, the so-called \u201cLomakin\u201d effect stems from the combined effect of the inlet loss of the axial flow into the seal and the axial pressure gradient across the seal. Figure 6.4 depicts the conceptual arrangement of the hydrostatic gas journal bearing radially supporting the rotor of the MEMS device shown in Fig. 6.2. The stiffness and damping coefficients of the bearings are indicated by springs and dashpots. The goal of the analysis is to appropriately describe and to model the stiffness and damping coefficients which can then be used in a rotordynamic analysis. To illustrate the dynamic behavior only the radial motion is considered here. Coupling phenomena between the axial and radial motion will be discussed later" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000131_bf02903530-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000131_bf02903530-Figure5-1.png", "caption": "Figure 5 Attaching the shaft close to the face, as shown in (a), strongly opposes putter-head rotation about the X-axis and reduces vertical gear-effect. Preferably, the shaft tip should be connected in line with the centre of mass, as shown in (b)", "texts": [ " Shaft coupling modifies the theory Initial qualitative tests on an early prototype putter gave encouraging confirmation that high topspin could be achieved, but more precise quantative tests using time-elapsed photography uncovered a major problem. The image-capture spin measurements turned out to be very consistent, but only about half the value predicted by theory. The reason for this was not immediately evident and one or two \u2018thought to be possible\u2019 causes were investigated. It then became apparent that the vertical gear-effect was being impeded by the axial stiffness of the shaft. With the shaft connected at the usual position forward on the putter head (as shown in Figure 5a), the topspin was only 50% of that obtained when the shaft connection was repositioned back from the impact face to align with the COM (Figure 5b). Thus, shifting the shaft connection to a point close to the heel\u2013toe axis (i.e. the desired rotation axis for vertical gear-effect) restored the imparted topspin to the value predicted by theory. To explore this phenomenon, an experimental putter head was made and tested. This is shown schematically in Figure 6 and comprises a 120 mm length of 32 mm square section aluminium with a standard putter shaft bonded into a 20 mm deep bore, 20 mm from the heel end, in the vertical centre-plane. Rectangular axes shown in Figure 6 follow the same system adopted in Figure 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002885_wnwec.2009.5335810-Figure15-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002885_wnwec.2009.5335810-Figure15-1.png", "caption": "Figure 15. FEA modeling after setting convective exchange coefficient", "texts": [ " 14, the red section denotes the torque applied in the spline; the blue section denotes the displacement constraints applied in the bearing. The temperature constraints were applied in the yellow section and the gradient temperature field was applied in the contact area of wheel and pinion. The atmosphere temperature is 25 . Other thermodynamics coefficients are as below: the thermal conductivity is 1.35E-3W/(m\u00b7K) and the thermal expansion coefficient is 1.12E-5/K. The convective exchange coefficient is set as Fig. 15. Analysis results of the bending and torsion deformations when considering thermal-mechanical coupling are listed in Tab.6. From Tab.6, we can see that: the bending deformations of the wheel and pinion are affected little when considering thermal-mechanical coupling. But comparatively speaking, it has almost no influence on the bending deformations of the wheel. However, both the bending deformation at the noncontact tooth and the maximum torsion deformation of pinion decrease slightly. It has little influence on the torsion deformation at the contact tooth, but it has great influence on the torsion deformation at the non-contact tooth when thermalmechanical coupling is considered" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002541_1.2967884-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002541_1.2967884-Figure1-1.png", "caption": "Fig. 1 Thread cross section along the bolt axis", "texts": [ " ournal of Pressure Vessel Technology Copyright \u00a9 20 om: http://pressurevesseltech.asmedigitalcollection.asme.org/pdfaccess.as coarse or a fine pitch. Then, the equivalent diameter is defined as the diameter of a circle that has the same area as the true cross section. The equivalent diameter is compared with such conventional diameters as stress area diameter and pitch diameter. 2 Analytical Expression of the True Cross Sectional Area of Screw Threads 2.1 Mathematical Expression of the Thread Profile. Figure 1 illustrates the cross section of the external screw thread along the bolt axis. Assuming that the thread root radius is part of a single circle with radius , the thread profile can be defined by dividing the cross section into three parts, namely, \u201cA-B: root,\u201d \u201cB-C: flank,\u201d and \u201cC-D: crest.\u201d The cross section profile of the external thread perpendicular to the bolt axis is obtained by expanding the thread configuration of one pitch height into the plane, as shown in Fig. 2. It is worth noting that the cross section profile perpendicular to the bolt axis is identical at any position along the axis and can be expressed in terms of radius r shown in Fig. 1. Figure 3 a shows the geometry around the thread root in detail. The coordinate of a point on the arc BAB that forms the thread root radius is expressed as r , P /2 . The root radius has an upper limit restricted by a minor diameter d1. Figure 3 b displays the case of the maximum root radius max. The radial coordinate of the points B or B , where the thread flank becomes the tangent of the arc, coincides with that of minor diameter d1. Considering the symmetry between 0 and \u2212 0, the cross section profile shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001470_sice.2006.315455-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001470_sice.2006.315455-Figure7-1.png", "caption": "Fig. 7 Trajectory tracking control.", "texts": [], "surrounding_texts": [ "This section explains about an obstacle avoidance based on the obstacle information which got in the Section 3.We consider a trajectory generation and trajectory tracking, to let the robot avoid the obstacle. 4.1 Trajectory generation In this section, we consider a trajectory generation in order to let the robot avoids the obstacle. About the trajectory generation, many methods have been proposed, for example, based on splines[7] or optimal control[l]. In this paper, we run on the mobile robot hypothetically[2]. And evaluating the given trajectory, we choose an optimal trajectory. If the robot detects a obstacle at (Xob, Yob) that closer to the robot than the safety distance dOA, the control of the rotation velocity of the robot is W = COA WOA (d)(OLob ObR) where 1 2 2 d = /(XR Xob) + (YR Yob) WOA (d) = /d2 and c/ob = 1i:- + arctan 2 (Yob -YR,Xob- XR)2 Hence, COA is a constant weight, and COA is choose experimentally. In addition dOA is given by dOA dob + dsafe where dOB is the size of the obstacle which goat in section 3,?and dsafe is a constant weight. On the other hand, if the robot leaves the obstacle enough, the control of the rotation velocity of the robot is given by W = bg -R where )ob = arctan 2 (yg -YR, Xg -XR) If we generate a trajectory, using the above way, two trajectories are generated due to the choice of plus or minus in eq. (3) as shown in Fig. 6. In this paper, when plus is choose, the trajectory is called positive trajectory, on the other hand, if minus is choose, the trajectory is called negative trajectory. We express the generated trajectory y = f(x). To decide the optimal trajectory, we treat a curvature of the trajectory. The curvature is given by f\"(x) (1 + (f/(X))312 where f(.) and f(.) denote first and second derivatives, respectively. For we suppose that it is difficult for the robot to realize the trajectory which has a large curvature, we choose the trajectory which has a smaller curvature, that is, we are minimizing the function as follows rx9 J j/ f (x) dx. For example, as for the trajectory of Fig. 6, J of the positive trajectory is 3.14, on the other hand the negative trajectory has J with 6.86. Therefore, we choose the positive trajectory as the optimal trajectory. 4.2 Trajectory tracking control We explain about a trajectory tracking control. A purpose of this control is that the robot track to the reference trajectory without error. We set a reference point (Xd, Yd) on the reference trajectory, the control values of the robot VR, WR are given by VR = . de WR= k ec+ Od (3) where -y, k are positive, and these parameter are set ex- perimentally. In addition de = e + x = d R, eC = Yd-YR, e- = Od-OR and Od = arctan 2(ey, ex). Reference point (Xd, Yd)" ] }, { "image_filename": "designv11_61_0002564_s10015-008-0560-2-Figure13-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002564_s10015-008-0560-2-Figure13-1.png", "caption": "Fig. 13. Motion of manipulator", "texts": [ " Figure 12 shows the experimental response, under the condition that initial position is (\u03b81i = \u03b82i = \u2212\u03c0/4, \u03b83i = 0), the fi nal position is (\u03b81f = \u03c0/2, \u03b82f = \u2212 \u03c0/4, \u03b83f = \u03c0/4), and the working time is 0.8 s. Figure 12a, b show the experimental response of angular velocity of motor which are calculated by angular displacement measured by rotary encoder. Broken line is theoretical result, and solid line is experimental result. In Figure 12c, it is shown that link 3 is arrived at desired position. And the motion is caused by inertia force of link 1 and 2. And, Figure 12d, e are applied voltage and electric current of motor 1. The theoretical result is similar to the experimental one. Figure 13 shows a strobe light image of the motion. The light fl ashes occurred every 0.02 s. The theoretical result (a) is similar to the experimental result (b). From these results, it is confi rmed that the modeling is effective. 5 Conclusions The results obtained in this work are summarized below. 1. It is considered that the trajectory for energy saving of manipulator with passive revolute joint is available for the work in vertical plane. 2. From experimental results, it is considered that modeling for simulation is effective" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000848_1.1866162-Figure15-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000848_1.1866162-Figure15-1.png", "caption": "Fig. 15 Bearing pressure field when the applied load is maximum \u201e0 deg rotation angle\u2014reference case\u2026", "texts": [], "surrounding_texts": [ "Downloaded From: http:/ Table 4 Main parameters for different load offsets \u201ebearing 1\u2026 Table 5 Main parameters for different applied loads \u201ebearing 1\u2026 l. 127, APRIL 2005 Transactions of the ASME /tribology.asmedigitalcollection.asme.org/ on 01/30/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use" ] }, { "image_filename": "designv11_61_0003643_ictee.2012.6208635-Figure26-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003643_ictee.2012.6208635-Figure26-1.png", "caption": "Fig. 26- Auto CAD model showing all rectangular faces of the square prism opened", "texts": [], "surrounding_texts": [ "A square prism edge of base 30 mm and height 50 mm is given .The prism is having its base in H.P. with two edges of base perpendicular to V.P. .Draw the development of lateral surface of prism. Explanation tools and their application \u2022 Auto CAD model shown in fig18,fig 19, fig 20,fig 21&fig 22 are used to explain three dimensional concepts \u2022 Drafting video explains the construction method \u2022 Animated power point presentation is used to display all writing work live Fig. 23-Auto CAD model shows the square prism whose surface has to be developed reference planes make visualization very clear cult to show reference planes and projections Drafting videos can be played at desired speed and at desired point Making drawing on board by instruments is time consuming and laborious Animated power point presentations increase the legibility and saves time. Writing on board is also good but consumes time Screen cast video lecture series can be developed in Camtasia at very low cost No possibility VI. CONCLUSION Comparison shown in table 3 helps us to conclude that modern teaching style should be adopted for teaching Engineering Drawing. A centralized agency should help in development of the tool for teaching engineering drawing and distribute it Scope for further work \u2022 Explore possibility of developing 3D animations in 3Dmax ,Pro E etc. \u2022 Explore possibility of better solid modeling software \u2022 Develop a video lecture series on Engineering Drawing ACKNOWLEDGEMENTS The author would like to thank, Dr Anil Kothari Prof. U.I.T RGTU Bhopal and Shri Vinay Thapar Prof. U.I.T RGTU Bhopal, Dr S.C Kapoor Director NIIST Bhopal ,Shri D. Subodh Singh Chairmen NRI group of institutions Bhopal for their kind support and guidance." ] }, { "image_filename": "designv11_61_0000267_papcon.2002.1015139-Figure14-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000267_papcon.2002.1015139-Figure14-1.png", "caption": "Figure 14: Typical Aluminum Die Cast Skewed Rotor", "texts": [ "96 percent 459 Effecfs of Unbalanced Volfages on the Performance of Polyphase lnduction Mofors (NEMA 74.36): When the line voltage applied in a polyphase induction motor are not equal, unbalanced currents in the stator winding will result. A small percentage voltage unbalance will result in a much larger percentage current unbalance. Consequently, the temperature rise of the motor operating at a particular load and percentage voltage unbalance. will be greater than for the motor operating under the same conditions with balanced voltages. When the derating curve of Figure 7 (NEMA Figure 14-1) is applied for operating on unbalanced vonages, the selection and setting of the overload device should take into account the combination of the derating factor applied to the motor and increase in current resulting from the unbalanced voltages. This is a complex problem involving the variation in motor current as a function of load and voltages unbalanced. When unbalanced voltages are anticipated, it is recommended that the overload devices be selected so as to be responsive to Imaximum in preference to overload devices responsive to laverage *FOOTNOTE: Frequently the operator does not know what the actual load is, nor can he control it", " There are a variety of other methods to predict winding temperature with respect to change in load. However, most of them are not very accurate for significant changes in load or at the extremes. The mathematics can be quite complicated and is usually not practical for fieid use. For a more detailed study of the issue, see reference 4 of this paper (pages 301 through 307) written by Richard Nailen. ROTOR TEMPERATURE CONSIDERATIONS The typical rotor less than 500 hp is die-cast with aluminum and has no air ducts. The rotor fan blades are normally cast as part of the end rings as shown in Figure 14. For these size motors, most will have a skewed rotor to improve acceleration and noise characteristics. A study of the thermal stability of the rotor involves these elements: the rotor lamination surface, the rotor bars and the rotor end rings. Table 9 shows the relationship of the elements for starting and stall conditions. Usually under running and overload conditions the stator winding will be the limiting factor and not the rotor. Under stall conditions it may be the rotor, especially if the motor is a two pole or four pole design" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001932_s00170-007-1295-2-Figure10-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001932_s00170-007-1295-2-Figure10-1.png", "caption": "Fig. 10 A gear manufactured by the SLS/CIP/HIP process", "texts": [ " So it testifies that the mechanical performance of hot isostatic pressed specimens is better than that of vacuum sintering specimens. \u03c3 \u00bc \u03c30 \u00fe Kd 1=2 \u00f04\u00de where \u03c30 and K are two different constants; \u03c3 is the yield strength; and d is the diameter of the crystal grain. As shown in Table 4, mechanical performances of HIPped AISI304 specimens are close to that of solution treatment AISI304 specimens. Consequently, dense metal parts with better mechanical performances can be manufactured by SLS/CIP/HIP. Figure 10 shows the gear manufactured by SLS/CIP/HIP, whose relative density is near to 98%. AISI304 selective laser sintering (SLS) green specimens are successfully fabricated by indirect SLS using bisphenol A type epoxy resin as the binder. Airproof canning has been easily achieved by the polyreaction of RTV-2, and SLS metal parts are not polluted by canning used in the SLS/ cold isostatic pressing (CIP)/hot isostatic pressing (HIP) process, so liquid RTV-2, Si(OEt)4, Bu2Sn(OCOC11H23)2, and CH3COOCH2CH3 are suitable for CIP canning raw materials" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001870_s11465-008-0002-9-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001870_s11465-008-0002-9-Figure1-1.png", "caption": "Fig. 1 Linear 2-DOF vehicle model", "texts": [ " Some simplifications are as follows: 1) the front-wheel angle is taken as the input and neglecting the influence of the steering system; 2) the suspension is neglected, namely the displacement along the z-axis, the elevation revolving the y-axis and the obliquity revolving the x-axis are zero; 3) the vehicle takes linear motion at a constant speed; 4) the sideslip characteristic of a tire is in a linear range; 5) the influence of the air is neglected; 6) for left and right wheels, the change of the tire characteristic due to the change of load is neglected. The effect on the outer obliquity by the aligning torque of a tire is also neglected. The result of the simplifications shows that the vehicle only has the two-degree side motion, along the y-axis and the yaw motion revolving the z-axis (Fig. 1). The centroid of a vehicle is taken as the origin of the vehicle coordinate; therefore, the parameters that are related to the distribution of mass, such as the moment of inertia, are constant. Some parameters in this article are defined as follows: M is the mass of the vehicle; IZ is the moment of the inertia revolve z-axis; v is the speed of the centroid; li is the distance between the centroid and the ith-axle (the direction of the vehicle motion is in positive direction); di is the steering angle of the ith-wheel; ki is the synthetic cornering stiffness of the ith-wheel; Lij is the distance between the ith and jth-wheel (the direction of the vehicle motion is in positive direction); Li is the distance between the ith-axle and the steering center line; u is the branch of the centroid speed on the x-axis; v is the branch on the y-axis of the centroid speed; vr is the angular velocity of the vehicle revolving the z-axis; b is the sideslip angle of the centroid; and kpi is the ratio of the ith-steering angle to front-wheel steering angle" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000902_roman.2004.1374761-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000902_roman.2004.1374761-Figure3-1.png", "caption": "Fig. 3 Experimental Field", "texts": [], "surrounding_texts": [ "I . INTRODUCTION There is much research on human-robot symbiosis. Therefore, many researchers have been studying humanaffinity and maneuverability of the robots. For example, Miwa has been studying robots with human-like faces [ 13. He investigated robot functions necessary for human-robot communication. Goetz has reported that robot feature and behavior strongly affected human-robot interaction [2]. 0-7803-8570-5/04/$20.00 02004 IEEE We propose a new approach that is different from these other researches. Through interaction experiments between rats, less sophisticated animals, and robots, we would like to create a model of symbiosis between creatures and robots. We have been conducting several interaction experiments based on animal psychology. On animal psychology, a lot of research has been conducted, and effective results have been reported [3]. These studies have contributed to clarifying human abilities for leaming and cognition. Since 1995, we have been developing experimental setups, and conducting interaction experiments. In 2003, we developed a robot (WM-6) with two levers (Fig. 1) and experimental setups (Fig. 2) that were able to operate without time restriction. Using these setups, we conditioned rats to push the levers on the robot to obtain food [4]. It was very difficult for the rats to learn to push the levers. Therefore, to show rats how to obtain food and how to push the levers, the experimenter pushed the levers on the robot and fed food pellets in front of the rats. In this paper, we tried to condition rats automatically to push the levers on the robot without experimenter showing. Refereeing to \u201coperant conditioning experiment\u201d conducted by Skinner (1904-1990), we developed a robot behavior generation algorithm that enabled the robot to autonomously show rats how to obtain food, pushing the levers. We considered that the rats which were conditioned using this algorithm could learn to push the levers faster - 229 - than the rats which were not. Therefore, we named this algorithm \u201cAccelerating Rat\u2019s Learning Speed Algorithm\u201d. We built it into the Control-PC shown in Fig. 2 (\u201cOperation generator\u201d). We then conducted an experiment and confirmed effects af this algorithm. Through this experiment, we propose a method that the robot autonomously shows creatures its functions. Weight g II. RELATED WORKS Skinner was a famous psychologist and the originator of behavior analysis. He had been studied \u201coperant conditioning\u201d [5]. Operant conditioning involves the modification of behavior by the reinforcing or inhibiting effect of its own consequences. He developed the \u201cSkinner box\u201d to conduct the operant conditioning experiments. In this box i(300 x 250 x 250 [mm]), a food feeder and a lever were set on a wall. The food feeder fed a food when the subjects pushed the lever or when the experimenter pushed the button outside of this box. In the operant conditioning experiment, he conditioned a rat to push the lever to obtain food. Pushing levers is not innate behavior of rats, and hence a difficult task for them to learn. Therefore, Skinner divided the learning process into some small steps. He was then changing these steps depending on the progress of rat\u2019s learning and succeeded in conditioning the rat that had never spontaneously pushed the lever. On this conditioning process, Skinner conditioned the rat to push the lever to obtain food. Today, this procedure of operant conditioning is very popular in psychology. 540 111. EXPERIMENTAL SETUPS Length mm 170 Width based on \u201cAccelerating Rat\u2019s Learning Algorithm\u201d are built in operation generator. According to these patterns, the operation generator generates a robot path and controls other setups. The robot controller controls robot movements using visual feedback. B. WM-6 We developed a small mobile robot with two levers and named it WM-6 (Waseda Mouse No. 6). The performance of this robot is almost equal to a mature rat as shown in Table 1. In addition, this robot is covered in plastic as shown Fig. 1. This robot is a powered steering robot, which has two DC motors to drive two wheels (right and left). Moreover, this robot has an onboard processor Microcontroller PIC, a Bluetooth wireless communication module and two Li-ion battery. Thus, this robot is controlled remotely by the Control-PC. In addition, this robot has two levers with touch sensors to detect the rat pushing the levers during interaction. WM-6 can two-way communicate with the Control-PC using Bluetooth. Each motor is controlled by the PIC according to instructions from the Control-PC. In addition, the PIC measures battery voltage and checks the state of the levers as being on or off, and sends this information to the Control-PC in real time. Therefore, it is possible to monitor the battery power on the Control-PC. WM-6 has two batteries, where one is main battery (Li-ion, 7.2 V, 1500 [mAh]) and the other is a standby battery (Li-ion, 3.6 V, 500 [mAh]) for Bluetooth module. This robot can maintain its operation for 2 hours with a fully charged main battery. Moreover, the robot can maintain its operation independently of battery capacity using the battery exchanger (as described latter). During the dead battery is exchanged, the robot uses the standby battery to maintain communication link via Bluetooth. C. Other Setups When the food feeder receives an instruction form the Control-PC, this machine feeds a food pellet (45 mg) into a feeding point (a plastic bowl) on the field. - 230 - When the water feeder receives an instruction from the Control-PC, this machine extends a mouthpiece connected to a water bottle into the field for three seconds. During these seconds, rats can obtain water. The battery exchanger is set on a wall of the field. This machine consists of an arm for exchanging batteries and a charger set outside the field. When the battery on WM-6 is dead, the robot moves to the front of the battery exchanger. Once the robot arrives, the exchanger starts its operation. Then the arm exchanges the batteries from the robot and the changer. This operation takes about one minute to complete. After the operation, the robot retums to the task that it was performing. The battery that was put into the charger charges until the next battery exchange. Thus, we can conduct the interaction experiments for long time independently of battery capacity. Iv. AUTOMATIC LEARNING SPEED ACCELERATION In the experiment that was conducted in 2003, we create a condition that the rats could not obtain food without pushing the lever as shown in Fig. 4. In other words, Fig. 4 shows the operational pattern of the experimental setups or robot\u2019s functions in this experiment. In this pattem, a rat could obtain a food by pushing the levers on the robot and moving to the front of the food feeder with the robot. In this paper, we tried to automatically condition a rat to push the levers on the robot to obtain food. We then considered that it is possible to accelerate rat\u2019s leaming speed by changing the robot behavior pattern step by step. Referring to the operant conditioning experiment conducted by Skinner, we developed a robot behavior generation algorithm that enabled the robot to autonomously show rats how to use its hnctions. It is harder for rats to leam to push the levers on the robot than on the wall near the feeding point. However, the robot has the advantage of mobility to gesture its functions over the lever on the wall. We then divided the learning process into three \u201cLevel\u201d Each \u201cLevel\u201d includes robot behavior pattern, that is to say Rat; touch the levers? -- +l t the food feedin the operational patterns of the experimental setups. Therefore, the robot autonomously conditioned the rats to push the lever by changing the \u201cLevel\u201d depending on how much the rats had leamed. I ) Level I : Level 1 corresponds to \u201cAdaptation\u201d and \u201cMagazine training\u201d as seen in Skinner\u2019s experiment. The purpose of Level 1 is to reinforce rat\u2019s motivation for movement. During this level, the food feeder routinely feeds ten food pellets every hour. In our previous experiments, we observed that the rat rarely moved in the experimental field and preferred to stay in each corner. However, we confirmed that the rats that had obtained food in the field moved actively than those that had not. Therefore, we considered that the routine feedings could reinforce the rat\u2019s motivation for movement. When we observe the rat moves actively, Level 1 is finished and Level 2 is started. 2) Level2: Level 2 corresponds to the beginning half of \u201cShaping\u201d as seen in Skinner\u2019s experiment. The purpose of Level 2 is to condition the rat to approach the robot. During this level, the robot routinely moves to the front of the food feeder ten times every hour. Every time the robot arrives there, the food feeder feeds a food pellet. Therefore, the robot autonomously shows the rat the connection between its movement and food feeding. It is then expected that the rat would be interested in the robot and hence approach it. The rat\u2019s approach to the robot is detected using the image processing, which worked when the rat moves within a radius, r from the robot. When the approach is detected, the robot moves to the front of the food feeder. To reinforce the rat\u2019s approach, the food feeder then feeds a food when the robot arrives there. After we observe the rat approaches the robot a lot of times, Level 2 is finished and Level 3 is started. 3) Level3: Level 3 was the final level and corresponded to the latter half of \u201cShaping\u201d as seen in Skinner\u2019s experiment. The purpose of Level 3 is to condition the rat to push the levers on the robot. During this level, the approach detection area narrowed every time the rat approached the robot. Therefore, the radius r (radius of approach detection area) was decreased every time the rat approached the robot. It was then expected that the rat would approach the robot and occasionally push the levers on the robot. When the rat pushes the levers, the robot moves to the front of the food feeder and the feeder feeds a food pellet. When we observe the rat repeatedly pushes the levers to obtain food, we finish this automatic conditioning. V. EXPERIMENT A. Experimental Procedure We conducted an experiment to confirm the effect of \u201cAccelerating Rat\u2019s Leaming Speed Algorithm\u201d. We used two male albino-rats, 70 weeks old, as the subjects. One was an experimental subject that was conditioned by this algorithm, and the other was a control subject that was not. This experiment had been conducted until the rat learned to push the levers. The experiment was divided into some trials, with a single trial being 24 hours. To keep the rats hungry at the start of each trial, it was conducted every 3 or 5 days. Before each trial of experi,mental subject was started, we selected the level in this trial depending on the result of the previous trial (as described in the following result section). During one trial, we had not changed the level that was first selected. In whole trials of control subject, the experimental setups operated in accordance with Fig. 4 which is that the rolbot never moved autonomously. In this experiment, we conducted 8 trials using the experimental subject and 5 trials using the control subject. B. Result The distances that the experimental sub 'ect and the robot moved every ten minutes in the Is', 4' , 5'h, and 7'h trials are shown in Fig. 5, 6, 7 and 8. In addition, the numbers of that the experimental subject pushed the lever and approached the robot in the 4'h, 5\", and 7'h trials are shown in Fig. 9, 10 and 11. The distance that the control subject and the robot moved every ten minutes in the 1\" and 5'h trials are shown in Fig. 12 and 13. In addition, Table 2 shows the level, the total distance the subjects moved, the total numbers of the subjects approached and pushed the lever, the change of the weight between the start and the end of the trial, and the correlation coefficient between the transition of distances that the subject and the robot moved in each trial. In addition, the cumulative records of the experimental subject that moved, pushed the lever and approached the robot are :shown in Fig. 14. Therefore, this shows the leaming process of the experimental subject. t? C. Consideration At first, we described the process of the experimental subject. We conducted the 1'' trial without \"Accelerating Rat's Leaming Speed Algorithm\" and confirmed this subject had not leamed to push the levers. We then conducted the 2\"d trial and selected Level 1. However, this subject had never moved actively in this trial due to the subject having not found the feeding point through the trial. We then conducted the 3rd trial and selected Level 1 again. At the start of this trial, we put the subject near the feeding point. We observed that the subject moved actively in the experimental field and obtained food just after the feeding. Therefore, we conducted the 4lh trial and selected Level 2. Fig. 6 shows the result of this trial. According to the robot routine movement, peaks of the robot movement were shown every single hour. While the robot was moving, the subject had been watching the robot. During the latter half of this trial, the subject approached the robot after every robot routine movement as shown Fig. 10. We then conducted the 5\" trial and selected Level 3. At the beginning of this trial, the subject approached the robot repeatedly. However, this approach could not be confirmed as the approach detection area narrowed. Therefore, the subject never approached the robot during the latter half of this trial, as shown Fig. 10. We considered the cause of this effect was shortage of the reinforcement for the approach. Thus, we conducted the 6'h trial and selected Level 2 again. On this trial, the subject had been approaching the robot repeatedly again. Therefore, we conducted the 7'h trial and selected Level 3. During this trial, as the approach detection area narrowed, the subject began to push the lever on the robot as shown Fig. 1 1. We then conducted the gth trial without the \"Accelerating Rat's Learning Speed Algorithm\" in the same way as the 1\" trial. On this trial, the subject had been pushing the robot lever. We had however, been conducting five trials using the control subject. Nonetheless, we could not confirm any changes on the behavior of this subject throughout these trials. The subject had been staying at each comer of the experimental field and never pushed the lever on the robot. In this experiment, the experimental subject could leam to push the robot lever, while the control subject could not. Therefore, we confirmed the \"Accelerating Rat's Leaming Speed Algorithm\" was effective. Due to only one subject being examined, it is necessary for us to conduct the same experiments using more subjects. However, according to the fact that rat's ability for leaming is uniform, we believe this algorithm is effective for almost all rats. In addition, the leaming ability of rats reaches its peak between 10 and 20 weeks old. After that it declines with age. Therefore, we hope that younger rats could leam the robot functions faster than the subject used in this experiment. It is necessary for us to confirm this however. 35 -Approach to the robot 1 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 ~ 0 0 0 0 0 0 0 0 0 ~ 0 0 0 \" e m w r - m m o - z , ~ Time min Fig. 6 Movement ofthe experimental subjeci in the 4 trial 40 E 35 8 30 +d 9 25 5 20 S 15 v) M 5 10 5 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 ~ m b m w r - w m o ~ m c f ~ 0 0 0 0 0 0 0 0 0 ~ 0 0 0 - - e - * Time min Fig. 7 Movement of the experimenial subject in the 5 I h irial 40 E 35 8 30 U 9 25 3 20 v) 00 .? 15 5 I O 5 0 r ............................................................................................................................................... I 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 - m m b m a r - c a m o - m m + - - - - c Time min Fig. 8 Movemeni OJ the e.uperiniental suhject in the 7Ih trial VI. CONCLUSION In this paper we proposed a method that the robot autonomously showed creatures its functions. We tried to automatically condition the rat to push the levers on the robot to obtain food. We then developed a robot behavior generation algorithm that enabled the robot to autonomously show the rats how to obtain food, pushing the levers. We named this algorithm \"Accelerating Rat's Leaming Speed Algorithm\". We conducted an experiment to confirm the effect of this algorithm. We then compared the learning process of 40 r ....................................................................................... I I 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 N M b m n l D P - W O \\ O N M b ~ 0 0 0 0 0 0 0 0 0 ~ 0 0 0 - - - - - Time min Fig. 9 Numbers of the experimental subject approached in the 4'\" trial 40 35 g 30 E .z 25 c- ; 20 e 5 15 I O 5 0 ....................................................................................................................................................... * .Leverpush 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 \" X R ~ z g P 2 g 8 \" ~ R ~ - - - c - Time min Fig. I I Number of'the experimental subject pushed the lever and upproached in the 7'\" trial - 233 - two subjects, one is conditioned by this algorithm and the other is not. We then confirmed that the subject that was applied the algorithm could only leam to push the levers to obtain food. component that enabled the algorithm to select the \u201cLevel\u201d automatically. It is necessary for robots which aim to live symbiotically with creatures to have abilities or functions to show the creatures their functions. Through the experiment, we confirmed that using operant conditioning it is possible that the robot autonomously shows creatures its functions. The algorithm that was developed by us in this paper is not widely applicable. Therefore, we would like to develop a widely applicable robot behavior generation model that enabled the robots to autonomously show or condition creatures. ACKNOWLEDGMENT Part of this research was conducted at the Humanoid Robotics Institute (HRI), Waseda University. The authors would like to thank to Professor Hiroshi Kimura, Waseda University for his support. Finally, we would like to thank to Solidworks Corp., Advanced Research Institute for Science and Engineering of Waseda University, for their support. REFERENCES [ I ] H. Miwa, T. Okuchi, H. Takanobu, A. Takanishi. \u201cDevelopmeni of a New Human-like Head Robot WE-4,\u2019\u2019 Proc. of the 2002 IEEE/RSJ Iniel. Conference of Intelligent Robots and Systems. 2002, pp.2443-2448. J. Goeiz, S. Kiesler, A. P0wer.s. \u201cMaiching Roboi Appearance and Behavior io Task io Improve Human-Robot Cooperaiion, \u201d Proc. of 12th IEEE Ini\u2019l Workshop on RO-MAN, 2003 Robert Boakes, Psychology and ihe Minds of Animals, Cambridge lJniver.sity Press. 1984. Hiroyuki Ishii, Tomohide Aoki, Masaki Nakawji, Hiroyasu Miwa, Aisuo Takanishi, \u201cExperimenial Study on Inieraciuion behveen a Rai and a Rai-robot Based on Animal Psycholog,\u201d, Proc. o f ihe 2004 IEEE International Conference on Roboiics and Automation, 2004 [ 5 ] B.F. Skinner, The Behavior of Organism : An Experimenial Analysis, Appleton-Century. 1938 [2] [3] [4] 5 r 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0" ] }, { "image_filename": "designv11_61_0003999_icma.2012.6285092-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003999_icma.2012.6285092-Figure3-1.png", "caption": "Fig. 3 Joint angle and spring stiffness", "texts": [], "surrounding_texts": [ "with its corresponding eigenvectors. Effectiveness of our method is demonstrated through numerical examples. Our method admits coexistence of frictionless sliding contacts and frictional rolling contacts in a grasp. The analysis of frictionless sliding contact is applicable for the case that the contact friction condition is unknown before grasping. In this case, optimum grasp location is determined from the grasp stability of frictionless sliding contact model to make stable grasp.\nII. PROBLEM DEFINITION\nWe consider that multiple objects are grasped by multifingered hands with revolute joints (Fig. 1). The analysis of this paper is based on Refs. [15] and [16]. A. Assumptions In order to clarify and simplify our problem, we analyze it under the following assumptions: (A1) The objects and the fingers are rigid bodies (A2) There is a single point contact between two bodies. (A3) Initial grasp configuration is given and in wrench (force and moment) equilibrium. (A4) The local curvature at each contact point is given. (A5) An infinitesimal configuration (position and orientation) displacement of the object occurs due to an external disturbance. (A6) Each finger is constructed with two revolute joints. The relation between the joint torque and the joint angle displacement is replaced with a virtual rotational stiffness model. In Assumption (A6), each joint is controlled by compliance control. B. Symbols\nWe use the following coordinate frames and symbols (Fig. 2):\nb : Base coordinate frame of the hand.\nLfk : Local coordinate frame fixed at the initial contact\nlocation on the k-th finger surface. Cfk : Contact coordinate frame on the k-th finger.\noj : Object coordinate frame fixed in the j-th object.\nLojk : Contact coordinate frame fixed at the initial contact\nlocation on the j-th object surface contacting with the k-th finger.\nCojk : Contact coordinate frame on the j-th object.\nA homogeneous transformation matrix of b with respect to a is defined by the following form:\n33\n21 10\nb a b a b a RT p , (1)\nwhere the vector 2 b a p and the matrix 22 b aR are the position and orientation components, respectively.\nThe vector 3],[: oj T ojoj x denotes an j-th object\nconfiguration (position and orientation) displacement, where 2 ojx and oj are the translation and rotation\ncomponents, respectively. )()()( ojrotojtransojoj boj TTT x , 3)0( IToj boj .\nwhere\n10 )(\n21\n2 x x I Ttrans ,\n10 0)Rot( )( 21 12 rotT ,\ncossin sincos )Rot( .\nC. Joint angle and joint torque Each finger is constructed with two revolute joints, and the angle of these joints is represented by 2\n2,1 ][ T kkk qqq . (2)\nThe joint angle is decomposed into the following three elements:\ndklcklnklkl qqqq , (3) where nklq is a natural angle of the virtual joint spring, cklq is an initial compression to generate the initial grasping force, and dklq is a compression due to the object configuration displacement. Stiffness of the virtual spring is denoted by kls and the initial joint torque is given by\n2 cklklkl qs , ],diag[ 21 kkk ssS . (4)\nIII. STIFFNESS MATRIX OF A SINGLE OBJECT\nWhen a grasped object is displaced due to an external disturbance, each finger passively moves on the object surface. The parameter 2\n21 ],[ dkdkdk qqq depends not only on an object configuration displacement and the contact friction condition (frictionless sliding contact and frictional rolling contact). In this section, we derive this relation. A. Contact Constraint between an Object and Each Finger\nWhen the hand grasps the object, each fingertip maintains contact with the object. We have the following contact constraint:\n),()(\n)()(\nojkCojk Lojk ojLojk b\nCojk Cfk fkCfk Lfk dkLfk b\nTT\nTTT q (5)\nwhere the matrices )( dkLfk bT q and )( ojLojk bT are local", "coordinate frames on the finger surface and the object surface at the initial contact point with respect to the base frame, respectively. Lojk oj ojoj boj boj b ojLojk b TTTT )()( ,\nLfk k dkk k dkk k k b dkLfk b TqTqTTT 2 22 1 11 0 0 )()()( q ,\n)()()( 1,1, dklrotcklnklbkl lk dklkl lk qTqqTqT . (6)\nThe matrices )( fkCfk LfkT and )( ojkCojk LojkT are current\ncontact coordinate frames with respect to the local coordinate frames, respectively. The parameters fk and ojk are arc\nlengths of the contact location displacement on the finger and object, respectively.\n,)(\n10 }){Rot()Rot()(\n1 21\n1 1 2\nfk Lfk fkfkrotfk Lfk\nfkfkfkfkfk fkCfk Lfk\nTTT\nIT\nu\n1)()( ojk Lojk ojkojkrotojk Lojk ojkCok Lok TTTT ,\n)( 1 1u fktransfk Lfk TT , )( 1 1 u ojktransojk Lojk TT ,\n3)0( ITCfk Lfk , 3)0( ITCojk Lojk , T]0,1[1 u . (7)\nThe constants fk and ojk are the local curvatures at contact\npoint. If the body surface is convex, concave, flat, then the curvature becomes positive, negative, zero, respectively. From )()(lim 2\n0 ufktransfkCfk Lfk TT fi , T]1,0[2 u , (8)\nthe definition (7) includes the shape of flat surface. From (A2), 0 ojkfk .\nB. Joint Angle Displacement and Potential Energy From the rotation component of (5), we have the following constraint:\n. 0 )}()({\n)()(]0,[\n2 21\n21\n122\nI I TT\nTTTI\nojkCojk Lojk Lojk b\nCojk Cfk fkCfk Lfk dkLfk b\nq (9)\nHence, we have ojkojkojfkfkdkdk qq 21 . (10)\nThe parameter fk is reduced by substituting (10) into (5).\nHence, the translational component of (5) is represented as\n. 1\n0 )()(]0,[\n1 0 )(]0,[\n12 122\n12 122\nfk Cojk ojkCojk Lojk ojLojk b\nfk Lfk dkLfk b\nTTTI\nTTI q (11)\nAlthough the constraint is nonlinear, it can be formally expressed as the following form:\n),( ojkojdkdk qq . (12)\nThe potential energy of each finger is given by\n)}.,({)},({\n),(\n2 1\nojkojdkckk T ojkojdkck\nojkojk\nS\nU\nqqqq (13)\nThe first and second order partial derivatives are given by\nfk b\nb fk ojk fk ojk\nfs k\nk\nk\nR W U U\nf\np ][,\n, ,\n, 0\n00\n][\n][ ][\n2 33\n,,\n,,\nojk T ojk\nojk fk b b fkT ojk fk\nT\nfk b b ojT fk oj\nT T fk b\nb fk ojk fk ojk\nfs k\nfkk T fk b\nb fk ojk fk ojk\nfs k\nkk\nkk\nR\nR\nJ R\nW\nSSJ R\nW\nUU UU\nv v fp\nvv fp\np\np\n(14)\nwhere ,kU is the first order partial derivative of (13) by oj .\n,iU is the second order partial derivative by oj and ojk .\nThe other symbols are defined as follows:\nb oj\nfk oj\nfs jk R\nI W\np 2 ,\n1 0 12 v ,\nk T fk b fk b J f ,\n2\n1\nk\nk k ,\nT\nb k fk k\nb k fk k\nT dk\ndkfk b\nfk b\nR R J 22\n11\n0| )(\np\np\nq\nqp ,\nCfk\nojk\noj Lojk\nCojk\nLfk\nfk ojk\nfk Object\nFinger\nFig. 4 Contact coordinate frames", ". 11 10 00 01\n| )}({\n2 2 1 1\n0\n2\nkfk k k bT fk b kfk k k bT fk b\nT dkdk\ndkfk bT fk b\nfk\nRR\nS\npfpf\nqq\nqpf\nAs shown in (14), the joint angle derivatives fk b J and\nfkS exist. These terms mean the effect of revolute joints\ndifferent from Ref. [15] C. Frictionless Sliding Contact\nIn the case of frictionless contact, each finger slides on the object surface and its potential energy will be local minimum (Fig. 6). Hence, we have\n0 ),(\nojk\nojkojkU\n, 0\n),(2\nojkojk\nojkojkU\n. (15)\nFrom (15) in the initial grasp state, we have 0, kU , 0, kU . (16)\nFrom the first condition of (15), the parameter ojk is given\nby a function of the object configuration displacement oj .\nThe potential energy with sliding contact is represented by ))(,()( oj fs ojkojkoj fs k UU . (17)\nFrom the first condition of (15), we have\n1 ,,0|\n)(\nkk\noj\noj fs\nojkfs k UUQ\n. (18)\nTherefore, the first and second order partial derivatives are calculated as\n,}{\n],[| )(\n33\n, ,\n, 30\nfk bfs k\nk k\nkfs k\noj\noj fs\nkfs k\nW\nU U U QI U G\nf\n(19)\n.\n| )( ],[],[\n| )(\n33 , 1 ,,,\n0\n2\n,3 ,,\n,, 3\n0\n2\nkkkk\nT ojoj\noj fs\nojk k Tfs k\nkk\nkkfs k\nT ojoj\noj fs\nkfs k\nUUUU\nUQI UU UU QI\nU H\n(20)\nD. Frictional Rolling Contact In the case of frictional contact, each finger rolls on the object surface (Fig. 7) and we have the following rolling contact constraint:\n0 fkojk . (21)\nFrom (10) and (21), the parameters ojk and fk are reduced\nand the potential energy of each finger is given by ))(,()( oj fr ojkojkoj fr k UU . (22)\nHence, the first and second order partial derivatives of (22) are given by\n}{| )( 0 Lfk bfr\nk oj\noj fr\nkfr k W\nU G f\n, (23)\n,}{\n1 1 }{ 1 1 }{\n}}{}{}}{{{\n| )(\n1\n0\n2\nT Lfk oj oj bT Lfk b\nT T\nLfk bfr k T Lfk bfr k\nfkojk\nLfk bT Lfk b\nTT Lfk bfr kLfkk T Lfk bfr k\nT ojoj\noj fr\nkfr k\nR\nJWJW\nR\nJWSSJW\nU H\nvvpf\nvv\nuf\n(24)\nwhere\nb oj\nLfk oj\nfr k R\nI W\np\n2 , i T Lfk b Lfk b J f ,\nT\nb k Lfk k\nb k Lfk k\nT dk\ndkLfk b\nLfk b\nR R J 22\n11\n0| )(\np p q qp ,\n. 11 10 00 01\n| )}({\n2 2 1 1\n0\n2\nLfk k k bT Lfk b Lfk k k bT Lfk b\nT dkdk\ndkLfk bT Lfk b\nLfk\nRR\nS\npfpf\nqq\nqpf\nIV. STIFFNESS MATRIX OF MULTIPLE OBJECTS\nWhen the configuration of i-th object frame is displaced due to an external disturbance, the configuration of j-th object frame is also displaced and the j-th object slides or rolls on the i-th object depending on the contact condition. In this section, we derive the relation between oi and oj .\nA. Frictionless sliding contact between two objects From Fig. 1, contact coordinate frame Coij is constrained\nby\n).()(\n)()(\noijCoij Loij Loij oi oioi boi boi b\nCoij Coji ojiCoji Loji Loji oj ojoj boj boj b\nTTTT\nTTTTT\n(25)\n),( ojkojkU\nojk\n),0( ojkkU\n0\nFig. 6 The potential energy condition for frictionless sliding contact" ] }, { "image_filename": "designv11_61_0001259_jmes_jour_1972_014_043_02-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001259_jmes_jour_1972_014_043_02-Figure2-1.png", "caption": "Fig. 2. A n example oj-a gear train and its schematic diagram", "texts": [ " For this purpose, it is convenient to make use of a schematic block diagram, of a type similar to that used by Macmillan (8) (11), in which a p.g.t. is represented by a box, as shown in Fig. la. Connections are represented schematically by circles, as shown in Fig. Ib, which shows a connection joining three internal shafts (numbered 1,2 and 3); the shaft numbered 4 is the external shaft associated with the connection. The schematic diagrams used by Molian (4) differ in appearance from those used here, but they are topologically equivalent. The example shown in Fig. 2 demonstrates the construction of the schematic diagram for a complete train; Fig. 2a is a diagram of a three-speed Wilson gearbox, and Fig. 2b is the corresponding schematic diagram. For the purposes of illustration the sun, annulus and planet carrier for each p.g.t. are labelled S, A and C,respectively, Journal Mechanical Engineering Science Vol 14 No 5 1972 at UNIV CALIFORNIA SAN DIEGO on April 20, 2016jms.sagepub.comDownloaded from with appropriate subscripts, and the external shafts are lettered a to e: a is the input shaft, b is the output shaft, and c, d and e the three control members. The determination of the schematic diagram for known numbers of connections of each type corresponds directly to the problem in graph theory of the enumeration of the graphs for known numbers of vertices of given degrees", " This latter problem is one for which a general solution is at present not known, although methods of solution which, are applicable in particular cases have been described by Woo (12)~ Freudenstein (rj) , and Buchsbaum and Freudenstein (3). The solutions for the present problem for values of rn up to 4 are illustrated in Fig. 3, where, for example, the group designation 3 1 1 1 (with m = 4) indicates that three of the shafts associated with the elemental p.g.ts remain unconnected, and there is one connection joining two internal shafts, one joining three and one joining fbur shafts. The three-speed Wilson gear train shown in Fig. 2 is seen to be of the type 2 2 1. In most cases, there is only one schematic diagram for each designation, the largest number of different arrangements for any of the cases shown is four for the 2 2 2 0 system with m = 4. Up to this point, no distinction has been made between the shafts in the schematic representation of the elemental p.g.t. This means that each \u2018box\u2019 in the diagrams can be interpreted as six different p.g.ts of the sun, annulus and planet carrier type. Whilst this may appear to introduce additional complications into the schematic diagrams, the problems during synthesis are more apparent than real since it should be noted that when a numerical value has been assigned to the basic speed ratio of an elemental p" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001142_cdc.2005.1582306-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001142_cdc.2005.1582306-Figure1-1.png", "caption": "Fig. 1. The TORA system.", "texts": [ " Nijmeijer are with Department of Mechanical Engineering, Eindhoven University of Technology, P.O. Box 513, 5600 MB, Eindhoven, The Netherlands. Alexey.Pavlov@itk.ntnu.no, B.R.A.Janssen@student.tue.nl, N.v.d.Wouw@tue.nl, H.Nijmeijer@tue.nl In Section V we present and discuss experimental results. Section VI contains conclusions. The results presented in this paper are part of the work [13]. Consider the so-called TORA-system (Transitional Oscillator with a Rotational Actuator), which is shown in Fig. 1. This system consists of a cart of mass M which is attached to a wall with a spring of stiffness k. The cart is excited by a disturbance force Fd. In the center of the cart, there is a rotating arm of mass m. The center of mass of the arm CM is located at a distance of l from the rotational axis and the arm has an inertia J with respect to this axis. The arm is actuated by a control torque Tu. The cart and the arm move in the horizontal plane and, therefore, gravity effects are omitted. The horizontal displacement of the cart is denoted by e and the angular displacement of the arm is denoted by \u03b8" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002258_b978-1-4831-9821-7.50021-1-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002258_b978-1-4831-9821-7.50021-1-Figure2-1.png", "caption": "FIG. 2 . Collision relationships.", "texts": [ " Collision Equations The computational procedures described above can be applied to every node station at every instant of time except when the resulting location of a node station lies beyond the rigid wall. In this case a collision without friction is assumed and the calculated re bound location is introduced as the corrected node location. During the calculation process, the resulting displacement increments \u0394 \u039a and AW are always treated first as trial values, which are denoted by superscript i. The correct values to be used in the subse quent analysis are determined according to the following relationships (see Fig. 2): (MV\\: when D-R < 0 AW,j=\\ (21) (AV\\j when D-R<0 \u0394\u039a\u00b7.,\u00b7=1 (22) In these equations Dij is the calculated radial location of the ith node and is given by R is the radius along the mid-line of the ring and e is the coefficient of restitution which may assume a value between 0 and 1. Dynamic buckling of a circular ring 2 8 9 The deflection of the point lying on the vertical axis of symmetry, nondimensionalized with respect to the ring radius, is plotted along the abscissa and a nondimensional load parameter, \u03bb = PR^/EI along the ordinate axis" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002513_detc2008-49084-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002513_detc2008-49084-Figure3-1.png", "caption": "Fig. 3 Pseudo-Cartesian coordinates, (\u03be, \u03b6), of a point, P, of the positive shell.", "texts": [ "asmedigitalcollection.asme.org/pdfaccess.ashx?ur So = sequence which the secondary IPs can be determined with; Sp = set of the primary IPs of an SM; Ss = set of the secondary IPs of an SM; \u03b5 = slope angle (this angle identifies a pencil of meridians, see Fig. 2); \u03b8 = signed dihedral angle that locates one great circle in a pencil of meridians with assigned slope angle (see Fig. 2 and \u00a7\u00a7 2.1); (\u03be, \u03b6) = pseudo-Cartesian coordinates (these coordinates locate a point on the positive hemisphere of the RS, see Fig. 3); In spherical mechanisms (SMs), the instantaneous motion between two links is a rotation around an instantaneous rotation axis (instantaneous pole axis). The instantaneous (first-order) kinematics of SMs can be fully described by analyzing the positions of these axes [1]. Instantaneous pole axes (IPAs) have properties that are the spherical counterparts of instant centers' ones. Instant centers properties are fully exploited to analyze the kinematic behavior of planar mechanisms (see, for instance, [2-4])", " Since the IPAs' positions are sufficient to fully describe the first-order kinematics of the SMs, the first-order kinematics of the SMs can be studied by using only one (either positive or negative ) shell of the RS. Hereafter the positive shell will be used. The intersection of an IPA and the positive shell is a point named instantaneous pole (IP) [1]. In the positive shell, the slope angle, \u03b5, belongs to the range ]\u2212\u03c0/2, \u03c0/2] (rad), and it is sufficient to locate the IGC points of the shell. The remaining points of the shell are the points of the positive hemisphere, and they need two coordinates to be located. With reference to Fig. 3, each of these points can be seen as the intersection of two great circles, one with zero slope angle and the other with \u03c0/2 (rad) slope angle, which form the dihedral angles \u03b6 and \u03be with the PGC and the DGC respectively. \u03b6 is taken positive if clockwise with respect to the y axis, whereas \u03be is taken positive if counterclockwise with respect to the z axis. Both \u03b6 and \u03be belong to the range ]\u2212\u03c0/2, \u03c0/2[ (rad), and will be called pseudoCartesian coordinates2 of the point. The pseudo-Cartesian coordinates will be used to locate the point of the positive hemisphere" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003015_cae.20257-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003015_cae.20257-Figure5-1.png", "caption": "Figure 5 Diagram for retrieving the data points of measured object.", "texts": [ " It can be used in the browsing of CMM scanning to set up the sequence of the actual scanning procedure (including initializations, inspecting region, step angle, etc.) through the interactive nodes. (3) Decide dynamic interference: First of all, retrieve the point data on the measured object within the visible field of the CCD lens. Since the surface of the object is of the NURBS-defined convex, the normal vector of any control point (T) on the surface together with two point data (for example, P and Q in Fig. 5) on the lens can establish a CCD plan. Moreover, the coordination of the points on the object surface corresponding to the points on the CCD plan is decided and the cosine values of PQ and PT vectors can be calculated as indicated in Equation (1). Finally, as in Equation (2), from the sine value of PT, the distance between the two points can be obtained. y \u00bc cos 1 \u00f0PTjPQ\u00de \u00f0jPTjPQj\u00de \u00f01\u00de d \u00bc jPT jsin y \u00f02\u00de The node of collision design, the \u2018\u2018Convex collision manager\u2019\u2019 node is used to provide information about the position and angle of the object and to transmit interference information through the output events" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003035_iccas.2008.4694544-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003035_iccas.2008.4694544-Figure2-1.png", "caption": "Fig. 2 Geometry relationship between CG, CB and CV", "texts": [ " Finally, writing the equations of motions of airship in state space form, ( ) ( ) ( ) T \u03b7\u23a1 \u23a4+ + =\u23a3 \u23a6M\u03bd C \u03bd \u03bd G \u03b7 R \u03b8 g U (4) The inertia matrix M is sum of the inertia of the rigid body ( rgM ) and the buoyancy air ( amM ). Diagonal terms aM and aJ matrices designate the added mass and added inertia, respectively. ( ) ( ) 6 6 cv a cg rg am cv cg a m m \u00d7 \u23a1 \u23a4\u2212 \u00d7 \u23a2 \u23a5= + = \u2208 \u23a2 \u23a5\u00d7\u23a3 \u23a6 M r M M M r J (5) where, 0 0 0 0 0 0 u a v w m X m Y m Z \u2212\u23a1 \u23a4 \u23a2 \u23a5= \u2212\u23a2 \u23a5 \u23a2 \u23a5\u2212\u23a3 \u23a6 M , 0 0 0 0 x xz a y xz z J J J J J \u2212\u23a1 \u23a4 \u23a2 \u23a5= \u23a2 \u23a5 \u23a2 \u23a5\u2212\u23a3 \u23a6 J , ( ) 0 0 0 0 0 z cv cg z x x a a a a \u2212\u23a1 \u23a4 \u23a2 \u23a5\u00d7 = \u2212\u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 r (6) m is the mass of the airship and the cross product matrix ( )cv cg \u00d7r has distance variables ( xa , za ) form CV to CG in Fig. 2. Coriolis matrix ( )C \u03bd is given in Eq. (6), which contains the nonlinear forces and moments due to centrifugal and Coriolis forces. The matrix ( )G \u03b7 contains the restoring forces and moments formed by the buoyancy and gravity. ( ) ( )( ) ( )( ) ( ) 6 6( ) cv a cg cv cg a m m \u00d7 \u23a1 \u23a4\u00d7 \u2212 \u00d7 \u00d7 \u23a2 \u23a5= \u2208 \u23a2 \u23a5\u00d7 \u00d7 \u00d7\u23a3 \u23a6 \u03c9 M \u03c9 r C \u03bd r \u03c9 \u03c9 J , ( ) ( ) ( ){ } 3 3 6 3( ) B cv cv cg B cb m m m m \u00d7 \u00d7 \u23a1 \u2212 \u2212 \u23a4 \u23a2 \u23a5= \u2208 \u2212 \u00d7 + \u00d7\u23a2 \u23a5\u23a3 \u23a6 I G \u03b7 r r (7) where, ( ) 0 0 0 0 0 z cv cb z x x b b b b \u2212\u23a1 \u23a4 \u23a2 \u23a5\u00d7 = \u2212\u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 r , ( ) 0 0 0 r q r p q p \u2212\u23a1 \u23a4 \u23a2 \u23a5\u00d7 = \u2212\u23a2 \u23a5 \u23a2 \u23a5\u2212\u23a3 \u23a6 \u03c9 , 0 0 g \u23a1 \u23a4 \u23a2 \u23a5= \u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 g (8) The cross product matrix ( )cv cb \u00d7r has distance variables ( xb , zb ) form CV to CB in Fig. 2. Finally, terms of the right hand side of the Eq. (4) are external forces uF and moments uT and defined in Eq. (9) : i.e. aerodynamic, gravitational, buoyant, and propulsive forces and moments, respectively. ( ) ( ) , , , , , , , , , , u T e r u T e r \u03bc \u03bc \u03b4 \u03b4 \u03b4 \u03b4 \u03b1 \u03b2 \u03b4 \u03b4 \u03b4 \u03b4 \u03b1 \u03b2 \u23a1 \u23a4 \u23a2 \u23a5= \u23a2 \u23a5\u23a3 \u23a6 F U T (9) In summary, kinematics of Eq. (1) and dynamics of Eq. (4) is utilized for the design of the tracking controller in next section. The state vector 6TT T\u23a1 \u23a4= \u2208\u23a3 \u23a6\u03bd v \u03c9 is composed of six state variables and the real control inputs are composed of thrust T\u03b4 , tilting angle \u03bc\u03b4 , elevator e\u03b4 , and rudder r\u03b4 " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002602_s00707-007-0564-3-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002602_s00707-007-0564-3-Figure4-1.png", "caption": "Fig. 4. A model of three DOFS with symmetry type C3V", "texts": [ " The mass and stiffness matrices are M \u00bc m1 m1 m2 m2 m1 m1 m2 m2 2 66666666664 3 77777777775 8 8 ; K\u00bc k1\u00fek2\u00fek4\u00fek6 k4 k2 0 k6 0 0 0 k4 k1\u00fe k2\u00fe k4 \u00fek6 0 k2 0 k6 0 0 k2 0 k2\u00fek3\u00fek5 k5 0 0 k3 0 0 k2 k5 k2\u00fek3\u00fek5 0 0 0 k3 k6 0 0 0 k1\u00fek2\u00fek4\u00fek6 k4 k2 0 0 k6 0 0 k4 k1\u00fek2\u00fe k4\u00fe k6 0 k2 0 0 k3 0 k2 0 k2\u00fek3\u00fek5 k5 0 0 0 k3 0 k2 k5 k2\u00fek3\u00fek5 2 66666666666664 3 77777777777775 8 8 : Using the symmetric modes, the matrix T is formed as T \u00bc 1 2 1 0 1 0 1 0 1 0 1 0 1 0 1 0 1 0 0 1 0 1 0 1 0 1 0 1 0 1 0 1 0 1 1 0 1 0 1 0 1 0 1 0 1 0 1 0 1 0 0 1 0 1 0 1 0 1 0 1 0 1 0 1 0 1 2 6666666666666664 3 7777777777777775 : For k1 = 1,000, k2 = 750, k3 = 500, k4 = 1,500, k5 = 1,750, k6 = 2,000, m1 = 50, m2 = 100 and P\u00f0t\u00de \u00bc f50; 0; 0; 0; 0; 0; 0; 0g sin ffiffiffi 5 p t; using T, the displacements are calculated as w \u00bc 0:0333 0:1000 0:0030 0:0005 0:0057 0:0011 0:0050 0:0030 2 66666666664 3 77777777775 sin ffiffiffi 5 p t) u \u00bc T w) u \u00bc 0:0098 0:0186 0:0477 0:0493 0:0178 0:0205 0:0511 0:0518 2 66666666664 3 77777777775 sin ffiffiffi 5 p t; which is the result of the classical modal method. Table 5. Representation table of symmetry group C2V in eight-dimensional space Element Gi Gi (u) v(Gi) e G1 {u1, u2, u3, u4, u5, u6, u7, u8}t 8 C2 G2 {u6, u5, u8, u7, u2, u1, u4, u3}t 0 rV G3 {u5, u6, u7, u8, u1, u2, u3, u4}t 0 r0V G4 {u2, u1, u4, u3, u6, u5, u8, u7}t 0 The symmetry group of C3V consists of six symmetry operations, which are obvious in the structure of three DOFs, shown in Fig. 4. The symmetry operations are: {e, C3, C @1 3 , ra, rb, rc}. C3V is one of the symmetry groups of higher order, since the first representation of dimension 2 appears in this group, resulting in some trivial subspaces. The character table is given in Table 6. Due to the existence of the irrep of dimension 2, the corresponding subspace has two duplicate roots, and finding one of these is enough to form T. Dimensions of these subspaces are n1 = 1, n2 = 0, n3 = 2 1. Thus the whole problem of three DOFs can be changed into two single-DOF problems" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003852_amr.655-657.578-Figure10-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003852_amr.655-657.578-Figure10-1.png", "caption": "Fig. 10 Bending Stress nephogram of Meshing Teeth Pairs ( 1z =99, 2z =100)", "texts": [ " Due to great contact stress, the contact stress of external teeth pairs of the two pairs of teeth is very small, indicating only slight contact occurs. 3. At meshing-in end and meshing-out end, the tooth top and tooth root of gear meshes, the internal gear of meshing-in end and external gear of meshing-out end will also bear great bending stress (usually not the maximum value). This phenomenon can be clearly understood through observing bending stress could chart and bending deformation graph, as shown in Fig.10 and Fig.11 respectively. Make a comprehensive analysis by combining three kinds of stress broken line graphs, contact stress vector diagram, bending stress cloud chart, bending stress deformation graph, the number of apparently deformed meshing teeth pairs can be determined. For example, observing FIg. 3 (b) , FIg. 9,FIg. 10 and FIg.11, the number of apparently deformed meshing teeth pairs can be determined as 6 pairs when 1z =99, 2z =100. This article names the apparently deformed teeth pairs undertaking main load as effective meshing teeth pairs. According to calculation, when the tooth number of internal gear 2z =40-200 and tooth number difference 12 zz \u2212 =1, the number of effective meshing teeth pairs is 3-15 pairs, as shown in Table 2. when the tooth number difference 12 zz \u2212 in increased to 5, the number of effective meshing teeth pairs will reduce accordingly and changed into 3-11 pairs, as shown in Table 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000979_memsys.1995.472580-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000979_memsys.1995.472580-Figure1-1.png", "caption": "Fig. 1 Stator fabrication process", "texts": [ " The examination reveals that the new process is applicable to the manufacture of electromagnetic micro-devices. 2. STATOR FABRICATION PROCESS The crucial points in making a breakthrough in the development of a highly efficient generator are the formation of a coil with a high density winding, a core die with a high aspect ratio, and a magnetic body in the core die. To fabricate a coil with a high density winding, the insulation layer must be thin, and to make a core economically with a high aspect ratio, i.e. with a great thickness, it is necessary to reduce the time required for processing. Figure 1 shows the process of forming the stator. Currently, planar processing using semiconductor equipment is indispensable for forming thin insulation layers. For this reason, in the new process of fabricating the cylindrical stator, the coil and the core die were made on separate substrates, and after the coil was removed from its substrate, it was inserted and installed in the core die. The remaining area was filled with a magnetic material by electroplating. 2.1 Fabrication of the coil Figure 2 shows the fabrication process of the coil" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003061_vppc.2008.4677496-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003061_vppc.2008.4677496-Figure2-1.png", "caption": "Fig. 2 Eddy current produced by the tangential components of flux density", "texts": [ " The equivalent conductivity in lamination regions is anisotropic, and can be expressed by a conductivity tensor \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 = min ][ \u03c3 \u03c3 \u03c3 \u03c3 00 0k0 00k lam lam a (6) where klam is the lamination factor, and \u03c3min, the minimum conductivity limit, is used to ensure that the system equation is non-singular. Hence, the field equation (4) becomes )]([)]([ 1 \u03a9\u2207+ \u2202 \u2202+\u00d7\u2207\u00d7\u2207 \u2212 TT \u03bc\u03c3 ta )]([ pcst HHH ++ \u2202 \u2202\u2212= \u03bc . (7) B. Effects of eddy current loss caused by the tangential components of flux density When the flux density components Bx and/or By, tangential to the lamination plane, alternate, the produced eddy current field is bounded in each lamination, as shown in Fig. 2. The effects of eddy current produced by the tangential flux components can be considered by means of an equivalent magnetic field component Hpc, which can be computed by eddy current core loss. The eddy current core loss per unit volume is normally given in the frequency domain as shown below 2 mcc fBkp )(= (8) where f is the frequency, Bm is the flux density amplitude, and kc is the eddy current loss coefficient which can be calculated from 6dk 22 c /\u22c5\u22c5= \u03c3\u03c0 . (9) In the time domain, the eddy current core loss is computed from tt ktp c c \u2202 \u2202\u22c5 \u2202 \u2202= BB 22 )( \u03c0 " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000117_135065003322620246-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000117_135065003322620246-Figure5-1.png", "caption": "Fig. 5 Examples of rigs used to simulate loads applied to connecting rod bearings during service: (a) rig used by Yamaguchi et al. [48]; (b) schematic electrical circuit used on the rig shown in (a) to measure minimum oil \u00aelm thickness by an electrical resistance method; (c) rig used by Masuda et al. [34] in which oil \u00aelm pressure signals from sensors mounted in the test journal are extracted via slip rings", "texts": [ " 2c), and indicates that the `added mass\u2019 of the linkage has the effect of increasing the inertia of the engine mechanism reciprocating components by 10 per cent. It is claimed that the minimum oil \u00aelm thickness is therefore reduced by only about 0.2 mm and there is a negligible effect on its position. The effect on the oil \u00aelm thickness at smaller eccentricities is shown to be much greater, however. Further research of connecting rod big-end bearing performance was reported by Yamaguchi et al. [48] following studies using a simulation test rig as shown in Fig. 5a. Figure 5 also shows simulation rigs used by other researchers. In these tests, the connecting rods were pinned at the small-end bearing and a test journal carrying the big-end bearing ran also in slave bearings, the test journal being driven at a range of speeds. Sinusoidal loading was applied to the big-end bearing from an actuator via a lever bar connected to the smallend bearing. The oil \u00aelm resistance technique (see Fig. 5b) was used to measure minimum oil \u00aelm thickness and to indicate bearing seizure. Lubricant supply temperature was maintained constant, in order to prevent any variation in resistance with temperature from affecting the results. However, the focus of the work was on detecting surface-to-surface contact rather than on measuring the minimum oil \u00aelm thickness. A simulation rig had also been used earlier by Sasaki et al. [49] in which mass loss and wear rate were measured at staged intervals in a long duration test", " Holmes concluded that pressure variation was in agreement with theoretical predictions, and that a temperature rise of about 15 K above the supply temperature occurred. J02701 # IMechE 2003 Proc. Instn Mech. Engrs Vol. 217 Part J: J. Engineering Tribology at Kungl Tekniska Hogskolan / Royal Institute of Technology on June 30, 2015pij.sagepub.comDownloaded from Experimental rigs were used to simulate the load conditions in a big-end bearing by both Hashizume and Kumada [51] and Masuda et al. [37] (see Fig. 5c) in independent studies. In both cases the rig incorporated a dummy crankshaft journal mounted in slave bearings, and an alternating load was applied via the con-rod. Reference [51] was concerned with assessment of oil \u00afow and its relationship to bearing temperature measured by transducers located against the bearing shell. Reference [37] was concerned with oil pressure distribution measured by means of transducers located in the journal, the signal being extracted by means of slip rings. Hashizume and Kumada [51] concluded that bearing temperature was inversely proportional to oil \u00afowrate, and that increased con-rod rigidity results in a decreased oil \u00afow", " Figure 3a is reprinted from Trans. ASME, J. Engng Power, April, 1967 by Westbrook and Munro, `The telemetering of information from a working internal combustion engine\u2019, pp. 247\u00b1254, with permission from ASME. Figure 3b is reprinted with permission from SAE 780980 # 1978 Society of Automotive Engineers. Figure 3c is reprinted with permission from SAE 960989 # 1996 Society of Automotive Engineers. Figures 5a and 5b are reprinted with permission from SAE 892114 # 1989 Society of Automotive Engineers. Figure 5c is reprinted from STLE Tribology Transactions, Vol. 35, 1, by Masuda et al. `A measurement of oil \u00aelm pressure distribution in connecting rod bearing with test rig\u2019, pp. 71\u00b176, 1992, with permission from STLE. 1 Grente, Ch., Ligier, J. L., Toplosky, J. and Bonneau, D. The consequence of performance increases of automotive engines on the modelisation of main and connecting rod bearings. In Proceedings of Leeds\u00b1Lyons Symposium on Tribology, 2000, Elsevier Tribology Series 39. 2 Booker, J. F. Dynamically loaded journal bearings\u00d0 mobility method of solution" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001517_j.physb.2006.04.007-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001517_j.physb.2006.04.007-Figure5-1.png", "caption": "Fig. 5. Major hysteresis loop showing the locations marked a, b, c and d of the minor reptation loops depicted in Fig. 4.", "texts": [ " shows the four characteristic sets of reptation loops starting from convex shape turning into flat wavy bottom to oval shaped and finally into concave form. These are corresponding to near positive saturation position (a), close to zero on the positive side (b), for small negative interruption value (c), and finally near to negative saturation (d). These shapes gradually blend into each other as the position of the first interruption moves between the two magnetization maxima. The positions of the reptation loops used are shown in Fig. 5. The axes of the sets were individually rotated to near optimum, to improve the visual presentation of the loops and to emphasize the differences in their shape. This is the first time that a model gives realistic shapes for these reptation loops confirming the real shapes obtained experimentally [13]. As with the transients shown before the model can predict the final position of the steady state, equilibrium loop from the parameters of the material and details of the process. The drift in the right hand side position of the reversal points of each of the sets is due to the effect of the viscosity" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002449_20080706-5-kr-1001.01308-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002449_20080706-5-kr-1001.01308-Figure1-1.png", "caption": "Fig. 1. Plane pendulum system.", "texts": [ " An energy-like function of the pendulum, called an energy function, is defined, and then energybased controls are obtained that decrease the energy function effectively. They control the actuated variables sinusoidally with constant amplitudes. When the oscillation of the pendulum becomes sufficiently small, each of the energy-based controls is taken over by a linear or saturating control, which also satisfies the constraint. Also, the effectiveness of the control laws is examined by simulations and experiments. Figure 1 shows the pendulum considered in this study, which consists of a rigid rod with a pivot and a weight (a point mass) and moves in the vertical plane; the pivot can be moved in the vertical plane and the weight can be moved along the rod. For simplicity, it is assumed that there is no friction in the rotation of the pendulum. Let \u03b8 be the angular displacement of the pendulum, x and y the coordinates of the pivot, and L the distance from the pivot to the weight. The quantities x, y, and L are variables being able to be actuated; \u03b8 is not an actuated one" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002001_12.709253-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002001_12.709253-Figure4-1.png", "caption": "Figure 4. Needle tip forces (a)-c)) and needle body forces (d)-f))", "texts": [ " These thresholds and constants can be used to define material properties like friction, stiffness, viscosity and penetrability by using them as transfer functions (e.g. k = Ktf (V ( xp))). Figure 3 illustrates the stepwise proxy movement as described in.6 First the proxy is moved (or not moved) in direction of the surface normal to simulate penetrability. Then the proxy is moved perpendicular to the surface normal to generate the friction effect. Finally a repositioning back to the initial proxy position xpO is done to simulate viscosity. The surface resisting force (fig. 4(a)), surface friction (fig. 4(b)), and viscosity (fig. 4(c)), that affect the needle tip, are calculated using the haptic volume rendering approach. We extended the method described in 2.1.2 so that it combines information from label data VL( x) with information given by original CT data VCT ( x). It is possible to ensure haptic feedback from structures that are essential for a realistic haptic sensation by segmenting these structures manually. This sensation is extended by the haptic feedback from structures that are represented in CT density data. We combine the label data VL( x) and CT data VCT ( x) to build up a new image V ( x) additively. We define the values that are used to mark structures like bone, skin, muscles, or fat in the label data VL( x) in such a way that a realistic haptic feedback is generated by the haptic volume rendering approach. The realistic behaviour of the needle is improved by restricting rotation (fig. 4(d)) and transversal motion (fig. 4(e)) of the needle inside the body. Furthermore a friction force (fig. 4(f)), that resists the needles body from sliding through the tissue is computed based on the penetration depth. The needle position and needle direction vectors p and d are stored at the moment the needle pierces the skin to calculate correction forces that restrict the user from rotating the needle inside the body and transversal needle movement. Transversal motion of the needle is restricted using a spring force technique. The position of the proxy xp after the calculation of the needle tip forces (2" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003195_978-3-540-89393-6_21-Figure21.10-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003195_978-3-540-89393-6_21-Figure21.10-1.png", "caption": "Fig. 21.10 Servo by Prox Dynamics AS (0.49 g) (a), currently the lightest classical servo controllable in position. Didel PolyBIRD (0.25 g) (b), controllable only in torque. Toki SmartServo RC-1 (0.8 g) (c), a shape memory alloy-based position-controllable actuator. Figures (a) and (c) reprinted with permission from N. Zimet and Toki Corp., respectively", "texts": [ " They can be used to accurately set the position of a control surface or the pitch of helicopter rotor blades. The torque to weight ratio of RC-servos is in the range of 30\u2013100 gcm/g, and the current draw, 100 mA for the lightest, occurs when they move. The last few years have yielded some very light servo designs based on a lead screw and linear potentiometer. The current record for the lowest weight for a servo is 0.49 g. Developed by Prox Dynamics AS13 for use on miniature rotorcraft, it has a 10 mm range of motion with a resolution of 0.10 mm (Fig. 21.10a). Servos of this size, however, are still hand-built under a microscope and are not available on the market. 13 See: http://www.proxdynamics.com. Wobble motors [7] could be an interesting solution for even lighter servos and have been researched for the last 10 years. Theoretically they can go down to 0.3 g, including required electronics, but they have not yet reached production status. Electromagnetic actuators are probably the most widely used actuators for platforms in the range of 1\u201310 g (see Chaps. 6 and 14). Based on a small magnet rotating inside a coil (Fig. 21.10b), they are often called BIRD (built-in rudder device) in the hobby field, are commonly available from many sources, and can weight anywhere from 2 g down to 0.1 g. Not only are they lighter than servos, they are also simple to control, fast, low cost, and easy to install. The torque of the actuator is proportional to the magnet weight (only the best NdFeB magnets are used) and to the magnetic field. The downside of these actuators is that they are controllable only in torque and not position, and thus their position is dependent on the external forces on the control surface, which can vary greatly during a flight", " At high temperature (typically 80\u25e6 C), NiTi wires take the shape learned during the thermal process, whereas at low temperature they return to their original shape. Frequently named muscle wires, they can shorten by about 4%. Though the force provided is substantial, the response time is relatively poor (0.1+ s). NiTi wires have been used for ultralight rudder actuators down to 0.1 g [15] (see Chap. 19). This custom solution remains delicate, however, since handling of 25\u03bcm wires that cannot be soldered is not an easy task. SMA actuators have some promise, with a commercial product weighing 0.8 g14 (Fig. 21.10c) already available with an announced torque of 15 g/cm for an average current of 30 mA. Electro-active polymers (EAP) [2] are actuators based on polymers that deform when subjected to an electric field. There are two main types of EAPs: ionic EAPs based on the mobility or diffusion of charged ions and electronic EAPs based on electrostatic forces between electrodes that squeeze the polymer. Ionic EAPs react at low voltages and can provide a good force but are generally slow (0.1\u201310 s reaction times) and more power consuming, since they require energy to remain at a given position" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001836_978-3-540-71967-0_2-Figure2.7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001836_978-3-540-71967-0_2-Figure2.7-1.png", "caption": "Fig. 2.7. Concept of an aging aircraft instrumented with active sensors for structural health monitoring", "texts": [ " Structural health monitoring (SHM ), condition-based maintenance (CBM ) and birth-to-retirement refer to the capability of using sensors throughout the life or an adaptronic structure to monitor its state of health and act ac- cordingly. The sensors would record the way the manufacturing process was implemented, and would remember the pristine state of the structure. At the same time, the sensory output will be used to optimize the fabrication process and ensure quality consistency. The network of sensors embedded in the adaptronic structure will be then used to monitor the structural behavior throughout its life (Fig. 2.7). A structural health bulletin will be produced on demand and life history of the structure will be gathered in the database. If needed, active measures will be taken to control and reverse the evolution of structural damage or modify the structures behavior or performance to elude damage. These sensors will monitor the structural aging process and will determine when the artifact should be repaired or even graciously retired. Thus, scheduled maintenance will be replaced by need-based maintenance, with associate savings in the life-cycle costs and increase in the structural safety and equipment availability" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003646_s12283-012-0108-5-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003646_s12283-012-0108-5-Figure1-1.png", "caption": "Fig. 1 The Camber-Ski roller ski, with the camber and adjustable grip function", "texts": [ " The roller skis were equipped with a forward and a rear wheel, one of which had a ratchet to enable grip on the surface (PRO-SKI, ratcheted rear wheel, \u00d867 mm, width 50 mm; Swenor, ratcheted rear wheel, \u00d868 mm, width 45 mm; Marwe, ratcheted forward wheel, \u00d880 mm, width 39 mm). These types of roller skis (with ratcheted forward or rear wheel) had no camber that needed to be pushed down to grip against the surface. Also, six roller skis with a camber and adjustable grip function were tested (Camber-Ski, Mid Sweden University, O\u0308stersund, Sweden), see Fig. 1. The functionality for this was applied to the forward wheel of the roller ski. When sufficient load to press down the camber was exerted (*3 mm), the forward wheel established contact with a ratcheted spool (Rs, \u00d820 mm, stainless steel, cross-knurling pattern size 1.6) situated above the forward wheel. The need to push down the camber was similar to how skiing on snow must be performed to provide grip on the snow in DS and DPKICK. The degree of grip was therefore dependent on the stiffness of the roller ski\u2019s camber, which was simply adjusted via a spring-loaded screw (SF-TFX 2691, Lesjo\u0308fors Stockolms Fja\u0308der AB) This construction used wheels of the same type as the non-ratcheted wheel of PRO-SKI C2" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001946_bf02844205-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001946_bf02844205-Figure2-1.png", "caption": "Figure 2 The equations of motion of the puck (eqn. 2) as well as the trajectories of flight are described in the earth coordinate system (X, Y, Z), with the gravity pointing in negative Z direction. The direction of flight is initially along the earth coordinate system\u2019s X axis. The puck body fixed coordinate system (X\u2032, Y\u2032, Z\u2032) has its origin in the puck\u2019s centre of mass. The Z\u2032 axis of the local coordinate system is the main spinning axis; it is pointing against gravity at the beginning of each shot. The wind coordinate system is defined by (V, N, S). V is the direction of flight, N is in the plane of Z and V, perpendicular to V. S is perpendicular to N and V and is normally pointing sideways.", "texts": [ " To calculate the state for the next time step the derivative of the state is required: state = ( ) = ( ) (2) Here, m and I are the mass and the inertia of the puck respectively, and E is the Euler parameter transformation matrix (Haug, 1989). The gravity force vector is g. The aerodynamic forces (F) are the sum of the drag and lift forces FD and FL, respectively (eqns. 4 and 5). The aerodynamic moment (T ) consists of two parts: the pitch moment (TM, eqn. 6) acting around the S axis of the puck (Fig. 2) and the spin down moment (TN, eqn. 7) generating a moment in the main spinning local z direction, opposite to the spinning v 1\u20442ET 1\u2044mF + g I\u20131(T \u2013 ( \u00d7 I )) r. . v. . d dt r v velocity of the puck. Direction vectors of the forces and moments e\u03bd, en and eS along the centre of mass velocity, normal to the velocity and sideward direction respectively, were defined according to De Mestre (1990) and shown in Fig. 2. The pitch moment as well as lift and drag are dependent on the factor q (eqn. 3), which is the dynamic pressure ( \u03c1\u03bd2). The area of the disc was defined as \u03c0 times the radius of the puck squared, with the radius of the puck as 3.81 cm (Pearsall et al., 1999). The air density is \u03c1 and the free-stream flow speed is \u03bd. Air density is 1.168 kg m\u20133, corresponding to an altitude of 500 m at a temperature of 11.7\u00b0C (De Mestre, 1990). The aerodynamic Robins\u2013Magnus side force and Magnus rolling moment act on the spinning puck, owing to the interaction of near-surface fluid structures" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003245_1.5061338-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003245_1.5061338-Figure5-1.png", "caption": "Figure 5: Principle of the PFO 3D. Two galvanometer drives position the beam in X and Y direction, while the movement of a lens is positioning the beam in Z direction with the same dynamics. All axes move the beam in less than 30 milliseconds from one end to the other.", "texts": [ " Combined, such systems are much more productive than conventional welding systems and therefore perfectly suited for high-volume and high mix welding applications. \u201cRemote Welding\u201d refers to working distances of 300 mm and larger. Large working distances reduce contamination of such scanner optics and prolong the lifetime of the protection glass, with minimal maintenance and low cost per part. The emergence of higher beam quality lasers helps to increase the field size of scanner optics. Depending on the scanner configuration, working fields can reach dimensions of almost one meter today. Figure 5 illustrates the principle for threedimensional scanner optics. Laser Materials Processing ConferenceICALEO\u00ae 2008 Congress Proceeding Page 394 of 909 s \u201cWelding-on-the-fly\u201d TRUMPF\u2019s scanner controller systems can be coupled with the robot motion controller to be fully synchronized with the axes of a robot. This allows extremely fast material processing while the scanner optics is being moved in space by a robot to enlarge the processing space and access the part in 3D. The technology of coupling two systems enables so-called \u201cprocessing-on-thefly\u201d" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003902_s10015-013-0135-8-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003902_s10015-013-0135-8-Figure2-1.png", "caption": "Fig. 2 Cooperation by stress antibody allotment reward", "texts": [ " In this research, a group of robots aims to carry luggage from initial positions to a destination. It is assumed that a robot carrying a luggage experiences stress. The robot obtains reward when it picks up a luggage; however, at the same time, the robot experiences stress. Then it is assumed that antibodies are created, similarly to stress reply mechanism in the living cell as mentioned above and handed over as reward to the robot that comes to cooperate. In such a way, cooperative behavior toward other robot can be promoted. We call the reward as \u2018\u2018stress antibody allotment reward\u2019\u2019. Figure 2 is an illustration of cooperative behavior achieved by stress antibody allotment reward. Robot A is carrying a luggage and meets Robot B. Robot A gives stress antibody allotment reward to Robot B and then Robot A carries the luggage to the destination by cooperation with Robot B with increased speed. 3.2 Switching of learning mode When the environment is complex and plural tasks are targeted, the learning module becomes complex and huge [4\u20136]. In such a case, opportunity for acquisition of actions by recycling the learning results and learning time can be limited" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003095_s0022112070001490-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003095_s0022112070001490-Figure5-1.png", "caption": "FIGURE 5 . Pressure coefficient for supersonic wing with supersonic edges ( M = 0.51, M , = 1.5, k = 0.75, ?J = 0, ~t = 1).", "texts": [ " it works also when the wing tip passes behind the shock. The program can also superpose several basic planforms. Numerical examples for all possible combinations of Mach numbers Ml and B0 and the swept back k and also for several composite planforms are given in Chow & Gunzburger (1969). Figures 5 and 6 show two of the numerical examples for the wings with a straight leading edge. The pressure distribution on the wing at various stations of x are shown together with the various domains in x, y, z space. In figure 5 the Mach numbers of the wing with respect to the stream ahead and behind the shock are both supersonic (MI > 1, Bo > 1). The characteristics of the pressure distribution in various regions are quite obvious. The discontinuities in the slope of the pressure curve as it crosses the boundaries of various domains, e.g. the sonic sphere, the Mach cone, are quite obvious. In particular, along the intersection of the wing with the shock, x = MCt, Diffraction of shock waves by a moving thin wing 603 the disturbance pressure is constant outside the Mach cone of the equivalent wing and is the value along a ray from the vertex T' of the conical solution" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001830_9780470061565.hbb079-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001830_9780470061565.hbb079-Figure5-1.png", "caption": "Figure 5. Schematic flow-through FIA cell setup of the capacitive EIS sensor (a); micromachined flow-through microcells fabricated by combining Si and SU-8 technologies, where the microchannels have been formed as thick SU-8 layers directly onto the already prepared Si\u2013SiO2\u2013Ta2O5 EIS structures: microcell with EIS sensor and Pt ion-generator (b), and microcell with two separate microchannels and EIS sensor for multisensor and/or differential setup applications (c).", "texts": [ " In this way, for example, pHsensitive and penicillinase-modified Ta2O5-gate EIS sensors have been integrated into a flowthrough cell with a variable internal volume from 12 to 48 \u00b5l that is combined with a commercial flow-injection analysis (FIA) system.84,85 The second platform favors a monolithic waferlevel integration of the flow channel together with the EIS sensor structure. In this case, the sensor represents an integral part of the whole flow-through microcell. Here, a micromachined flow-through microcell with integrated EIS sensor was realized by combining Si and SU-8 technologies;85,95 the flow-through micro-channel has been formed in a thick SU-8 layer directly onto an already prepared p-Si-SiO2\u2013Ta2O5 EIS structure (see Figure 5). In order to extend the functional possibilities of this microcell, two thinfilm Pt microelectrodes have been deposited onto the same chip for additional amperometric and flow-velocity measurements. The EIS structures have been integrated in a pH- and penicillinsensitive configuration in flow-through and FIA mode, respectively, yielding a comparable sensor behavior as for the current single pH- and penicillin-sensitive EIS structures.85 In contrast to the capacitive EIS sensor, where the measured value of the analyte concentration is an average value over the whole sensing surface in contact with the analyte, the LAPS measurement has the advantage of being spatially resolved" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002061_j.jsv.2006.11.032-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002061_j.jsv.2006.11.032-Figure2-1.png", "caption": "Fig. 2. 21 MVA generator FE model.", "texts": [ " The effect of EM interaction on other modes (in all 13 modes) included into the analysis was minor. The frequency reduction was up to 3.2% (rotor bending mode 1). The damping ratio was 1.0% for the mode induced by the cage current variables. The remaining modes had negligible damping ratios. The normal frequencies and the electromagnetic damping factors of the undamped EM system as function of rotational speed are depicted in Fig. 4. The results depicted in Fig. 4(a) show that the gyroscopic effect has no effect on the rotor bending modes (Fig. 2). The unbalanced mass frequency-receptance plot is shown in Fig. 3(a) with phase lags in x- and y-directions depicted in Fig. 5. The modal damping ratios of 16% and 4% for the lower and higher rotor bending modes, respectively, were used. For the other modes damping ratio of 1% was applied. The unbalanced mass excitation was located at the rotor shaft center node and the results evaluated on the same node, as well. ARTICLE IN PRESS A. Laiho et al. / Journal of Sound and Vibration 302 (2007) 683\u2013698 691 The receptance is given as length of the semi-major axis of the rotor center elliptical locus", " The second numerical example is a 21MVA synchronous generator. The main parameters of the generator are given in Tables 2 and 3. The parameters b and c were evaluated by means of the products b c by the approach presented in Ref. [23]. The mounting of the generator was modeled imitating the vibration test conditions carried out for the actual machine. The generator was suspended by elastic springs. The natural frequencies of the suspension system and the machine were below 5Hz in the numerical model. The FE model mesh of the generator (see Fig. 2) consists of 8905 elements and 9582 nodes with 6 dofs each. The system order reduction was carried out by 30 lowest modes covering the frequency range 0\u2013100Hz. The EM coupling was implemented by using 13 slices. The 21MVA generator modal data calculated from the undamped (C \u00bc 0 in Eq. (11)) system at rated speed is presented in Table 5. The normal frequencies f and f 0 are for the EM model and mechanical model, ARTICLE IN PRESS A. Laiho et al. / Journal of Sound and Vibration 302 (2007) 683\u2013698694 respectively, with electromagnetic damping ratios x" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002001_12.709253-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002001_12.709253-Figure1-1.png", "caption": "Figure 1. Sensable\u2019s Phantom Premium device with 6 degrees of freedom", "texts": [ " The haptic component renders the forces that affect the needle during the insertion. The forces are calculated in real-time depending on needle position, needle rotation, insertion angle and local tissue properties. The haptic device is then used to return the rendered forces to the user. A force feedback device with six degrees of freedom (Sensable Phantom Premium 1.5) is used for the haptic I/O. This device enables haptic feedback in three directions in space and in three rotation axes of the pen-like end effector (fig. 1). The nominal position resolution of this device is 860 dpi for translation sensing, 0.0023 degrees for yaw and pitch rotation and 0.008 degrees for roll rotation. The translational force output is limited to 8.5 N maximum (peak force) and 1.4 N continuous force output. The maximum exertable torque is 515 mNm for yaw and pitch and 170 mNm for roll. The continuous exertable torques are 188 mNm and 48 mNm respectively. Proc. of SPIE Vol. 6509 65090F-2 Downloaded From: http://proceedings.spiedigitallibrary" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003565_978-3-642-13769-3_39-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003565_978-3-642-13769-3_39-Figure2-1.png", "caption": "Fig. 2. First model. The two IRs, and the press machine with the pair of dies.", "texts": [], "surrounding_texts": [ "5.2 Second Model" ] }, { "image_filename": "designv11_61_0000287_sensor.2003.1215310-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000287_sensor.2003.1215310-Figure1-1.png", "caption": "Figure 1: Schematic view of the magnetic microbeads manipulation principle. Two permanent magnets (a) generate a large and uniform magnetic field Bo in a microfluidic channel (b), enhancing the magnetization of the magnetic beads. In this case, the small time-dependent magnetic field of a set of simple planar coils (c) allows transporting the beads in the channel.", "texts": [], "surrounding_texts": [ "Magnetic beads coated with biologically active layers play an important role in chemicalbiological analysis [1,2]. Therefore, separation [3,4], transport [5 ] , and positioning [6,7] of such beads are important processes in diagnostics, microanalysis and microsynthesis. In case there is no interaction between particles, the force from tbe field (B) on the particle is of the form [3] F = (m x V) B , with m the magnetic moment of the bead ( m = V x , H, withx, the susceptibility and Y the magnetic volume of the particle). Since the effective relative magnetic susceptibility xm of the (super) paramagnetic beads is rather weak (typically xm < 1, due to demagnetization effects of the particle) and the magnetic volume of the particles is small, it is clear that a high absolute value of the magnetic field is beneficial for generating a large field gradient and a large value of the magnetic moment. Generally, the magnetic field of a simple planar coil is too small to induce a large moment of the magnetic microbead, but it is easy to change the size and the sense of such magnetic field. On the other hand, the magnetic induction generated by a permanent magnet is sufficiently large, hut it is static and cannot be varied in time in a simple way. Arrays of planar wires [3,4] have been demonstrated to be a useful solution for magnetic microbeads micromanipulation and separation (in separation, heads are retained magnetically hut are transported by a liquid flow). However, the magnetic field gradient generated by such devices is not large enough to allow magnetic microbeads transport over long ranges. An unresolved problem is to combine the separation step as well as the transport of the separated magnetic microbeads in a single microsystem. In this work, we propose a system for magnetic microhead separation and transport based on the actuation of a simple array of planar coils. The principle of the proposed system is based on the use of a static and uniform magnetic field component, the role of which is to impose a strong magnetic moment to the microheads. The very small magnetic field gradient generated from an array of planar coils is then suficient to cause a displacement of the microbeads over several millimeter distances. Arranging adjacent coils with spatial overlap and actuating them in a specific three-phase scheme assures the long-range displacement of the microbeads. EXPERIMENTAL We have realized simple Printed Circuit Board (PCB) coils (100 pm winding width, 35 pm height, 200 TRANSDUCERS '03 The 12th International Conterence on Saiid State Senwm. Acluatom and MiCmSysLBmS. Basan , June 8-12, 2003 0-7803-7731-1/03/$17.00 02003 IEEE 292 262.3 pm winding pitch) with a low number of windings (N410). They are distributed over two functional PCB layers, so that a strong spatial overlap between adjacent coils is realized. A coil typically generates a magnetic field gradient of about 5 mTimm for a maximum allowed current density of 400 Aimz. To enhance the magnetic moment (and thereby the magnetic force) of the heads, two bar-shaped NdFeB (40 mm x 15 mm x 8 mm) permanent magnets are placed on top of a soft magnetic sheet, generating a uniform field BO = 50 mT at the position of the microfluidic channel, as schematically shown in Figure I . The magnetic microbeads (1 pm diameter, kom Promega [SI) are suspended in water and are manipulated in a simple glass capillary (1 mm outer diameter, 0.5 mm inner diameter). RESULTS AND DISCUSSION The actuation of the microheads by the proposed planar coil system is inspired kom the magnetic field profile generated from a single coil. In the presence of the large uniform transverse magnetic field (B&O mT) generated @om the permanent magnets, magnetic beads will be polarized in the transverse direction, so that only the B, component of the magnetic field generated tlom a coil (of the order of 5 mT) is contributing to the magnetic forces. Figure 2 is a finite element simulation of the magnetic field generated hy a planar coil, using Femlah@ electromagnetic module software. The grey scale is a measure for the magnetic induction (maximum induction of 7 mT for a coil current of 1 A). In figure 3(a), one can see the typical form of the transverse magnetic field component generated from a single coil at a distance of 0.25 mm above the coil plane, corresponding with the position of the microheads at the bottom of the capillary. It is important here to emphasize that the shape of the magnetic field distribution depends strongly on the distance from the coil plane. In fact, as this distance becomes small, the magnetic field is affected by the \u201cgeometrical details\u201d of the coil as shown in figure 3(b). More important is the effect due to the connecting via in the center of coil, which has a dimension of about 0.5 mm in our PCB technology. Such local magnetic field singularity due to the geometrical details of the coil is no longer effective at a\u2019distance above the coil plane of typically half the size of the via area. The large and uniform magnetic field Bo, will impose a pmpnent magnetic moment (p) to the microbeads. In the presence of the coil-generated magnetic field gradient, the polarized magnetic microbeads are subjected to a magnetic force along the axis of the capillary (x-axis), given by: (1) F P x - P o ax This force is positive or negative depending if the coilgenerated magnetic field (&) is parallel or antiparalle1,to the uniform field component Bo. Therefore, by just shifting the sense of the magnetic field (BJ tlom up (parallel to the uniform field component Bo) to down, it is possible to drive the microbeads &om the center to the periphery of the coil. This single coil displacement effect suggests the combination of this effect using neighboring coils to generate time- and position-dependent magnetic forces. However, the magnetic field of a coil is localized TRANSDUCERS \u201803 The 12th International Conference on Solid Slate Sensors, Actllinors and Minosystem. Boston, June 8-12, 2003 X position (mm) Figure 4: Magnetic field of overlapping coils distributed in huo layers derived from a finite element simulation. By combining the time varying polarization of adjacent coils one can create a travelled magnetic field maximum, which will propulse the magnetic microbeads (a cloud of beads represented by the black circle) . over an &ea of the order of the coil width; simple juxtaposition of coils in an array will not enable microbead transport, hut merely allows separation of the microbeads [4]. As a solution to this problem, we propose to use this single coil displacement effect in an array geometry of overlapping adjacent coils, distributed over two functional layers of a PCB circuit. When a good overlap is realized between adjacent coils, there is never a local magnetic energy minimum in between two coils, which would hinder the transport of the microbeads. Figure 4 shows how one can combine the magnetic fields from adjacent coils properly in time to create a magnetic field maximum displacement, which Is effectively propulsing the microbeads in the capillary. Figure 5 shows video sequences on the experimentally observed transport o f a cluster o f beads inside the capillary, following the proposed actuation mechanism. Transport from one coil to a neighbouring one typically takes place on time scales of the order of 1 second." ] }, { "image_filename": "designv11_61_0000131_bf02903530-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000131_bf02903530-Figure6-1.png", "caption": "Figure 6 Schematic diagram of the experimental putter head", "texts": [ " With the shaft connected at the usual position forward on the putter head (as shown in Figure 5a), the topspin was only 50% of that obtained when the shaft connection was repositioned back from the impact face to align with the COM (Figure 5b). Thus, shifting the shaft connection to a point close to the heel\u2013toe axis (i.e. the desired rotation axis for vertical gear-effect) restored the imparted topspin to the value predicted by theory. To explore this phenomenon, an experimental putter head was made and tested. This is shown schematically in Figure 6 and comprises a 120 mm length of 32 mm square section aluminium with a standard putter shaft bonded into a 20 mm deep bore, 20 mm from the heel end, in the vertical centre-plane. Rectangular axes shown in Figure 6 follow the same system adopted in Figure 1. The putter head has equal principle moments of inertia of 415 kg mm2 about the Z and Y-axes and zero impact face loft. 90 Sports Engineering (2003) 6, 81\u201393 \u00a9 2003 isea Ballistic measurements were carried out using a smooth surfaced ball that was equivalent in all other respects to a standard balata covered golf ball. As a further precaution, the ball was placed in nearly the same angular orientation before impact, to ensure high repeatability. The aim of the experiment was to measure the velocity versus offset characteristics for heel\u2013toe offsets across the usual putter face and then compare these with the equivalent characteristics for impacts off the sole" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001397_1.2346686-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001397_1.2346686-Figure2-1.png", "caption": "Fig. 2. The rod went through the slot. Observations from the co-moving frame of the slot.", "texts": [ " 2 The relation between the angles is tan = tan , 3a tan = tan , 3b where = 1 1 \u2212 vx 2 + vz 2 /c2 . 4 In the following we consider three cases. Case I: = . An observer co-moving with the slot observes the rod landing with both their axes aligned. In this 999 999Am. J. Phys., Vol. 74, No. 11, November 2006 C. Iyer and G. M. Prabhu This article is copyrighted as indicated in the article. Reuse of AAPT content is subject to the terms at: http://scitation.aip.org/termsconditions. Downloaded to IP: 169.230.243.252 On: Thu, 26 Feb 2015 15:32:52 case see Fig. 2 the rod goes into the slot with its axes aligned and the rod is smaller than the slot all observations are from frame S . An observer co-moving with the rod frame F observes this same case differently. Consider the relation between and when = . If we incorporate the inequalities of Eq. 2 with this equality, we obtain = ; thus we have when = . In other words, when an observer in frame S observes an aligned landing, an observer in frame F observes a landing with the leading edge of the rod tilted toward the slot see Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002934_icelmach.2008.4799964-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002934_icelmach.2008.4799964-Figure4-1.png", "caption": "Fig. 4. Under load simulations isd = 1.5 A. (a) \u03bbrq 6= 0; (b) \u03bbrq = 0", "texts": [ " The convergence process is sketched in Fig. 3, where the linearity of the iron in the q\u2013axis direction is shown. The motor torque can be computed from the Maxwell\u2019s stress tensor from the field solution. Alternatively, the torque can be computed as: \u03c4 = 3 2 p\u03bbrdirq (11) The slip speed, in electrical radians per second, is achieved from (6). Multiplying both numerator and denominator by 3 2 irq, such a speed becomes \u03c9sl = PJr 3 2\u03bbrdirq (12) where PJr is the Joule loss in the rotor cage that is easily computed from the field solution. Fig. 4(a) shows the flux lines under load operating condition with isd = 1.5 A, isq = 1.5 A and \u03bbrq 6= 0. In this case, the FOC condition is not satisfied and this is evident since the flux lines are also in the q\u2013axis direction. On the contrary, Fig. 4(b) shows the flux lines under load operating condition with isd = 1.5 A and \u03bbrq = 0. Now the FOC condition is satisfied and in the rotor the flux line are only in the d\u2013axis direction. In Fig. 5 the stator and rotor flux linkages versus q\u2013axis stator current curves for isd = 1.5 A are shown. It could be observed that the FOC condition, that is \u03bbrq = 0, is kept for each q\u2013axis stator current. Moreover, the d\u2013axis flux linkages are constant for any value of the q\u2013axis current. Let us note that \u03bbrd results higher than \u03bbsd since the rotor winding is characterized by a higher number of turns" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001440_epepemc.2006.4778571-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001440_epepemc.2006.4778571-Figure1-1.png", "caption": "Fig. 1. Windings of the Double-star induction machine (DSIM).", "texts": [ " The typical structure of such systems is three-level inverter threephase electric machine system. The parallel circuit dual to the multi-level system is, essentially, the concept of the multi-phase inverter fed electric machine drive system. In a multi-phase machine drives systems, more than three phase windings are implemented in the same stator of the electric machine, and in the most common such stricture, two sets of three-phase windings are spatially phase shifted by 30 electrical degrees (figure 1). In such systems, each set of three-phase stator windings is excited by a three-phase inverter (figure 2), therefore the total power rating of the system is doubled [1]. The main advantage of multi-phase drives is the possibility to divide the controlled power on more inverter legs [2]. That will reduce the current stress of each switch, compared with a three-phase drives. Multi-phase drives possess several advantages over the three-phase ones such as: reduced torque pulsations at high frequency, reduced rotor harmonic currents, reduced current per phase without increasing the voltage per phase, higher torque per rms ampere for the same machine volume" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003984_imece2013-62657-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003984_imece2013-62657-Figure7-1.png", "caption": "Figure 7 Modeling methods for spline joint", "texts": [ " load\u2019s increasing. (2) The range of stiffness and divergence becomes wider with a maximum of stiffness, 11.9%, when there is a small load/deformation. (3) The stiffness of the spline joint seem to be more stable at big load/deformation, for both the range and the deviations of the stiffness are narrow when the force is bigger than 900N. Comparison between simulation and experiments Three modeling methods for centering surface were proposed to simulate it based on the FEM software, as shown in Figure 7. (1) Integral consolidation model: the outer and inner cylindrical centering surface is fixed. (2) Detailed contact model: based on nonlinear contact elements. Because the surface is face-face contact, contact elements, TARGE170 and CONTA174 were selected. The coefficient between the metal material was selected for the coefficient of sliding friction, the value is 0.22. (3) Parametric model: parameters of connecting part materials were optimized. The modulus of elasticity was assumed to be 70% compared with analysis above" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002096_robio.2007.4522345-Figure15-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002096_robio.2007.4522345-Figure15-1.png", "caption": "Fig. 15. Detecting steps", "texts": [ " In this system, passive RF tags (13.56 MHz) were positioned in a regular pattern such that the tags formed the vertices of equilateral triangles, as shown in Fig. 12. These RF tags were detected by a reader attached to a wheelchair, as shown in Fig. 13, in order to determine the position and direction of the wheelchair. The RF tag reader is attached beneath the wheelchair, as shown in Fig. 14. The actual operation proceeds as follows: First, the RF tag antenna is rotated to the rear-most position on the wheelchair, as shown in Fig. 15 (a); this point becomes the origin of the antenna. Next, the wheelchair is moved forward while the antenna remains fixed at the origin. The wheelchair stops at the position indicated in Fig. 15 (b) when RF tag A is detected, and it proceeds to the stop operations. The dotted circles shown in Fig. 15 indicate the ranges within which the respective RF tags can be detected. Once the midpoint of the antenna is inside one of these circles, the RF tag at the center of the circle can be detected. After detecting RF tag A, as shown in Fig. 15 (b), the wheelchair moves forward by an additional distance of approximately 0.1 m before it stops completely; however, this movement will not take the antenna outside the detection range due to the proximity RF tags used in this system. After the wheelchair stops completely, the antenna reads the information regarding the neighboring RF tags that is stored inside RF tag A. The antenna then performs one full rotation in order to read the ID numbers of the neighboring RF tags B and C, as shown in Fig. 15 (d), and to measure the antenna angles at which these tags are detected. After the antenna rotation has completed, the system computes the direction of the wheelchair using a direction estimation algorithm. In the direction estimation algorithm, it is assumed that an RF tag is positioned halfway between the angle at which the RF tag is first detected and the angle at which the tag can no longer be detected. The relationship between tags A, B, and C is shown in Fig. 16; in this figure, O represents the center of rotation" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001246_iros.2006.281695-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001246_iros.2006.281695-Figure8-1.png", "caption": "Figure 8. Definition of coordinate system and vectors.", "texts": [ " For the control method of this robot, a model based walking control method, based on ZMP criteria, is used [10]. This algorithm consists of the following main parts: 1. Modeling of the robot 2. Derivation of the ZMP equations 3. Computation of approximate waist motion 4. Computation of strict waist motion by iteratively computing the approximate waist motion Let us assume the walking system as follows: (1). The robot is a system of particles. (2). The floor for walking is solid and cannot be moved by any force or moment. (3). A Cartesian coordinate system is determined as shown in Figure 8. The X- and Y-axes form a plane identical to that of the floor. (4). The contact region between the foot and floor is a set of contact points. (5). The friction coefficient for rotation around the X, Y and Z-axes is zero at the contact point. First, we define an approximation model of the waist and position vectors, as in Figure 8. The moment balance around point P on the floor can be expressed as follows: all _ particles Ym1(r -rp)x(r +G)+T = 0 (6) If the point P is defined as ZMP, T = 0. We denote the position vector of P as PZMP (XZMP, YzMP ,0) . To consider the relative motion of each part, a moving coordinate O - XYZ is established on the waist of the robot parallel to the fixed coordinate 0 - XYZ. Q(xq, Yq' Zq) is the position vector of the origin of O -XYZ from the origin of 0-XYZ. Using the moving coordinate frame, Equation (6) can be modified as follows: all _ particles m(r-irmp)x{r+Q-G i +\u00b1x\u00b1ri42ix\u00b1x+ (o)xri)} 0O where r mp is the position vector of ZMP with respect to the 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003466_iccas.2010.5670299-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003466_iccas.2010.5670299-Figure7-1.png", "caption": "Fig. 7 Geometry of Guidance", "texts": [], "surrounding_texts": [ "Among those candidates, one trajectory can be selected by using costs. After selecting that finite optimal and admissible trajectory, just one step of whole trajectory is used for the collision avoidance of a vehicle. However, if the waypoint from the selected trajectory is applied to an actual vehicle, it may break the stability of the system because the distance between the present position of a vehicle and the desired waypoint is very close so that it cannot be regarded owing to the error from the navigation system such as GPS or INS. Therefore, a modification for the selected waypoint should be performed, and this process is based on the line-of-sight guidance law. Fig. 6 shows how this procedure is applied to the real vehicle, and it is also expressed as Eq.(8). cos ' sin waypoint waypoint waypoint waypoint R \u03c8 \u03c8 \u23a1 \u23a4 = + \u23a2 \u23a5 \u23a3 \u23a6 x x (8)" ] }, { "image_filename": "designv11_61_0000086_ptg-48029-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000086_ptg-48029-Figure2-1.png", "caption": "Figure 2. Bull Gear / HSP Configuration", "texts": [ " A verification strain gage test with proper contact showed no modulation in tooth engagement. The first reduction mesh in a turbo prop gearbox consists of a 27 tooth pinion gear and a 145 tooth bull gear. The gearbox is shown in Figure 1. The input speed is 41,730 RPM. The bull gear is connected to the sun gear of a planetary system that drives the prop shaft at 1591 RPM. The bull gear, which is laterally supported by two roller bearings, is axially positioned in the gearbox by thrust collars on the pinion gear bearings (Figure 2). There is a flat ground on the rim of the bull gear at a 15\u03bf angle from the centerline plane of the gear. This surface contacts a 1.0-inch radius on the inner ring of the pinion bearing. The contact is close enough to the pitch diameter of the gear so that the sliding velocity is similar to that of the gear mesh and an EHD film is generated to separate the surfaces. m: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx? 1 Copyright \u00a9 2003 by ASME url=/data/conferences/idetc/cie2003/72062/ on 02/06/2017 Terms of Use: http://www" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003392_9780470876541.ch5-Figure5.1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003392_9780470876541.ch5-Figure5.1-1.png", "caption": "Figure 5.1 (Continued)", "texts": [ " The instantaneous torque of the electric machine comes from the cross product of the flux linkage vector and the line vector, where the current flows. And to control the torque instantaneously, the flux linkage and the line vector should be controlled instantaneously. In this section, from the modeling of a DC machine, it is described how the instantaneous torque control is possible in a DC machine. Also, the principle of the torque control is extended to permanent magnet AC machines. And, finally, I describe how the principle is implemented in the induction machine. As shown in Fig. 5.1a, the excitation (field) flux is regulated by the current of the field winding, and the armature current, from which the torque comes through the interaction with the flux, is regulated by the armature voltage. The armature current may distort the field flux by the armature reaction. However, under the assumption that the armature reaction is fully prevented by the commutating poles and the compensation winding as shown in Fig. 2.5, the flux linkage to the armature current can be fully regulated only by the field winding current", " Hence, the instantaneous torque ofDC machine is simply the product of the magnitude of the armature current and that of the excitation flux, where the polarity of the torque is decided by the polarity of the flux and by that of the current. If the magnitude and the direction of the flux are kept constant, then torque is solely proportional to the armature current as seen from (2.20) in Section 2.3. Hence, by controlling the armature current instantaneously, the torque can be regulated instantaneously. The same torque control capability can be obtained if the relative position of the armaturewinding and field winding is maintained even if bothwindings are rotating as shown in Fig. 5.1b. The outer fieldwinding and the inner armature windingmay be rearranged as shown in Fig. 5.1c. In this arrangement, if the relative position of the two windings and the magnitude of magnetic motive force (MMF) of both windings are maintained, then the same torque control capability can be achieved. As described in Section 2.5, the equivalent MMF of outer armature winding in Fig. 5.1c can be obtained by three-phase symmetry winding as shown in Fig. 5.1d, which is a structure of the AC synchronous machine in Section 3.3. Now the armature winding is stationary, the MMF by the outer armature windings is rotating. Also, the instantaneous position of the equivalent MMF of thewinding is the decided by the instantaneous three-phase AC current in the winding. In Fig. 5.1d, to have the same position ofMMF in Fig. 5.1c, the current in a-phasewinding is null and the magnitude of current in b- and c-phase winding is identical, but the polarity of the current in b-phase winding is positive but that in c-phase winding is negative. In this way, the instantaneous torque control principle of a DC machine can be applied to an AC machine if the relative position and the magnitude of magnetic motive force (MMF) of both windings aremaintained. As seen in Fig. 5.1, the instantaneous torque control of a synchronous machine is exactly the same as in the case of a DCmachine, except the rotation of MMF of armature winding and MMF of the field winding. However, in the case of the induction machine, the instantaneous torque control is rather difficult to understand, where there is no separate field winding and no SN fi r\u03c9 ai ai fi (a) (b) ai ai fi S N fi r\u03c9 e\u03c9 Figure 5.1 Relative position of field flux and armature current. (a) Separately excited DC machine; Stationary field and armature windings (b) Rotating field and armature winding: outer field winding, inner armature winding. (c) Rotating field and armature winding: outer armature winding and inner field winding. (d) Rotating field and armature winding: stationary outer armaturewinding by threephase symmetry winding, as well as rotating inner field winding. permanent magnet. But, the rotating flux linkage, whether it is rotor flux linkage, air gap flux linkage, or stator flux linkage, can be defined" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003909_s11581-013-0956-4-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003909_s11581-013-0956-4-Figure2-1.png", "caption": "Fig. 2 Cyclic voltammograms of the CMWCNT-CPE in a 0.1 M phosphate buffer (pH 7.0) solution in the absence (a) and presence (b) of 0.25 mM AD. (c) as (a) and (d) as (b) for a CCPE, and (e) as (b) for MWCNT-CPE. In all cases, the scan rate was 10 mV s\u22121", "texts": [ " Seem to the presence of MWCNT inside the coumestan lead to significant enhancement the rate of electron transfer between coumestan and CPE which this case is well obvious from comparing this work and our previous work [31]. This is probably due to the MWCNT dimensions (of the tubes, the channels that are inherently present in the tubes), the electronic structure, and the topological defects present on the tube surface [32]. Electrocatalytic oxidation of AD at CMWCNT-CPE In order to test the potential electrocatalytic oxidation of different modified electrodes, the cyclic voltammetric responses of CMWCNT-CPE, CCPE, andMWCNT-CPE were obtained in the absence and presence of a 0.25 mM AD solution (Fig. 2). Voltammograms (a) and (b) of Fig. 2 show the cyclic voltammograms of CMWCNT-CPE in the absence (voltammogram a) and the presence of 0.25 mM of AD (voltammogram b). As expected for electrocatalytic oxidation, there was an increase in the anodic peak current of CMWCNT-CPEox/CMWCNT-CPEred redox couple in the presence of AD, and also the cathodic peak current disappears. This actually shows the efficiency of the catalytic oxidation. Under the same experimental conditions, the cyclic voltammograms of CCPE were recorded in the absence (curve c) and in the presence (curve d) of 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003745_auto.2012.1040-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003745_auto.2012.1040-Figure5-1.png", "caption": "Figure 5 The VPP concept. (a) Self-stable upright body posture in the purely mechanical roly-poly toy. (b) Transfer of the pivot-point idea to a conceptual VPP model based on the spring-mass model extended with a rigid trunk (Maus et al., 2010). (c) Application to humanoid robots, like the BioBiped robot. In order to achieve an upright trunk posture (body alignment, Fig. 4) during locomotion, the direction of leg force needs to be controlled during the contact of the segmented leg while preserving the axial spring-like leg function. Picture provided by Christophe Maufroy.", "texts": [ " g. PD control). Alternatively, also proprioceptive information from leg muscles could support upright body orientation [28]. Consider the case that the orientation of the trunk cannot be directly measured or that this information is distored (e. g. due to a vestibular disorder). Is there an alternative to keep the trunk aligned in a certain orientation with respect to the ground? Is there any mechanical strategy to solve this problem? Let us consider a simple mechanical toy, the roly-poly toy (Fig. 5a). Here the spherical shape of the lower body results in a pivot point (center of rotation) above the center of mass (COM). In the case of a perturbation from the vertical position, the COM is lifted as it rotates around the pivot point and the body responses similar to a pendulum. Consequently, the COM oscillates around the neutral position (COM directly below the pivot point) until the resting position is reached. It is important to note that a mechanical pendulum does not provide an asymptotic stable resting position without damping. Still it provides natural stability, which can easily be transferred into asymptotic stability (e. g. using damping). Here we are interested if this idea of creating a virtual pendulum can also be applied to legged locomotion. For this we extend the spring-mass model with an upper rigid body (Fig. 5b) and deviate leg force to a virtual pivot point (VPP) by applying the corresponding hip torque. This simple hip control scheme can provide postural trunk stability in both walking and running [29]. The hip torques predicted by the VPP model are similar to those observed in human walking with extending hip torques at the first half of stance and hip flexion torques after midstance. In the VPP model, the amount of hip torque depends on merely two sensory inputs: the amount of leg force and the orientation of the leg with respect to the trunk. This information can be derived based on sensory signals in legged robots, such like the BioBiped robot (Fig. 5c). The VPP-based control of trunk orientation may reduce the dependency on supraspinal pathways based on vestibular or visual information. As a result, the trunk becomes oriented vertically without the need to measure the direction of gravity or the horizontal ground. This also permits the use of the trunk as a reference for rotational leg control (e. g. leg orientation at touchdown, [26]). The hip torque patterns predicted by the VPP model may not only provide postural trunk stability but may equally support rotatory leg function (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000541_j.jappmathmech.2005.11.002-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000541_j.jappmathmech.2005.11.002-Figure1-1.png", "caption": "Fig. 1", "texts": [ " During the first two decades of the twentieth century these questions were discussed by mathematicians in the French and German scientific press. The work of the Russian mathematician, Moscow University Professor Ye. A. Bolotov [2], in which it was demonstrated that a system experiences \"impact generated by friction\"~ in the first of the above-mentioned cases, apparently remained unknown to the participants in these discussions. 1. C O N T A C T W I T H O U T E X T E R N A L F O R C E S Suppose a wheel of radius R (bodyA in Fig. 1) rotates around a fixed axis O 1 at a certain angular velocity coo. A rectangular brake pad B, which can also rotate about a fixed axis 02 (with a unilateral constraint), leans against the wheel. If the brake shoe is not in contact with the wheel, it can rotate in any direction but, when it makes contact with the wheel, it can only rotate away from the wheel. tPrikl. Mat. Mekh. Vol. 69, No. 6, pp. 912-921, 2005. ~: Professor Bolotov's work was highly appreciated by N. Ye. Zhukovskii, who not only gave it an honorable mention but also draw attention to it in his famous jubilee speech \"Mechanics at Moscow University over the Last Half Century\"", " Moreover, the thesis \"If two bodies under given conditions do not exert a pressure on one another, being absolutely smooth, then they also do not exert pressures when their surfaces are rough\" was used by Painlev6 and others. The dependence of N on the parameter fa/b, according to Eq. (1.5) is shown in Fig. 2(a). The \"intersection of two lines\" which is shown here does not occur in a number of so-called \"rough\" geometrical forms and inevitably breaks down when any \"Perturbation\" is introduced into the system of equations (1.1)-(1.3). We will no \"perturb\" the system with an additional external force P (Fig. 1) which, for simplicity, acts along the same line as N. The left-hand side of the equilibrium equation of body B (1.1) is then supplemented with a term bP which, when account is taken of the friction law (1.3), gives b N = b _ f a P (2.1) First, since the reference solution N = 0 held almost everywhere when P = 0, it seems natural to expect that we shall have a small N in the case of a sufficiently small value of P. We will now verify this assumption using formula (2.1). First, when P > 0 and the additional force clamps the brake show onto the wheel, the solution of Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000563_009-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000563_009-Figure1-1.png", "caption": "Figure 1 B, cleavage blade; C, blade holder rod; D, UHV flange adaptable to FC 38; E, three threaded removal rods at 120\"; F, removal bellows; G, bellows; H, locking screw; I, nut; J, UHV flange", "texts": [ " Besides the steric problem of the cleaving mechanism position, the specimen holder, which must allow rotation (0, @), vertical and transverse translations, heating and crystal temperature measurement, is necessarily fragile due to geometrical requirements (minimum bulkiness compatible with LEED pattern observation). Owing to this, the cleavage must be effected with a minimum stress upon the sample holder and in conditions of access to the crystal that are often difficult (anvil solid with the cleaver). Once cleavage is effected, the block anvil knife must leave the area in front of the other functions fitted on the UHV vessel. The cleaving mechanism thus must be removable. In order to meet all these requirements we have devised and achieved (figure 1) a cleaver adaptable to a UHV flange (FC 38). The block anvil A and knife B, C, which are fixed with respect to flange J, are mobile relative to the fixed flange D fitted to an FC 38 flange of the UHV vessel. Motion is ensured by means of the three threaded removal rods E. The bellows G, allowing the knife to slide within A, transmit the impulse necessary for cleavage due to their elasticity and the pressure difference. A cleaving operation thus consists of the following. Schematic drawing of the apparatus: A, anvil; (i) Setting in position of the block anvil A and blade B by means of the threaded removal rods E, by compressing the bellows F" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003386_978-3-642-39348-8_44-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003386_978-3-642-39348-8_44-Figure3-1.png", "caption": "Fig. 3 Torque-measuring rotating flange and its stator for contactless signal transmission", "texts": [ " Triaxial accelerometers are used to obtain alternating motion for each gear in 6 of freedom, performing vector composition of accelerations [12]. Analysis of gear motion can be performed in conjunction with shaft and bearing orbits analysis. Orbits are obtained using pairs of 5 mm diameter proximity probes SKF MC SS65 with sensitivity of 7.87 mV/lm, oriented at 90 and pointing towards each rotating shaft at the desired locations. Torque measurement can be performed using a rotating flange mounted on one test shaft. Signals are transmitted via a contactless interface across a gap between the rotating flange and the stator (Fig. 3), allowing misalignments to be accommodated. The torque sensor covers three functions: 1. Measure and verify torque preload retention in the recirculation loop; 2. Measure instantaneous torque transmitted by the test gears; 3. Verify tooth root strain measurements. Verification that the torque preload maintains constant average value is required to ensure that no (micro) slip is happening in the clutch flange and in the expansion bushings used to fasten the gears. It is worthwhile to point out how, besides isolating the test and the reaction sides of the test rig, flexible couplings ensure that the twist fluctuations caused by test gears TE excitation do not determine significant fluctuation of the torque preload" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001473_s00604-005-0341-8-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001473_s00604-005-0341-8-Figure4-1.png", "caption": "Fig. 4. Cyclic voltammogram of physcion. Start potential: 0.0 V (a) 1.0 10 7 mol L 1 physcion, 300 mV s 1, in the presence of oxygen; (b) 1.0 10 6 mol L 1 physcion, 10 V s 1, without oxygen", "texts": [ " A longer accumulation period was required to reach the adsorption equilibrium at a lower concentration. The experimental results showed that the peak current (Ip) of physcion increased with the increasing potential scan rate ( ). As shown in Fig. 3, there is a linear relation between Ip and in the range of 50 300 mV s 1. When >300 mV s 1, the peak current decreases with increasing potential scan rate. The above features are characteristic of an electrode reaction in which the reactant is absorbed on the surface of the electrode [15]. Figure 4 shows the cyclic voltammograms for 5.0 10 7 mol L 1 and 1.0 10 6 mol L 1 physcion after accumulating at 0.0 V for 60 s under stirring. In Fig. 4a, the cathodic peak (Pc) appears at 0.74 V in the cathodic branch. Reverse scanning produces a peak (Pa) with the same shape as Pc at the same potential in the anodic region. The phenomenon is more pronounced when the scan rate is lower. This feature is typical of a parallel catalytic reaction with reversible charge transfer [16]. The oxidant is the dissolved oxygen in the solution. The effect of the dissolved oxygen mentioned in the above section also proves this thesis. Moreover, the fast cyclic voltammogram (Fig. 4b) shows that this is a reversible redox reaction, since the cyclic curve is symmetric. The cyclic voltammogram of physcion at a static mercury electrode is shown in Fig. 5. The catalytic reaction does not yet occur in the presence of the dissolved oxygen. The pair of symmetrical peaks indicates that it belongs to a typical reversible adsorptive redox reaction. The results clearly show that the kinetics of the electrode reaction is dependent on the electrode surface properties. It is obvious from the structure of physcion that the electroactive group is quinonyl" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000182_s0263574703005071-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000182_s0263574703005071-Figure5-1.png", "caption": "Fig. 5. The structure schematic drawing of the direct-drive robotic manipulator.", "texts": [ " Since \u03081 1 (Xe, e ) =0, \u03081 2 (Xe, e ) =0, \u03081 \u03071 (Xe, e ) =0, \u03081 \u03072 (Xe, e ) =0, \u03082 1 (Xe, e ) =0, \u03082 2 (Xe, e ) =0, \u03082 \u03071 (Xe, e ) =0, \u03082 \u03072 (Xe, e ) =0, \u03081 1 (Xe, e ) = c ac b2 cos2( 2e 1e) , \u03081 2 (Xe, e ) = b cos( 2e 1e) ac b2 cos2( 2e 1e) , \u03082 1 (Xe, e ) = b cos( 2e 1e ac b2 cos2( 2e 1e) , \u03082 2 (Xe, e ) = a ac b2 cos2( 2e 1e) , the linearized equation is \u02d9\u0302 1 \u02d9\u0302 2 \u00a8\u0302 1 \u00a8\u0302 2 = 0 0 0 0 0 0 0 0 1 0 0 0 0 1 0 0 \u03021 \u03022 \u02d9\u0302 1 \u02d9\u0302 2 + 0 0 c Me b cos( 2e 1e) Me 0 0 b cos( 2e 1e) Me a Me \u03021 \u03022 (24) where \u03021 = 1 1e, \u03022 = 2 2e, \u02d9\u0302 1 = \u03071 \u03071e = \u03071, \u02d9\u0302 2 = \u03072 \u03072e = \u03072, \u03021 = 1 1e = 1, \u03022 = 2 2e = 2, and Me =ac b2 cos2( 2e 1e). Since ac>>b2, Me ac. (24) can be written as: \u02d9\u0302x=Ax\u0302+Bu\u0302 (25) where A= 0 0 0 0 0 0 0 0 1 0 0 0 0 1 0 0 , B= 0 0 1 a b ac cos( 2e 1e) 0 0 b ac cos( 2e 1e) 1 c , x\u0302=(\u03021 \u03022 \u02d9\u0302 1 \u02d9\u0302 2) T and u\u0302= (\u03021 \u03022) T. It is seen from (25) that matrix B is the linear function of cos( 2e 1e), where 2e 1e is the angle between joint 1 and joint 2 in Figure 5. This angle decides the dynamic characteristics of (25). Practically, the measured values of 2 and 1 can be regarded as the equilibrium points to linearize the system (21). Thus, along with the changes of 2 and 1, (25) can be regarded as a LPV system with respect to cos( 2 1). This article uses the values of cos( 2 1) at different angles to design the gain scheduled controller so as to improve the control performance. However, it is difficult to design a gain scheduled controller by using the polytope technique for a LPV system with a structure of (25)", " By combining (6) and (25), the augmented system can be expressed as x\u0307= A\u0303x+B\u0303u (26) 0 0 1 0 0 0 0 0 0 1 0 0 where A\u0303= 0 0 0 0 l a bl ac cos( 2e 1e) 0 0 0 0 bl ac cos( 2e 1e) l c 0 0 0 0 h 0 0 0 0 0 0 h 0 0 0 0 and B\u0303= 0 0 . 0 0 d1 0 0 d2 It is seen that after the introduction of the filter, A\u0303 becomes a linear function of cos( 2e 1e). For such a structure, the polytopic technique can be used to design the gain scheduled controller. Let = bl ac cos( 2e 1e). Such a selection of makes the system have only one varyingparameter so that the design is simplified. From Figure 5, 2e 1e [ , 0]. Thus, [ min, max]= \u2013 bl ac , bl ac . Since max max min + min max min =1, let 1( )= max max min and 2( )= min max min . Such a selection of 1 and 2 makes the system have a convex property. After a convex decomposition, the augmented system (26) can be expressed in a polytopic form as x\u0307= 2 i=1 i( )Ai x+B\u0303u (27) 0 0 1 0 0 0 0 0 0 1 0 0 where A1 = 0 0 0 0 1 a min) 0 0 0 0 min 1 c 0 0 0 0 h 0 0 0 0 0 0 h 0 0 1 0 0 0 0 0 0 1 0 0 and A2 = 0 0 0 0 1 a max . 0 0 0 0 max 1 c 0 0 0 0 h 0 0 0 0 0 0 h It is seen clearly that 1( ) and 2( ) are polytopic coordinates, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002168_isie.2007.4374753-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002168_isie.2007.4374753-Figure8-1.png", "caption": "Fig. 8. Trigonometric ratios of the sheet\u2019s rotor.", "texts": [ " It has been indicated that in those newly rotor, the sheets don't have any slot, so that in order to have them perfectly joined, and not leave air spaces between them, its shape must have a perfect definite profile. This profile is reached by the shaping with a tool or matrix to deform every piece or sheet, giving the precise shape that is described as follows. If we consider a rotor of interior radio R, formed by n sheets of a thickness e, being this thickness very small compared to R, and with an outside radio RE, we will have the trigonometric ratios in the figure 8. If we make equal the perimeter of the sheets' growth in the inside radio R, and outside radio RE, we can present an equation system such as the following: (1) Besides, from the differential triangle, knowing that both angles \"\u03b1\" are the same, because of the perpendicular sides, we can write: So That (2) If we integrate now, we get: and (3) This is an expression that relates the length of every sheet with the interior radio R and the exterior one RE. Despite the knowledge of the relationship between length, and inside and outside radios, is not enough to determine the exact shape that the sheets acquire" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003722_atee.2013.6563405-Figure10-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003722_atee.2013.6563405-Figure10-1.png", "caption": "Fig. 10. Flux density color map \u2013 steady state, no-lo", "texts": [ " The mechanical 7 helps in predicting the slip es occur and with the help of sitic torques could be deeply torque/speed curve occurs at s effect is rapidly exceeded. t the synchronous speed the sm by the reluctance torque. the rotor rotates under the e [1]. has a great influence in the achine. A great value of the gap is equivalent to a greater desirable to have an airgap s possible, which is however mical reasons [1], [7]. Fig. 8 y for two consecutive poles, degrees. the airgap flux density is x density corresponding to ircuit can be seen in Fig. 10. wo consecutive poles. The airgap for each slice separately and their mean of all slices. p flux density. The summed effect of the axial variation density in the case of skewing by a certain content of harmonics by increasing t fundamental [7]. FEM proves to be successful also in p paths and to determine the parts of the mag saturation occurs. V. RATED LOAD START-UP AND For this simulation the same conditions no-load were used. The only difference torque was set at the rated torque value, i and the momentum of inertia of the load adding it to the momentum of inertia machine, resulting in an overall momentum 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001682_inmic.2004.1492982-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001682_inmic.2004.1492982-Figure1-1.png", "caption": "Figure 1: Helicopter process configuration", "texts": [], "surrounding_texts": [ "1. Introduction and Problem Formulation\nSome nonlinear control problems can be tackled nicely using feedback linearization. There are a few well known feedbnck linearization methods. 1) Exact linearization via sfatic state feedback 2) Approximate linearization via static state feedback 3 ) Exact linearization via dyiiamic state feedback.\nA two step exact linearization of the twin rotor modcl is described in [I]. The idea is to divide the dynamics of the system into two subsystems, exact state feedback linearization of subsystem 1 is done and desired state for subsystem 2 is obtained. Then, a servo controller is designed to track the desired state. The important factor over here is delay caused by servo controller and some nonlinearity left in the system. We will consider two scenarios. In first scenario, we will use an approach similar to described in [I], In second case, subsystem 2 will be considered in steady state and dynamics of subsystem 1 will be modified accordingly. This will allow exact state feedback linearization of whole system.\nIn this paper, a simplified model o f the twin rotor system will be considered. Firstly, we will design the feedback linearizing law and then, based on resulting linear system, state feedback controller will be designed. Two scenarios are considered 1) Partial linearization of exact model 2) Exact linearization of simplified model. Simulations results are shown for both cases.\n2. Plant Dynamics and Description\nThe dynamics for the twin rotor system, considered here, are derived in [Z] for the ETH helicopter process using Euler-Lagrange approach. A similar model is also derived in [3]; and control using LQG approach is designed. A schematic of the helicopter process configuration is shown in figure I .\nThe helicopter consists of a vertical axle (A), on which a lever arm (L) is connected by a cylindrical joint. The helicopter has two degrees of fieedom: the rotation of the vertical axle (angle ) with respect to the fixed ground, and the pivoting of the lever arm (angle ) with respect to the vertical axle. Two rotors\n0-7803-8680-9/04/$20.00 02004 IEEE. INMIC 2004", "are mounted on the lever arm: R, and R,, with the resultant aerodynamic forces giving rise to moments in the B and q directions respectively. The voltages U, and u2 to the rotor motors are the inputs to the system.\nThe dynamics for ETH helicopter model are: d . - lb=d dt\nU\" - -B =e dt\nWhere\n(4)\nL, = cos28J , -2hcosBsinBml,+ h2sin2Q8m+J,\nL5 = J , h2m\nThe values for different parameters are given in [ 2 ] . Due to complexity of nonlinear terms, exact state feedback linearization of (1) to (6) is not possible. Therefore, the model i s simplified by reducing the\nAfter inserting values of various parameters, the resulting dynamics of twin rotor system are:\ni2 = 1.16 x I O - ~ X : sec(x3) + 1 . I X ~ O - ~ X ; sec(x3)\n+2x2x, tan(+) ( 8 )\nWe can divide the dynamics in two subsystems. Subsystem I contains equations (7) to (10) whereas subsystem 2 consists of equations ( 1 1 ) and (12). Subsystem 1 represents the position of twin rotor system whereas subsystem 2 represents the velocity of main and tail rotor.\n3. Analysis and Feedback Linearization\nLooking at the dynamics reveals, subsystem 1 is the only nonlinear part of the system. A linearizing feedback law is derived as described in [4]." ] }, { "image_filename": "designv11_61_0002760_ac7020486-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002760_ac7020486-Figure1-1.png", "caption": "Figure 1. Photograph (A) and circuit diagram (B) of 48-channel biamperometry array.", "texts": [ " Anal. Chim. Acta 2006, 573-574, 419-426. (24) Brock, B. J.; Gold, M. H. Arch. Biochem. Biophys. 1996, 331, 31-40. (25) Srivastava, M. K.; Ahmad, S.; Singh, D.; Shukla, I. C. Analyst 1985, 110, 735-737. (26) Matsumoto, K.; Baeza, J. J. B.; Mottola, H. A. Anal. Chem. 1993, 65, 1658- 1661. (27) Qiu, J.-D.; Peng, H.-Z.; Liang, R.-P.; Li, J.; Xia, X.-H. Langmuir 2007, 23, 2133-2137. Analytical Chemistry, Vol. 80, No. 8, April 15, 2008 2989 Instrumentation. The 48-channel biamperometry array is shown in Figure 1A. Each electrode pair consists of Pt wires pressfit into grooves present on the bottom of a Teflon stud. The bottom of the Teflon stud is a conical shape, to prevent trapping of bubbles, and the Pt wires are located on opposite sides of each stud. The wires proceed through the interior of the stud where they are electrically connected to the bottom of a 1 mm thick fiberglass printed circuit board (PCB) using standard tin-lead solder. PDMS was used to seal the interior channels for the Pt wires, to prevent liquid entry and possible sample carryover", " Currents are sampled on the PCB by the microprocessor and the ADCs, and values are sent to the computer via the communications circuit, through a standard RS232 serial cable. The PCB has its own power supply, which is connected to a standard 120 V outlet. The PCB was designed inhouse and was manufactured by Advanced Circuits (Waterloo, Ontario) using standard PCB manufacturing techniques. The electronic components were assembled on the PCB by Moreltronics (Waterloo, Ontario). A schematic of the current measurement and data acquisition circuit is shown in Figure 1B. In this instrument, each channel (electrode pair) has a maximum current of 1 \u00b5A. The applied voltage rise time is less than 1 \u00b5s, and the maximum sampling rate is 100 Hz per channel. Integration of current-time data is performed by software following data acquisition, to provide a single charge value for each run. For the applications described herein, parallel 2 min biamperometric measurements were made using an applied potential of 100 mV; measured currents were integrated between 30 and 120 s" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002480_j.jde.2009.01.014-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002480_j.jde.2009.01.014-Figure1-1.png", "caption": "Fig. 1. Left: a bead sliding without friction along a rotating bar. Right: trajectories with non-decreasing radius.", "texts": [ " In this case we have \u0393 (q, u) = { {p \u2208 R; p 0} if q = 0, {0} if q = 0. Hence, the map q\u0303(t) = t \u2212 1 provides a solution to (2.5). However, for every C 1 map t \u2192 u(t) the corresponding solution of (6.1) satisfies q(t) < 0 for all t 0. Hence the map q\u0303 cannot be approximated by smooth solutions of (6.1). Next, we illustrate a simple application of Theorems 1 and 2. Example 2. Consider a bead with mass m, sliding without friction along a bar. We assume that the bar can be rotated around the origin on a horizontal plane (see Fig. 1). This system can be described by two Lagrangian parameters: the distance r of the bead from the origin, and the angle \u03b8 formed by the bar and a fixed line through the origin. The kinetic energy of the bead is given by T (r, \u03b8, r\u0307, \u03b8\u0307 ) = m 2 ( r\u03072 + r2\u03b8\u03072). (6.2) We assign the angle \u03b8 = u(t) as a function of time, while the radius r is the remaining free coordinate. Setting p = \u2202T /\u2202 r\u0307 = mr\u0307, the motion is thus described by the equations{ r\u0307 = p/m, p\u0307 = mru\u03072. (6.3) Observe that in this case the right hand side of the equation contains the square of the derivative of the control" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003386_978-3-642-39348-8_44-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003386_978-3-642-39348-8_44-Figure2-1.png", "caption": "Fig. 2 Accelerometers carrier and its centring on the gear collar", "texts": [ " Stator and rotor of the high-resolution encoders are instead floating with respect to each other; since the rotor is mounted on the shaft it is intrinsically aligned. To align the stator, measured positioning references are taken on the low cost encoders to form a single assembly. Together with low-cost encoders, two or four uniaxial tangentially-mounted accelerometers can be used to measure the alternating TE components in dynamic conditions [15] by double integration of angular acceleration. Four seats are prepared on an accelerometers carrier (Fig. 2) to be fastened on each of the two gears being tested. Particular care has been taken to ensure symmetry of the accelerometers around the gear body, since the acceleration signals from the different accelerometers need to be combined. Measured distances from the axis of the carrier bore are accurate up to \u00b15 lm, perpendicularity of seat surfaces lies within 0.1 , coaxiality between the gear collar and the gear axis lies within 0\u20135 lm. Triaxial accelerometers are used to obtain alternating motion for each gear in 6 of freedom, performing vector composition of accelerations [12]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002900_s11465-009-0064-3-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002900_s11465-009-0064-3-Figure4-1.png", "caption": "Fig. 4 Structural dimension of axle (unit: mm)", "texts": [], "surrounding_texts": [ "\u00bdHB \u00bc \u00bdH J\n\u00fe \u00bdTT \u00f01\u00de bb \u00fe \u00bdTT \u00f02\u00de bb \u00bdTR\u00f01\u00de bb \u00fe \u00bdTR\u00f02\u00de bb\n\u00bdRT \u00f01\u00de bb \u00fe \u00bdRT \u00f02\u00de bb \u00bdRR\u00f01\u00de bb \u00fe \u00bdRR\u00f02\u00de bb\n2 4\n3 5, (12)\nwhere \u00bdH J in Eqs. (11) and (12) is the matrix of the joint region to be identified.\n3 Identification of dynamic stiffness matrix of bearing joint\nSuppose that the axle is substructure 1, the bearing housing is substructure 2, and the assembly of the axle, bearing, and bearing housing is the whole structure 3. The assembly structure is shown in Fig. 2. The bearing is simplified into a five-dimensional stiffness and damping matrix [2,3,7], as shown in Eqs. (13) and (14).\n\u00bdK \u00bc kxx kxy kxz kx y kx z kyx kyy kyz ky y kz z kzx kzy kzz kz y kz z k yx k yy k yz k y y k y z k zx k zy k zz k z y k z z 2 66666664 3 77777775 , (13)\n\u00bdC \u00bc cxx cxy cxz cx y cx z cyx cyy cyz cy y cy z czx czy czz cz y cz z c yx c yy c yz c y y c y z c zx c zy c zz c z y c z z 2 666664 3 777775 , (14)\nwhere x, y, z, y, and z denotes the x-axle direction, the y-axle direction, the z-axle direction, the y-axle rotational direction, and the z-axle rotational direction, as shown in Fig. 3. b, represents the intermediate point of two bearings, and a symbolizes the point beyond the joint region. The assembly is fixed upon the experimental table with great stiffness through the bearing housing, and the stiffness of the bearing housing is much greater than that of the bearing. Therefore, suppose that the bearing housing is stiff and fixed and, in Eq. (2), [H(2)] equals 0, Eqs. (11) and (12) can be simplified into\n\u00bdTT \u00f03\u00de aa \u2013 \u00bdTT \u00f01\u00de aa\n\u00bc \u00bdTT \u00f01\u00de ab \u00bdTR\u00f01\u00de ab h i HB \u2013 1 \u2013 \u00bdTT \u00f01\u00de ba\n\u2013 \u00bdRT \u00f01\u00de ba\n2 4\n3 5, (15)\n1 screw; 2 gland; 3 locknut; 4 left axle housing; 5 rolling ball bearing; 6 axle housing; 7 bearing housing; 8 axle", "HB\n\u00bc \u00bdTT \u00f01\u00de bb \u00bdTR\u00f01\u00de\nbb \u00bdRT \u00f01\u00de\nbb \u00bdRR\u00f01\u00de bb\n2 4\n3 5\u00fe \u00bdH J , (16)\n\u00bdH J \u00bc \u00bdK \u00fe i\u03c9\u00bdC , (17)\nwhere \u00bdH J is the FRFs matrix of the bearing joint region. Equation (15) can be further improved into\n\u00bdY \u00bc \u2013 \u00bdTT \u00f01\u00de ab \u00bdTR\u00f01\u00de ab h i \u2013 1\u00bd\u0394TT aa , (18)\n\u00bd\u0394TT aa \u00bc \u00bdTT \u00f03\u00de aa \u2013 \u00bdTT \u00f01\u00de aa , (19)\n\u00bdH J \u00bc \u00bdTT \u00f01\u00de\nba \u00bdRT \u00f01\u00de\nba\n2 4\n3 5\u00bdY \u2013 1 \u2013 \u00bdTT \u00f01\u00de bb \u00bdTR\u00f01\u00de bb\n\u00bdRT \u00f01\u00de bb \u00bdRR\u00f01\u00de bb\n2 4\n3 5, (20)\n\u00bdPJ \u00bc \u00bdH J \u2013 1: (21)\n\u00bdPJ is the dynamic stiffness matrix of the bearing joint region. Equations (18), (19), (20), and (21) are identification models of dynamic stiffness matrices of the bearing joint region, where such FRFs matrixes related to the translational degree of freedom as \u00bdTT \u00f03\u00de aa , \u00bdTT \u00f01\u00de\naa , \u00bdTT \u00f01\u00de\nab , \u00bdTT \u00f01\u00de ba , and \u00bdTT \u00f01\u00de\nbb can be directly measured through function, while such FRFs matrixes correlated with the rotational degree of freedom as \u00bdRR\u00f01\u00de bb , \u00bdTR\u00f01\u00de\nab , \u00bdRT \u00f01\u00de\nba , \u00bdTR\u00f01\u00de bb , and \u00bdRT \u00f01\u00de\nbb cannot be directly measured but indirectly estimated with finite element model, in virtue of difficulty in measuring the angular displacement and applying the moment.\n4 Experimental researches\nThe axle is suspended free of any constraint, and electrodynamic shaker is applied to excite p (p \u00bc a,a1,a2) and b in the x-axle direction, y-axle direction, and z-axle direction, respectively. A three-axis accelerometer is employed to collect the vibration data, and FRFs\nmatrixes correlated with translational degree of freedom (degree of freedom in the x-axle, y-axle, and z-axle directions) at p (p \u00bc a,a1,a2) and b are measured. Finally, FRFs matrixes, i.e., \u00bdTT \u00f01\u00de pp , \u00bdTT \u00f01\u00de pb , \u00bdTT \u00f01\u00de bp , and \u00bdTT \u00f01\u00de bb (p \u00bc a,a1,a2) are obtained. The axle, bearing, and bearing housing are then assembled into a whole, and the FRFs matrix \u00bdTT \u00f03\u00de pp of assembly at p of the axle is measured. During experiments, it is very difficult to measure the angular displacement and apply the moment. Moreover, the measurement of connection points may also be hindered by clamps or other joint mechanism. Therefore, these FRFs matrixes related to rotational degree of freedom like \u00bdTR\u00f01\u00de pb , \u00bdRT \u00f01\u00de bp , \u00bdTR\u00f01\u00de bb , \u00bdRT \u00f01\u00de bb , and \u00bdRR\u00f01\u00de bb (p \u00bc a,a1,a2) in Eqs. (18), (19), (20), and (21) can just be indirectly estimated based on the unconstrained finite element model. To provide relatively high precision for FRFs derived from the finite element model, the finite element model should be accurately calibrated. The 3D beam element is used for finite element modeling for the axle to calculate its natural frequency of the axle. Figure 3 shows the structural dimension of the axle, and Table 1 lists the property of the axle materials. The calculation results of the finite element model and the analysis results of the experimental mode are shown in Table 2, which indicates that the natural frequency calculated through finite element is of high precision. The effect of modal damping is involved in FRFs measured in experiments. To enhance the accuracy of the finite element model in calculating FRFs, the effect of damping should be taken into account in the finite element model. The first six-order modal damping ratio of the axle is obtained through modal tests, as shown in Table 3. The\nTable 1 Property of materials\nmaterials No. 45 quenched and tempered steel\nyoung modulus/Pa 2.1E11\npoisson ratio 0.3\ndensity/(kg/m3) 7850", "damping ratio of the first six-order models is introduced into the finite element model, the exciting load is imposed upon a and b in the z-axle direction, respectively, the displacement and FRFs of a and b in the z-axle direction is calculated, and the comparison is conducted between the calculated FRFs and that obtained through experiments, as shown in Figs. 5 and 6. It is necessary to know that the acceleration FRFs is obtained in the experiments, whereas the displacement FRFs is attained in finite element calculation, and the acceleration FRFs should be multiplied by \u2013 1=\u03c92 before it is transformed into the displacement FRFs (\u03c9 is angular frequency). Figures 5 and 6 show that the FRFs curve obtained by the calibrated finite element model agrees well with that measured through experiments in the frequency range of 30 to 2000 Hz. Thus, there is comparatively high precision and reliability in FRFs matrixes related to the rotational degree of\nfreedom, i.e., \u00bdTR\u00f01\u00de pb , \u00bdRT \u00f01\u00de bp , \u00bdTR\u00f01\u00de bb , \u00bdRT \u00f01\u00de bb , and \u00bdRR\u00f01\u00de bb (p \u00bc a,a1,a2), which are indirectly estimated through the calibrated finite element model. The calibrated finite element model is applied to calculate FRFs matrixes, i.e.,\u00bdTR\u00f01\u00de pb , \u00bdRT \u00f01\u00de bp , \u00bdTR\u00f01\u00de bb , \u00bdRT \u00f01\u00de bb , and \u00bdRR\u00f01\u00de\nbb (p \u00bc a,a1,a2), respectively. Then, the FRFs matrices of a and b and that of a1 and b are used to identify the dynamic stiffness matrixes of the bearing joint region, i.e., Pab\nJ and Pa1b J under the help of Eqs. (18)\u2013(21).\nThe dynamic stiffness matrix of the bearing joint region is\nPJ \u00bc \u00f0Pab J \u00fe Pa1b J \u00de=2: (22)\nEquation (22) can reduce the identification error of the dynamic stiffness matrix of the bearing joint region. More measurement points will make higher identification precision. Identification results of PJ are shown in Fig. 7. The dynamic stiffness matrix of the bearing joint region identified and FRFs matrixes of a2, i.e., \u00bdTT \u00f01\u00de a2a2 , \u00bdTT \u00f01\u00de\na2b , and \u00bdTT \u00f01\u00de\nba2 are introduced into Eq. (15) to estimate the FRFs of the freedom degree of the assembly at a2 in the zaxle direction (by launching excitation and collecting vibration in the z-axle direction, estimated FRFs), and the comparison is conducted between the estimation results and the experimental results, as shown Fig. 8. Figure 8 shows that the estimation results of FRFs of the freedom degree of a2 in the z-axle direction are in relative harmony with the experimental results, and the difference between them may be caused by the calculation error of the estimation model and experimental error. Therefore, the dynamic stiffness matrix of the bearing joint region identified, i.e., PJ, is of relatively high reliability." ] }, { "image_filename": "designv11_61_0003195_978-3-540-89393-6_21-Figure21.1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003195_978-3-540-89393-6_21-Figure21.1-1.png", "caption": "Fig. 21.1 Aerovironment\u2019s Black Widow (56 g) (a), David Liu\u2019s triplane (10 g) (b), and Martin Newell\u2019s Shark (0.495 g) (c). Figures reprinted with permission from M. Keennon, D. Liu, and M. Newell, respectively", "texts": [ " The last decade has seen an explosion in the field of micro-aerial vehicles (MAV) with the design of ever- A. Klaptocz ( ) Lab of Intelligent Systems, EPFL, Lausanne, Switzerland e-mail: adam.klaptocz@epfl.ch smaller platforms capable of flying longer and more robustly than ever before [11, 21]. The main driver in the outdoor MAV field was the Defense Advanced Research Projects Agency (DARPA), which defined the maximum dimension of an MAV at 15.24 cm (or 6 in.) and funded many successful projects. The 56 g Black Widow (Fig. 21.1a) designed by Matt Keennon and the team at Aerovironment [11] was an impressive success in 2000, optimizing the aerodynamic performance of the platform for the size constraints. Flying outdoors, however, implies high speeds and empty space, and thus recent emphasis has been on larger platforms with bigger payloads to get better autonomy, more powerful cameras, and better sensors. In confined environments, the toy and radiocontrolled (RC) hobbyist market contributed the largest advances in MAV technology. The advent of new energy sources, such as lithium-polymer (Li-Po) batteries and increasingly miniaturized actuators, fueled mainly by the RC market, has only recently yielded ultralight MAVs capable of useful mission indoors. Tele-operated flight in confined, room-sized spaces, which implies low speed and sharp turns, was first demonstrated in 2002 by David Liu\u2019s design of a 10 g triplane1 (Fig. 21.1b). The lightest indoor flyer as of 2008, the Shark, was built by Martin Newell2 (Fig. 21.1c) weighing in at a mere 0.5 g. The Shark cannot carry any payload, however, since saving weight with tricky construction was required just to have it stay airborne and remain controllable. 1 see: http://www.didel.com/DavidLiu.html. 2 see: http://mnewell.rchomepage.com. 299D. Floreano et al. (eds.), Flying Insects and Robots, DOI 10.1007/978-3-540-89393-6_21, \u00a9 Springer-Verlag Berlin Heidelberg 2009 The ability to fly indoors has many applications. Search and rescue in damaged buildings can best be done from the air, due to the large amount of debris blocking ground-based robots and the better viewpoint", " Traditional model and glider constructions with balsa ribs, spars, intrados, and extrados are too complicated and fragile at small scales. Wings cut from a sheet of foam, such as the expanded polypropylene (EPP) type used for toy planes, are a good solution for outdoor planes but is generally too heavy for indoor flyers with a small wingspan. A 3-mm-thick EPP weighs around 400 g/m2 (depending on the type of EPP). The best solution currently available is to use carbon rods for the perimeter of the wing and a thin plastic foil (such as Mylar that can weigh only 2 g/m2) as the wing surface (such as the wing of M. Newell\u2019s plane, Fig. 21.1c). Such a construction can result in a 20 cm wing of only 0.1 g, though it is unlikely that any new materials will produce significantly lighter constructions at this scale. At smaller scales microelectromechanical systems (MEMS) can be used to build small flat surfaces inspired from bat and insect wings [34]. Precision stereolithography may be a solution for 3D pieces used to link hinges, carbon rods, and MEMS. Handassembling wings becomes difficult and impractical for wings below 10 cm in size, at which point microconstruction techniques such as stereolithography to produce wing spars become more practical" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002222_978-0-387-46283-7-Figure7.12-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002222_978-0-387-46283-7-Figure7.12-1.png", "caption": "Fig. 7.12 Sliced part showing compensation zone", "texts": [ " This method entails subtracting a tailored volume (Compensation Zone) from underneath the CAD model in order to compensate for the increase in the Z dimension that would occur due to print-through. By controlling the process parameters, including the thickness of the Compensation Zone, it is possible, in theory, to eliminate the printthrough errors completely. The primary process variables under user control are: \u2022 Thickness of the Compensation Zone, given by the function Zc(x ,y) \u2022 Thickness of every layer given by the function LTk(x ,y), where LTk(x ,y) is the thickness of the kth layer from bottom \u2022 Exposure supplied to cure every layer, given by function Ek(x ,y) In Fig. 7.12, it can be seen that the thickness of every layer is not constant. For example, the bottom layer of the boss protrusion is cylindrical and has a non-zero compensation zone to compensate for extra exposure leaking through from layers above. Hence, we denote the layer thicknesses by functions of lateral coordinates as LTk(x ,y). The exposure distribution for the bottom layer of the protrusion must correspond to the desired thickness profile. The central region does not have a compensation zone since there are no layers directly above it" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000089_rspa.2002.1105-Figure11-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000089_rspa.2002.1105-Figure11-1.png", "caption": "Figure 11. Subrope profiles in the Liverpool splice.", "texts": [ " At some point in the vicinity of the splice both subropes contribute equally, and here the two subropes form a two-component twisted \u2018strand\u2019. Load transference is by friction (mode 1) between the two strands. Proc. R. Soc. Lond. A (2003) Modelling and analysis of splices used in synthetic ropes 1655 The path geometry is shown in figure 10; the subropes are twisted around each other in the anticlockwise direction and the pair of subropes is wound clockwise around the splice axis. (a) Analysis of the Liverpool splice The detail of the construction of the R\u2013S \u2018strand\u2019 is shown in figure 11. At the crotch both components are at opposite positions and are equally loaded from the \u2018strand\u2019 axis. Moving away from the crotch, and towards the rope, the R subrope becomes more loaded and the S subrope sheds load; the R subrope moves towards the R\u2013S \u2018strand\u2019 axis and the S subrope moves away; thus, the R subrope spirals into the \u2018strand\u2019 core and the S subrope spirals out. At any point between the eye and rope in this intermediate structure, the path can be determined by ensuring contact between the rope and S subropes" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001241_j.mechatronics.2005.04.001-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001241_j.mechatronics.2005.04.001-Figure1-1.png", "caption": "Fig. 1. Layout of the transverse flux linear induction motor showing each path shape.", "texts": [ " Additionally, the decouple control algorithm in which the slip frequency is uniquely determined according to the corresponding command input has been tried, using a dependency of the slip frequency on the relative ratio of the thrust force and the normal force [6]. Although the slip frequency is obtained uniquely according to the relative ratio, the irregularity meaning that the frequency value oscillates severely for a perturbation of the ratio at the transition state should be solved. As one of solutions to decouple both thrust and levitation for control index in this paper, a novel approach using DC-biased multi-phase power different from the existing input for TFLIM (transverse flux linear induction motor) shown in Fig. 1 is taken, which results in realizing the independent control of both motions. In the proposed method, the normal force is controlled by the biased DC and the thrust force by multi-phase current instead of the slip frequency of the multi-phase current and magnetic flux (is directly proportional to the current), although both inputs take a similar form in the finally impressed state. Although this method has been tried for LIM with the secondary composed of only the back yoke, a request of the thrust force was relatively very small as its application target was a conveyance of the heavy steel plate [7]", " Summarizing the primary contents discussed in this paper, the magnetic force of TFLIM is firstly analyzed through the use of the distributed-parameter magnetic field theory. And then, when DC current is biased in multi-phase input applied to the primary coil, the distortion of magnetic force is quantified. Based on the continuous solution, the time-average values of the thrust force and the normal force are converted into the lumped parameter types. Finally, the active compensation algorithm of couple between both forces and control inputs is proposed, including the experimental results. As described in Fig. 1, TFLIM making its closed magnetic path with the direction of traveling field and the distribution plane of eddy current orthogonal is a modified form of LIM developed to decrease an edge effect of the general induction motor. Particularly, each electromagnet composing its primary can be controlled to manipulate the normal force, and it is the most distinct peculiarity compared with the general LIM. In this chapter, when the multi-phase input is applied to the primary coils of TFLIM, the thrust and normal force characteristic is modeled, using the distributed-parameter field theory under a suitable assumption" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002222_978-0-387-46283-7-Figure4.25-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002222_978-0-387-46283-7-Figure4.25-1.png", "caption": "Fig. 4.25 Alumina coating, approximate thickness of 10 nm, on nominally 3 m thick polysilicon MEMS microgears. The sacrificial silicon dioxide has been removed to release the mechanism from the substrate. (b) Cross-section of the hub regions showing the contact surfaces, and (c) TEM micrograph of regions labeled (1), (2) and (3) and shown in (b). Thickness ranges from 10 nm to 10.5 nm [72]", "texts": [ " Hesketh Cantilever beams, MEMS actuators, capacitors, and resonators have been successfully coated by ALD. Coated cantilever beam structures showed no change in their radius of curvature, indicating that the coating was uniform on both the top and bottom side of the beams. A small shift in resonant frequency was observed, which corresponds to the added mass of the ALD alumina coating (see Fig. 4.24). Another example is an ALD [72] of 10 nm-thick alumina, deposited at 168\u25e6 C with a TMA precursor, at a pressure of 1 Torr (625 Pa), and coated on MEMS polysilicon micromotors for wear-resistance. Figure 4.25 shows the thickness of the layer deposited on the top surface (1). The hub (2) was 10 nm, compared with that produced under the gear (3), which was 10.5 nm. The surface roughness was 0.2 nm, indicating that adequate coating uniformity can be achieved with this process. An ALD of 4 Nano/Microfabrication Methods 91 alumina has also been applied to optical MEMS [73]. MEMS device applications are reviewed by Stoldt and Bright [74] Wear reduction via ALD tribological coatings in MEMS actuators is a highly desirable process" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001836_978-3-540-71967-0_2-Figure2.1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001836_978-3-540-71967-0_2-Figure2.1-1.png", "caption": "Fig. 2.1. The bio-inspired approach to adaptronic structures: a active materials, b induced-strain actuators, c integrated active sensors; d multifunctional composites, e microcontrollers", "texts": [ "daptronic structures (also referred to as smart materials or intelligent structures) are defined in the literature in the context of many different paradigms; however, two are prevalent. In the technology paradigm, adaptronic structures are seen as an \u2018integration of actuators, sensors, and controls with a material or structural component\u2019, see Fig. 2.1. In the science paradigm, adaptronic structures are \u2018material systems that have intelligence and life-like features integrated in the microstructure of the material in order to reduce to total mass and energy and produce an adaptive functionality\u2019. The vision and guiding analogy of adaptronic structures is that of learning from nature and living systems in such a way as to enable man-made artifacts to have the adaptive features of autopoiesis we see throughout nature. This leads to the description of the anatomy of an adaptronic material system: actuators or motors that behave like muscles; sensors that have the functionality of the five senses (hearing, sight, smell, taste, and touch); and communication and computational networks that represent the nerves, brain, memory, and muscular control systems [1]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001290_05698190590929080-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001290_05698190590929080-Figure2-1.png", "caption": "Fig. 2\u2014Schematic diagram of the cylindrical roller bearing.", "texts": [ " Such noise will be transmitted to the inside of the outer race, and then it will be D ow nl oa de d by [ U ni ve rs ity o f Il lin oi s C hi ca go ] at 2 0: 30 2 3 N ov em be r 20 14 reflected and/or transmitted to the outside boundary of the outer race. Finally, the noise radiates from the outside of the outer race. That is, it was assumed that the variation of the oil-film pressure generated between the roller and the outer race is a source of the roller-bearing noise. In other words, one of the noise sources in rolling element bearings is a transmission phenomenon resulting from the variation of oil-film pressure caused by motions between the rolling elements and races. The schematic diagram of a cylindrical roller bearing is shown in Fig. 2a. The lubrication problem between the cylindrical roller and the outer race can be considered as the lubrication problem between the equivalent cylindrical roller and the plane, as shown in Fig. 2b. In this case, the equivalent radius of the roller can be expressed as 1 re = 1 rr \u2212 1 ro [1] where re is the equivalent radius of the roller, rr is the radius of the roller, and ro is the inner radius of the outer race. For an infinitely long cylindrical roller bearing operating under zero external load, the Reynolds equation for a Newtonian lubricant in laminar flow can be written as \u2202 \u2202x ( h3 \u2202p \u2202x ) = 12 \u00b5um \u2202h \u2202x [2] where h is the film thickness, p is the oil-film pressure, \u00b5 is the oil dynamic viscosity, and um is the mean surface velocity or the entraining velocity in the x direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002200_6.2007-6442-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002200_6.2007-6442-Figure1-1.png", "caption": "Figure 1. Local vertical, local horizontal frame definition", "texts": [ " For the purposes of this work, it is not important to discuss the type of device and its properties. For our equations of motion, we will define several sets of coordinate axes. The axes we will define are the inertial reference frame (ECI) which will be denoted by {n\u0302}, the local vertical local horizontal orbital frame (LVLH) which will be denoted by {o\u0302}, and the spacecraft body frame which will be denoted by {b\u0302}. There are several definitions of an LVLH frame. The definition used in this work follows the diagram shown in figure 1. The o\u03021 axis corresponds to the radial direction of the spacecraft from the body it is orbiting, the o\u03022 axis is along the velocity vector (for circular orbits), and the o\u03023 axis is normal to the orbital plane. For this problem, these three frames are the only frames required. To analyze the attitude of the spacecraft body frame, we will define the following transformation that relates the coordinate axes of {b\u0302} to those of {o\u0302}. To represent the attitude of {b\u0302} with respect to {o\u0302}, a 3-2-1 Euler angle sequence is used" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002107_acc.2007.4283107-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002107_acc.2007.4283107-Figure1-1.png", "caption": "Fig. 1. Invariant sets for the controlled system", "texts": [ " Before proceeding to prove Theorem 4.1, we first state some lemmas that are instrumental in constructing the proof. Lemma 4.2: Assume that the trajectory of the closed-loop system (5)-(6) exists for all t \u2265 0. For any given small positive number \u01eb and any initial condition |\u03b8(0)| \u2264 R, there is a finite time T0,\u01eb such that |\u03b8(t)| \u2264 L1+\u01eb for any t \u2265 T0,\u01eb. Remark: Note that the behavior of \u03b8(t) is independent of the other two states (x(t), y(t)), and the set I\u03b8 = {(x, y, \u03b8)|\u03b8 \u2208 [\u2212L1 \u2212 \u01eb, L1 + \u01eb]} (see Fig 1(a)) can be viewed as an invariant set for the dynamics after t > T0. After the internal dynamics \u03b8(t) has been forced to evolve within the above compact set, the follower is almost aligned with the desired moving direction of its virtual agent. Then, the goal of driving all states into the origin is achieved by manipulating the translational speed and adjusting the angular velocity within a very small interval. This essentially mimics human pilot/driver\u2019s behavior in a typical leaderfollowing scenario", " Notice that since the domain of all possible values of \u03b8(t) is (\u2212\u03c0, \u03c0], under closed-loop control \u03b8(t) converges monotonically into the invariant set I\u03b8 and stays inside afterwards. Lemma 4.3: Assume that the trajectory of the closed-loop system (5)-(6) exists for all t \u2265 0. Then, the x-subsystem is ISS with respect to input y and with no restriction on initial state x(0). In particular, given any \u01eb > 0, there exists a time T1 = T1(\u01eb, x(0)) such that the norm of x(t) satisfies |x(t)| \u2264 \u03c9d M + L2 k1vd \u2212 \u01eb |y(t)|, \u2200 t \u2265 T1. Notice that the set Ixy = {(x, y, \u03b8)| |x| \u2264 \u03c9d M + L2 k1vd \u2212 \u01eb |y|} (see Fig 1(b)) can be viewed as an invariant set after t > T1. This relationship suggests that |y(t)| \u2192 0 implies |x(t)| \u2192 0 after the trajectory (x(t), y(t)) has entered Ixy . Lemma 4.4: Assume that the trajectory of the closed-loop system (5)-(6) exists for all t \u2265 0. Define the set Ixy\u03b8 = {x2 + y2 + \u03b82 \u2264 C\u22c6} \u2229 Ixy \u2229 I\u03b8. Then, for any given small positive number \u01eb and any initial condition, Ixy\u03b8 is a controlled invariant set after t \u2265 T2 = max{T0, T1} if the following inequalities are satisfied L2 > vd MC\u22c6 k2L1 > vd MC\u22c6 " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003928_icelmach.2012.6349964-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003928_icelmach.2012.6349964-Figure5-1.png", "caption": "Fig. 5. Appearance of BLH230K", "texts": [ "7g/cm^3 50H800 about permeability that there is a big difference in a magnetic steel sheet and SMC. From Fig. 4, SMC shows that it is not based on density but there is no big difference in iron loss. In addition, whereas a quadratic function the slope of a magnetic steel sheet, the slope of SMC is not seen a big change. This shows that SMC is low eddy current loss. III. TARGET MOTOR [4] The motor used by this study is brushless DC motor BLH230K made from Oriental Motor. The appearance of BLH230K is shown in Fig. 5. The rating of the motor is shown in Table IV. The motor size is shown in Table V. Induced voltage is measured in order to clarify magnetic property by the difference in the fabrication density of SMC. AC servo motor is driven at 500rpm - 6000rpm of number of revolutions, and the voltage induced between lines was measured. The measuring device of no-load property is shown in Fig. 6. The induced voltage E can be expressed as follows, when the number of turns N and magnetic flux \u03a6 are used. + = , (10) + \u221d & (11) The measurement result of the induced voltage of fabrication density 7" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000576_icems.2005.202825-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000576_icems.2005.202825-Figure4-1.png", "caption": "Fig 4 DTC Sector partition", "texts": [ " Three-phase stator voltage can be expressed as: UAN = UAO - UNO UBN = UBO - UNO , UCN = UcO UNO From (1 1),stator voltage space vector can be expressed: ( 14) 2 J27t3 14ff3 U5(t) =-[uAN + UBNe 3+ UCNe M]3 ==- EUAO +UBe I+u(Ie ] 3 where UAO UBO,UCO are the voltages between A,B,C and the center of the DC linkage, whose values are / 2-i Equ.(15) can be rewritten as: (15) Bang-bang control method are adopted in DTC algorithm. Torque and stator flux are limited in certain error scope. In practice, stator flux linkage can be limited in a round ring band ( 2A yV5 ),and stator flux chain can be divided- into six sectors: a(1), a(2)...a(6) ,as fig.4 shows. Us (Sa ISb,Sc) = -E[(d1)S,+l + (I)Sb+'ej 3 + ( 1)s+' ej4 ] 3 ,where Sa ISbhISc = 0,1. In a(1) sector, there have four distinct space voltage vectors: Us2, US3 , U55 Us6 'To useus2 ,flux chain can de- crease, and torque can decrease, to use U$3, flux chain can increase, and torque can decrease, to use35,5flux linkage increases, and torque increases, to use us6 ,flux linkage decreases ,and torque increases, as fig.5 shows. , A a] US5(F,TI Sect r t US3(1, TD) US6(FD TI) Y / X \\ \\ ,--~u.s2(FD),TD') Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000162_s0263574703005058-Figure18-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000162_s0263574703005058-Figure18-1.png", "caption": "Fig. 18. Left: Final parts of 8 homing trajectories obtained in the landmark-navigation experiments (for the configuration in Figure 17, including trajectories A and B shown there). The trajectories and final positions relate to the position of the camera axis. Circles indicate the positions where the robot stopped as it reached the termination threshold. Small dots are placed on the trajectories at points where the disparity dropped below a value of 1.5 times the termination threshold. Dashed lines depict iso-disparity curves. Right: Disparity in the vicinity of the snapshot location. Grey regions are restricted by equidistant iso-disparity curves; disparity decreases when approaching the snapshot position.", "texts": [ " At the end of the outward journey, control was handed over to the visual homing algorithm, which performed the extraction of a sectorized horizontal view from the camera image, the alignment of this view with the compass direction, and the computation of the home vector by matching the aligned view to the snapshot using the PV model. The resulting home vector was used to set the direction of movement. As soon as the disparity between snapshot and current view became lower than a threshold, the experiment was stopped (see Figure 17, right). Eight runs (including trajectories A and B in Figure 17) with different angles and distances of the outward journey were performed with this landmark configuration. The final parts of the 8 trajectories are shown in Figure 18 (left). The deviation between final and initial position of the camera axis was between 7 cm and 20 cm. This deviation does not reflect the ultimate precision of the homing mechanism, though, since the final positions are locations where the disparity reached the termination threshold, but not locations with zero disparity. In addition to the final positions, those points on the trajectories are marked, where the disparity reaches 1.5 times the threshold. For comparison, the right part of Figure 18 visualizes the disparity for points in the vicinity of the snapshot location obtained in a simulation. Final positions as well as the points where 1.5 threshold was passed are in accordance with the disparity plot, since they can be thought to be located on roughly oval bands around the snapshot position, with all final positions (except one) on the inner band. Since the band of the final positions and the band of the intermediate (1.5 threshold) positions are not overlapping, noise induced to the disparity from the sensor signal seems to be in a range below the termination threshold" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001139_iros.1992.587374-Figure15-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001139_iros.1992.587374-Figure15-1.png", "caption": "Fig. 15: Collision free movement in dynamic varying environment.", "texts": [ "re stat,ionary. For one update of the c-space model using the OCMEM-algorithm introduced for quasi-st>ationary obstacles a cornput,at,ion time of about 1.7 seconds is needed. By utilization of the OCMEM algorit-lim for dyna- 45 1 mic obstacles this t ime is reduced to less the 400 msec. in average. Thus a real-time update of t,he e-space model is reached even on a 80386 processor for comparatively slow object motion. A path-planning task in a varying environment solved by the system is illustrated by fig. 15. I t is obvious tha t the robot considers the stationary environment consisting of several complex shaped objects as well as the moving box. The average computation-time for path-planning is about 50 msec. per sampling period. The maximum time for a new path-planning which was necessary because of the moving of the object was about 1.5 seconds. ficient promptness of reaction can be reached for more rapid motion of robot and objects. -4CKNOWLEDGMENT This research has been supported by the \"Deutsche Forschungsgemeinschaft (DFG)\" " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002511_imece2009-12688-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002511_imece2009-12688-Figure3-1.png", "caption": "Fig. 3 Motor-Tachometer Setup", "texts": [ " This experimental setup is useful in illustrating concepts such as time constant, system order, linearity, and the effect of different control actions on the behavior of the system. The experimental hardware consists of a small DC motor Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/01/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use 4 Copyright \u00a9 2009 by ASME (Transicoil 1121-110 DC Servo Motor Tachometer from Servo Systems, Inc.) with a built in tachometer (see Figure 3). The control input to the motor is supplied from the PWM output of the micro controller through the H-Bridge amplifier. The speed of the motor is measured from the tachometer using the 10-bit A/D converter on the micro controller. Using the User-Interface Program, the student can do the following: Perform a calibration test to relate the steady state speed of the motor to the input voltage. This is done by selecting the Motor I/O experiment from the experiment list. This test will reveal any nonlinearities in the response such as those caused by friction" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000443_s00170-005-0009-x-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000443_s00170-005-0009-x-Figure4-1.png", "caption": "Fig. 4 A spatial parallel mechanism with 3-PUU", "texts": [ " The numerical solutions are: xc \u00bc 0:0000 yc \u00bc 0:0000 zc \u00bc 3234:5257 \u00bc 1:11023852799 63:6120: 8>< >: However, if we use the Newton-GMRES algorithm, the similar precision values can be obtained only in about 12 steps when the initial values are also give as xc=10, yc=10, zc = \u221210 and \u03b2=1, which is shown in Table 5. Generally, we can utilize the Newton\u2212Raphson method to solve the forward displacement problems. However, if it costs too many steps (e.g. more than 50 steps) or terminates for the singularity of Jacobian matrix, we had better turn to the Newton-GMRES method for help just as the above example shows. 3.2 The forward and inverse displacement of a kind of 3-UPU parallel manipulator Another kind of spatial parallel manipulator, shown in Fig. 4, is made up of 3-UPU (one prismatic joint and two universal joints) kinematic chains. Firstly, we will create an absolute coordinates oxyz as Fig. 4 shows. The plane xoy is parallel to the fixed base B1B2B3 and z-axis is perpendicular to B1B2B3. Assuming that the origin of oxyz is superposed with the geometric center and the radius of the circumcircle of triangle B1B2B3 is R, there are: B1 0 R 0\u00f0 \u00de; B2 ffiffiffi 3 p 2 R 1 2 R 0 ; B3 ffiffiffi 3 p 2 R 1 2 R 0 : If we presume that the geometric center of manipulator M1M2M3 is C, the local coordinate system ocxcycz will be shown in Fig. 4, whose original is superposed with point C. The vertexes of the manipulator are denoted by Mi,(i=1,2,3), and the prismatic pairs are denoted by Pi(i=1,2,3). In the absolute coordinate system, the vector coordinates of Mi,(i=1,2,3) are rM1 ; rM2 and rM3 ; the vector coordinates of C are (xc,yc,zc). With the similar process of 3.1, we can find the DoF of the manipulator in Fig. 4 is three. According to the screw theory and the analysis, the three orthogonal rotations of the manipulator will be forbidden by the reciprocal screws and the manipulator can only execute three independent translational movements along x, y, and z axes. Therefore, the manipulatorM1M2M3 will always be parallel to the base B1B2B3; and therefore, the transform matrix A=I, where I denotes the unitary matrix. If we presume that the radius of the circumcircle of triangle M1M2M3 is r, the coordinates of the three vertexes of the manipulator in the local coordinate system are as follows: M1 0 r 0\u00f0 \u00de;M2 ffiffi 3 p 2 r 1 2 r 0 ; M3 ffiffi 3 p 2 r 1 2 r 0 : Considering the geometry characteristics, we can obtain the following equation: rBiMi \u00bc roc \u00fe ArocMi rBi \u00bc roc \u00fe rocMi rBi; \u00f0i \u00bc 1; 2; 3\u00de (22) rBiMik k \u00bc li; \u00f0i \u00bc 1; 2; 3\u00de (23) where li,(i=1,2,3)\u2014the lengths of the limbs", "001 (Unit: 103 mm) rB1 \u00bc 0 R 0 2 4 3 5; rB2 \u00bc ffiffi 3 p 2 R 1 2R 0 2 4 3 5; rB3 \u00bc ffiffi 3 p 2 R 1 2R 0 2 4 3 5 (24) So, if the absolute coordinates of point oc are (xc yc zc), then the vector of point oc can be denoted as: roc \u00bc xc yc zc 2 4 3 5 (25) rocB1\u00bc 0 r 0 2 4 3 5;rocB2\u00bc ffiffi 3 p 2 r 1 2 r 0 2 4 3 5;rocB3\u00bc ffiffi 3 p 2 r 1 2 r 0 2 4 3 5: (26) Therefore: x2c \u00fe yc r \u00fe R\u00f0 \u00de2\u00fez2c \u00bc l21 xc \u00fe ffiffi 3 p 2 r ffiffi 3 p 2 R 2 \u00fe yc \u00fe 1 2 r 1 2R 2\u00fez2c \u00bc l22 xc ffiffi 3 p 2 r \u00fe ffiffi 3 p 2 R 2 \u00fe yc \u00fe 1 2 r 1 2R 2\u00fez2c \u00bc l23 8>>< >>: : (27) The inverse displacement solutions of the manipulator can be obtained: l1 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x2c \u00fe yc r \u00fe R\u00f0 \u00de2\u00fez2c q l2 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi xc \u00fe ffiffi 3 p 2 r ffiffi 3 p 2 R 2 \u00fe yc \u00fe 1 2 r 1 2R 2\u00fez2c r l3 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi xc ffiffi 3 p 2 r \u00fe ffiffi 3 p 2 R 2 \u00fe yc \u00fe 1 2 r 1 2R 2\u00fez2c r 8>>>< >>>: : (28) To solve the forward displacement of the manipulator, we can presume: X xc yc zc 2 4 3 5;G X\u00f0 \u00de \u00bc x2c \u00fe yc r \u00fe R\u00f0 \u00de2\u00fez2c l21 xc \u00fe ffiffi 3 p 2 r ffiffi 3 p 2 R 2 \u00fe yc \u00fe 1 2 r 1 2R 2\u00fez2c l22 xc ffiffi 3 p 2 r \u00fe ffiffi 3 p 2 R 2 \u00fe yc \u00fe 1 2 r 1 2R 2\u00fez2c l23 2 6664 3 7775: Therefore, J3 3 \u00bc G0 X\u00f0 \u00de \u00bc 2 xc yc r \u00fe R zc xc \u00fe ffiffi 3 p 2 r ffiffi 3 p 2 R yc \u00fe 1 2 r 1 2R zc xc ffiffi 3 p 2 r \u00fe ffiffi 3 p 2 R yc \u00fe 1 2 r 1 2R zc 2 64 3 75: (29) The determinate of J3\u00d73 is: J3 3j j \u00bc 8 xc yc r \u00fe R zc xc \u00fe ffiffi 3 p 2 r ffiffi 3 p 2 R yc \u00fe 1 2 r 1 2R zc xc ffiffi 3 p 2 r \u00fe ffiffi 3 p 2 R yc \u00fe 1 2 r 1 2R zc \u00bc 12 ffiffiffi 3 p R r\u00f0 \u00dezc: (30) Therefore, the singularity criterion of the Newton\u2212Raphson method to solve the forward displacement of the manipulator shown in Fig. 4 are R= r or zc= 0, which are also the singular criteria of the machanism. In fact, R> r and zc> 0, so the iterative process with Eq. 7 can be executed without any singularity. Because the manipulator shown in Fig. 4 only has three translational DoFs, we can solve its analytical solutions of the forward displacement problems directly. To obtain the analytical solutions, Eq. 27 can be equivalently transformed into: x2c \u00fe yc r \u00fe R\u00f0 \u00de2\u00fez2c \u00bc l21ffiffiffi 3 p xc \u00fe 3yc \u00bc l21 l22 R r ffiffiffi 3 p xc \u00fe 3yc \u00bc l21 l23 R r 8>< >: : (31) Considering zc \u2265 0, we can obtain the analytical solutions of the forward displacement problem as: xc \u00bc l23 l22 2 ffiffi 3 p R r\u00f0 \u00de yc \u00bc 2l21 l22 l23 6 R r\u00f0 \u00de zc \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi l21 x2c yc r \u00fe R\u00f0 \u00de2 q 8>>< >>: : (32) Therefore, if the manipulator only has translational DoFs, the forward displacement problems can also be solved with analytical method" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003613_detc2011-47030-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003613_detc2011-47030-Figure2-1.png", "caption": "Figure 2: Measurement points", "texts": [ " Some preliminary tests showed that, if the wheel is removed, some particular modes appear with the tips of the blades that modify their distance; actually these modes are not present in the bicycle owing to the presence of the hub. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 11/22/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 3 Copyright \u00a9 2011 by ASME For modal testing a tri-axial accelerometer was mounted on the tip of the left blade (point 1) and hammer impact was applied in turn on points 1, 2 and 3 of the left blade and on point 18 of the right blade, see figure 2. The tests were carried out both with out-of-plane excitation (z) and in-plane excitation (x). Figure 3 shows the measured FRFs of fork A in the low frequency range (<60 Hz), which is the most interesting from the point of view of bicycle stability and safety. Each FRF is labelled by a string of four alphanumeric characters, the first number is measurement point, the first letter measurement direction, the second number excitation point the second letter excitation direction. With out-of-plane excitation there is only one low frequency peak at 21", " The bicycles were equipped with the same slick tires inflated at 7 bar. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 11/22/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 4 Copyright \u00a9 2011 by ASME The bicycles were excited by an out-of-plane hammer impact (z direction) on point 100 near the hub of the rear wheel. Accelerations were measured by means of a tri-axial accelerometer, which was placed in turn on 10 points of the right side of the frame and on 10 points of the fork blades (see figure 2). The larger number of measurement points in the forks is motivated by the importance and complexity of this subsystem. The high repeatability of the tests assured the compatibility of the accelerations measured in the different mesh points. Results In figures 6,7,8 and 9 the overlays of the moduli of the 60 FRFs measured in the four bicycles (made from the two forks and the two frames) are reported. with fork A and frame B Frequency ranges from 0 to 130 Hz, since higher frequencies are scarcely interesting from the point of view of bicycle stability and handling" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002336_1.2999933-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002336_1.2999933-Figure5-1.png", "caption": "FIGURE 5. Sketch of the \"humping\" regime, with corresponding cross sections of melt flow at two positions (a) and (c). Typical view of melt pool (scale: 1mm) and cross-section of an hump (scale:0.5 mm).", "texts": [ " e) Welding Speeds Above 20 m/min: \"IHumping\" Regime. For our operating conditions, above 20 m/min, we reach the very characteristic humping regime that can be defined by the occurrence of a weld seam with very strong undercuts, composed of solidified large swellings of quite ellipsoidal shape, separated by smaller valleys. The melt flow that is coming from the keyhole region presents very particular characteristics [3]. As in the previous regime, the main flow is always emerging from the bottom of the keyhole front wall (see Fig. 5); it is strongly deflected rearwards and it rises up to a level that is much lower than the surface sample. It stays attached to the sidewalls along a distance of about 2 mm and then a central part of this flow is detached and forms a thin strip liquid jet that propagates rearwards at high velocity. It is along this strip liquid jet that the humps are generated at a certain distance from the detachment point [3]. In fact, one observes at a certain point, typically 2 mm from the detachment point, a shrinkage of this strip melt jet due to the Rayleigh instability that is driven by surface tension" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001154_sice.2006.315039-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001154_sice.2006.315039-Figure7-1.png", "caption": "Fig. 7 Leader and follower", "texts": [ " Note that although MPC is applied to the control of y and z in the same way, MPC is not essential in control of y and z in this example, since the constraints are not tight for small initial deviations. We also define a leader vehicle described as: Xr = vr cOS or Yr rsinOr, 0r Wr, (20) where vr and Wr are the linear and angular velocities respectively, and (xr, Yr, Or) denotes the measurable coordinate with respect to the global frame. The reference position of the vehicle i in (19) is given as a constant vector (ri, 1i) in a local frame on the leader vehicle in (20) (see Fig. 7). In other words, the reference trajectory for the vehicle i is given with respect to the global frame as (21)d .= xr + ri sin Or + li cos Or zi .- Yr - ri cos 0, + li sin 0, We refer to the vehicle i(= 1, ... , n) in (19) as the \"follower i\", and the leader vehicle in (20) as the \"leader\". Our goal is to control the each follower's position with a given offset d, defined as z. := yt j y +dsinOi 1 (22)\u00be: i yI [ + dcosi Oi to the reference trajectory z4d in (21) without collisions. We assume that a sufficient condition for collision avoidance between the followers i and j is given from the size of the vehicles as follows: z- zioc >4 , Vj#i" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001178_sice.2006.315341-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001178_sice.2006.315341-Figure3-1.png", "caption": "Fig. 3 A schematic diagram of the rotary inverted pendulum", "texts": [ " Using 5 0 0 0 Q= 0 20 0 0 , R= I 0 0 0 0 L0 0 0 0i , where A1 is Hurwitz. Therefore, if the matrix (27) is stable, the control input (26) has the low pass filter property (25) as well as stabilizes the given system (1). As an application of the passivation approach we present a solution to set-point regulation problem of a rotary inverted pendulum in the next section. INVERTED PENDULUM In order to show the effectiveness of the proposed controller, this section will show experimental results on the rotary inverted pendulum shown schematically in Fig. 3. Using the Lagrangian method, the equation of motion of rotary inverted pendulum can be derived as follows [17]: (J + Mr'2)> + mLr sin(a)d-mLr cos(a)&d = T-Bml 43mL2 _ mLrcos(o)fl-mgL sina= O (28) the LQR control gain u -Kx can be computed as Since CB = 0 and the state feedback law can be rewritten as U =-F1Cx -F2CAx (31) where F is chosen as (see Remark 1) Moreover, letting -y = 1 and R = 1 leads to an output feedback control y Lu [-2.2361 20.7462] y + u [-1.9974 2.8088] y + T1 -(7y (32) where ( > 0 should be sufficiently large to guarantee the stability of the whole system" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000266_robot.1988.12184-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000266_robot.1988.12184-Figure4-1.png", "caption": "Figure 4: Geometric description of part to be welded", "texts": [ " (v) 0 R T ~ ~ ~ TorgRTbl TblRp,tPartRsur = II. Geometric Model of the Weld Contour and the Positioning Table Figure 3, indicates the relative location of the positioning table and the part with respect to a reference frame 0 (world origin frame). The transforms are defined in the figure. Geometric Description of Weld Part We adopt cylindrical coordinates to describe the position of the weld contour on the surface of the part, with respect to a part reference frame. We assume the shape of the part is arbitrary. Figure 4, defines the variables used in the geometric modelling of the part. If partEur is a position vector which is located on the weld contour it is defined as: where r, cy and z are subject to a surface equation: P a r t ~ u r = (r cosa, rsincy, z ) ~ surf(r, cy, z) = 0 (1) (2) Let be the direction of the surface normal, and sur be the direction of the surface tangent aligned along the weld contour. The normal of the surface is then given as: c* -s* 0 S+ C+ 0 (11) 0 0 1 Here the normalization constant a" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000366_cec.2004.1330998-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000366_cec.2004.1330998-Figure1-1.png", "caption": "Fig. 1. Cooperative transportation", "texts": [ " It would be expensive to build a system that carries out optimal operation using this information. of motions: forward, backward, rightward, and leftward. The experiment was performed 10 times for each motion. Each of these experiments was repeated twice using the same robot and setup of the initial position. 111. I'ROBLEM IN TRANSPORTATION BY HUMANOID The study [ 111 shows that mutual shifts in position occur when humanoid robots perform the cooperative transportation. In the study, it is assumed that two robots use a stretcher with an object on it as shown in Fig. 1. The shifts in position may cause the following problems. The mechanical damages caused by the excessive stress applied to the arms. . The fall due to the loss of body balance. . The breakage of a transporting object. As the main cause of the problem, we can think of errors due to the manual setting of the initial positions of each joint's motors. In order to investigate the error of the initial setting, we performed several experiments in the basic mode We conducted the experiment using a humanoid robot: HOAl-1 (manufactured by Fujitsu Automation Limited)", "AI I1 . .. i . 'Y' I ( c ) Earlier trajectory (Patted). (d) Acquired ttajectory (PanemZ). Therefore, it may be possible to acquire relatively efficient behaviors by Q-learning in easier situations. However, its effectiveness will become lower as the number of states increases. VII. DISCUSSION AND COMPARISON The performance of Classifier System was compared with the one of Q-learning. The comparison was conducted for the average values of success ratios in the past I O times. As can be seen from Fig. 1 I(a). the success ratio for Qleaming is higher than the one for Classifier System. This will be explained by the hill-climbing nature of Q-leaming. However, the success ratios in the leaming process for complicated Pattem2 are almost the same as for simpler Paneml (Fig. ll(b)). Q-learning can bring out high performance in a simple environment, but that the leaming becomes more difficult in a complicated environment. Therefore, Classifier System approach is considered more suitable in order to cope with a complicated situation" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000252_robot.1992.220183-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000252_robot.1992.220183-Figure1-1.png", "caption": "Figure 1: Satellite Platform/Manipulator Arm.", "texts": [ " In this paper, we develop a numerical approach for computing near-optimal control of a platform/manipulator system. The approach is generic and works for a large class of systems known as Nonholonomic Motion Planning (NMP) systems. It does not assume any specific structure of a system, and most importantly, it can be automated computationally. We have successfully tested the numerical algorithm on several systems including the one described in this paper. +Research supported in part by NSF Grants MSS-9010900, CDA-9018673, and IRI-9003986. 2 Dynamic Model Consider the platform/manipulator system shown in Figure 1. Label the platform as body 1 and the moving links of the manipulator as bodies 2, 3 and 4, respectively. Choose coordinate frames as shown and assume that the mass center of each body coincides with the origin of the respective body frame. Denote by mi = massofbodyi, m = mi, system\u2019s total mass Ii E Px3, moment of inertia tensor of body i ri R, E S0(3), orientation of body i T E 1123, position of the system\u2019s mass center rf = ri - T, 4 i=l E @, position of frame Ci in the inertia frame 2 4 E a n d & = ( % ) , d i = ( 3 0 0 ) %, length of linki, i = 2, " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003960_14644193jmbd255-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003960_14644193jmbd255-Figure3-1.png", "caption": "Fig. 3 Terminal body force and moment definitions", "texts": [], "surrounding_texts": [ "A Newtonian approach is used to form the necessary dynamic equations for the tether model. A total of 2N vector equations are assembled where these equations will consist of N force equations and N moment equations. In order to begin forming these equations, the forces and moments acting on terminal and nonterminal bodies are shown in Figs 3 and 4, respectively. Each body has weight, Wj , and an external force, F Dj , associated with it, both defined in the inertial frame. A reaction force, \u2212Rj , on body bj , defined in the bj frame, occurs at the jth joint for all j except for the terminal body. An equal, but opposite reaction, Rj , is present on body bj+1. Similarly, a moment, \u2212Lj , on body bj , also defined in the bj frame, occurs at the jth joint for all j except for the terminal body. An equal but opposite moment, Lj , is also present on body bj+1. The only Proc. IMechE Vol. 224 Part K: J. Multi-body Dynamics JMBD255 at East Carolina University on July 8, 2015pik.sagepub.comDownloaded from limitation on Lj is that since the spin dynamics are neglected, the joint cannot impart a twisting moment. Finally, there is an external load, F A, applied to the end of the terminal body. Dynamic equations are formed by summing forces and moments for individual links with the moment equation expressed in the j body frame and the force equation expressed in the j \u2212 1 body frame. The two vector equations can be put in a recursive form where moving through the tether from the terminal link towards the root link, equations for the j \u2212 1 links contain terms from the jth link. Formation of the recursive dynamic equations is developed below first for a terminal link and then for a non-terminal link." ] }, { "image_filename": "designv11_61_0003015_cae.20257-Figure12-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003015_cae.20257-Figure12-1.png", "caption": "Figure 12 Procedure for wiring circuit of charging-system output test\u2014practical mode. (a) Question and single-choice item; (b) the question for next step; (c) the question for another step; (d) right answer and finish this job. View this article online at wileyonlinelibrary.com.", "texts": [ " The light is bright when wiring is exact; otherwise, an alarm message is automatically shown and requests learners to do it again. The right wiring procedure is positive terminal of battery! fuse (Fig. 11a)!multi-function switch (Fig. 11b)! relay (Fig. 11c)! terminals of head light (Fig. 11d)! ground. Take another example of practice mode for testing the performance of alternator (includes maximum current and a specified voltage in the vehicle lights are turn on, crankshaft speed is 2500 rpm as well as air conditioning is open). To measure charging-system output, the exact process is illustrated in Figure 12. Connect an ammeter (positive lead, \u2018\u2018\u00fe\u2019\u2019) at the alternator BAT terminal (as shown in Fig. 12a). Follow the operating instructions for the ammeter you are using, link the negative lead of ammeter and positive terminal of battery (Fig. 12b). Connect a voltmeter from the BAT terminal (Fig. 12c) to ground (Fig. 12d). In summary, this system allows learners to study and operate a practical task module for troubleshooting through the Internet. Hence, the main contributions of this system to the users, including students in the department of vehicle engineering, freshmen or novices, technicians in the automotive repair/ maintenance are: (1) providing references in detection of electrical problems to promote repairing skills for technicians; (2) building a web-based troubleshooting architecture of automotive electric that can enhance learners\u2019 knowledge and techniques; (3) developing a prototype environment for distance learning in automotive engineering education to eliminate the limitation of space and time, for example instructors and trainees in the automotive academic department and industry" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002291_bfb0119397-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002291_bfb0119397-Figure2-1.png", "caption": "Figure 2. A schematic of a typical dragline showing the main elements of the machine. (a) front view, (b) side view.", "texts": [ " Considerable work was still required to robustly find the ropes in the presence of spurious targets such as raindrops and insects (which are a t t racted to the dragline lights at night). This retains the key advantage of non-contact position sensing, the devices are rugged and low in cost, and the difficult problem of robust scene segmentation is side stepped. 2.1. D e t a i l s Our final rope finding system provides estimates of the swing angle, 8, and the hoist angle, r at a rate of 5 Hz. The system uses two of the ubiquitous PLS sensors manufactured by Sick Opto-electronics, Germany. The PLS returns range, R, and bearing, w, data across a 180 ~ field-of-view. Figure 2 shows the location of the rope sensing system on the dragline. The scanners are mounted in a weatherproof unit (Figure 3) that is located near the boom tip. Two sensors are used for redundancy to hardware failure and sun dazzle. A schematic of the hoist rope angle measurement system is shown in Figure 3. The main processing steps are: 1. Rope finding, which uses a priori knowledge to robustly discriminate rope targets. It comprises: (a) Range gating, where only targets in a specified range are accepted" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002886_med.2009.5164718-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002886_med.2009.5164718-Figure1-1.png", "caption": "Fig. 1. Small helicopter model with reference frames.", "texts": [ " The technique works for a wide class of nonlinear systems and guarantees that the output feedback controller achieves the performance of the state feedback controller when the observer gain is sufficiently high. The rest of the paper is organized as follows. Section II presents the nonlinear model of the small-scale helicopter with emphasis on modeling the main rotor flapping dynamics and stabilizing bar dynamics. In Section III we describe the controller design. Section IV provides simulation results. Finally, we draw conclusions in Section V. This section presents a dynamic model of a single main rotor and a single rear rotor helicopter equipped with a Bell-Hiller stabilizing bar, as shown in Fig. 1. The helicopter dynamics 978-1-4244-4685-8/09/$25.00 \u00a92009 IEEE 1251 is derived from first-principles to describe a six degrees of freedom rigid body model driven by forces and moments that explicitly include the effect of the main rotor, stabilizing bar and rear rotor. The reference frames used are shown in Fig. 1. The flapping frame, xF , yF , zF has its origin at the center of mass of the blade, xF is aligned with the blade chord, zF points upward and yF completes the right hand coordinate system. The rotor frame xR, yR, zR is centered at the rotor hub, xR is aligned with the blade and is rotating with it, zR is directed upward and yR completes the right hand coordinate system. The hub frame xH , yH , zH is similar to the rotor frame but is fixed to the aircraft; the rotor frame rotates around the zH = zR axis with respect to the hub frame" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003942_s13369-012-0287-1-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003942_s13369-012-0287-1-Figure5-1.png", "caption": "Fig. 5 The fourth mode", "texts": [], "surrounding_texts": [ "1. Jeffcott, H.H.: The lateral vibration of loaded shafts in the neighborhood of a whirling speed: the effect of balance. Philos. Mag. 37, 304\u2013314 (1919) 2. Dimentberg, F.M.: Flexural Vibrations of Rotating Shafts. Butterworth, London (1961) 3. Tondl, A.: Some Problems Of Rotor Dynamics. Chapman & Hall, London (1965) 4. Ruhl, R.L.; Booker, J.F.: A finite element model for distributed barometer turborotor system. J.Eng. Ind. 94, 126\u2013132 (1919) 5. Nelson, H.D.: A Finite shaft element using Timoshenko beam theory. J. Mech. Design 94, 793\u2013803 (1980) 6. Ozguven, H.N.; Ozkan, Z.L.: Whirl speeds and unbalance response of multibearing rotors using finite elements. J. Vibrat. Acoust. 106,72\u201379 (1984) 7. Greenhill, L.M.; Bickford, W.B.; Nelson, H. D.: A conical beam finite element for rotor dynamics analysis. J. Vibrat. Acoust. 107, 421\u2013430 (1985) 8. Genta, G.; Gugliotta, A.: A conical element for finite element rotor dynamics. J. Sound Vibrat. 120, 175\u2013182 (1988) 9. Mohiuddin, M.A.; Khulief, Y.A.: Modal characteristics of rotors using a conical shaft finite element. Comput. Methods Appl. Mech. Eng. 115, 125\u2013144 (1994) 10. Genta, G.: Whirling of unsymmetrical rotors: a finite element approach based on complex coordinates. J. Sound. Vibrat. 124(1), 27\u201353 (1988) 11. Jei, Y.G.; Lee, C.W.: Modal analysis of continuous asymmetric rotor-bearing systems. J. Sound Vibrat. 152(2), 245\u2013262 (1992) 12. Mohiuddin, M.A.; Khulief, Y.A.: Dynamic response analysis of rotor-bearing systems with cracked shaft. ASME J. Mech. Design 124, 690\u2013696 (2002) 13. Suh, J.H.; Hong, S.W.; Lee, C.W.: Modal analysis of asymmetric rotor system with isotropic stator using modulated coor- dinates. J. Sound Vibrat. 284(3), 651\u2013671 (2005) 14. Zorzi, E.S.; Nelson, H.D. Finite element simulation of rotor-bearing systems with internal damping. J. Eng. Power 99, 71\u201376 (1977) 15. Kane, K.; Torby, B.J.: The extended modal reduction method applied to rotor dynamic problems. ASME J. Vibrat. Acoust. 113, 79\u201384 (1991) 16. Khulief, Y.A.; Mohiuddin, M.: On the dynamic analysis of rotors using modal reduction. Int. J. Finite Element Anal. Design 26(1), 41\u201355 (1997) 17. McKee, R.J.; Simmons, H.R.: Blade Passing Related Problems in Centrifugal Pumps. Presented at the EPRI Symposium on Power Plant Pumps, New Orleans, Louisiana, March 1987 18. Guo, S.; Maruta, Y.: Experimental investigations on pressure fluctuations and vibration of the impeller in a centrifugal pump with vaned diffusers. JSME Int. J. Ser. B 48(1), 136\u2013143 (2005) 19. ANSYS User\u2019s manual, version 11.0 (2007)" ] }, { "image_filename": "designv11_61_0001616_icsmc.2004.1400778-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001616_icsmc.2004.1400778-Figure3-1.png", "caption": "Figure 3. Regions around the obstacle", "texts": [ " So we implemented an easy method of teaching the system the position of immobile obstacles by hand in advance. Here is the procedure: 1. We put an ultrasonic emitter on four comers of an obstacle in the order from I to 4 in F ig2 The obstacle is modelled as a rectangular object and its position is stored ip the system. 2. 3. Iterate the procedure 1 and 2 for all immobile obstacles. The obstacle which we modeled using this procedure is shown below. Considering that an obstacle shape is rectangle, we set each vertices of it with bo, br, bt, and b3. And we set robot\u2019s width to Wr. Fig. 3 shows four regions around the obstacle. We assign the region numbers 1 to 4 to the yellow, blue, green, and red regions, respectively. The center of the circle Oc is shown below (I). R1: oc=; i ] (1) R2: The radius of a circle r gives the following circle [\u201d -;I +(y -:I s Y2 n Y gl n Y 2 g, (7) equations (3). R3: - When bobz is defined as a straight-line g l and b3b, is defined as a straight-line g2, we obtain equations of linear formulas. - - When the robot enters RI, it rotates at that position and moves parallel to bob, until it goes out the region" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003926_s12221-011-0268-0-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003926_s12221-011-0268-0-Figure3-1.png", "caption": "Figure 3. A typical extension cycling test curve.", "texts": [ " The relevant areas for 1st and 25th cycles were measured using Qwin image analysis software available with Leica image analysis system. The permanent deformations after 1st, 10th, and 25th extension cycle were also noted down. Ten observations for each sample were taken and the average was calculated. These average values are then expressed in percentage of the A1 A n \u2013 A1 --------------- 100\u00d7 breaking extension of respective samples and termed as permanent deformation percentage. A typical extension cycling curve is shown in Figure 3. A high decay and permanent deformation would mean more loss of energy and hence poor structural integrity. The load elongation diagram of the fibres used in the study is shown in Figure 4. Both the fibres were tested at a gauge length of 25 mm. The cross head speed was 2.5 mm/min. It is evident from the figure that though polyester is stronger (14 cN) and more extendable (3.5 mm) than cotton (10 cN and 2 mm) it\u2019s initial modulus appears to be less than cotton. However, the breaking energy of polyester is definitely more than cotton" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003483_978-94-007-2069-5_16-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003483_978-94-007-2069-5_16-Figure1-1.png", "caption": "Fig. 1 Degrees of freedom of a helical element in global coordinates", "texts": [ " To determine static axial tip deflections of helical springs with large pitch angles due to the static axial force, the authors used analytical expression, which takes into account for the whole effect of the stress resultants such as axial and shearing forces, bending and torsional moments. I\u0307. Kacar ( ) \u2022 V. Yildirim C\u0327ukurova University, Adana, Turkey e-mail: ikacar@cuksurova.edu.tr; vebil@cukurova.edu.tr J. Na\u0301prstek et al. (eds.), Vibration Problems ICOVP 2011: The 10th International Conference on Vibration Problems, Springer Proceedings in Physics 139, DOI 10.1007/978-94-007-2069-5 16, \u00a9 Springer ScienceCBusiness Media B.V. 2011 119 120 PI. Kacar and V. Yildirim The global stiffness matrix, K, incorporates the pre-loading effects (Fig. 1). The global mass matrix, M, is in the diagonal form. In the solution of a large-scale eigen-value problem, the subspace-iteration method is employed. The numerical fundamental natural frequencies of such springs,\u00a8(rad/s), are presented in this work at first time. Apart from those, the critical buckling loads are obtained based on the dynamic approach. .K \u00a82M/D D 0 (1) For a helical bar, the Frenet unit vectors associated with the bar axes are t D dr=ds D . cos\u2019 sin \u2122/i C .cos\u2019 cos \u2122/j C sin \u2019 k n D ", " sin \u2019 cos \u2122/j C cos\u2019 k (2) Natural Frequencies of Composite Cylindrical Helical Springs Under Compression 121 where \u2122 is the horizontal angular displacement, \u2019 is the helix pitch angle, and the infinitesimal length of the bar is ds D .R=cos\u2019/d\u2122. Frenet unit vectors are related to each other with the following relations d t=ds D n; dn=ds D \u00a3b t; db=ds D \u00a3n (3) For a cylindrical helical bar, denoting the centerline radius of the helix by R (DD/2), the curvature is D cos2 \u2019=R, and the tortuosity is \u00a3 D sin\u2019 cos\u2019=R. The governing equations of helical rods made of an anisotropic material and subjected to the initial force To D . Po sin \u2019; 0; Po cos\u2019/ and initial moment Mo D . PoR cos\u2019; 0; PoR sin \u2019/ are obtained by Yildirim as (Fig. 1) dU ds A0 T C t x D 0 d ds D0 M D 0 (4) dT ds C .D0 M/xTo C p D 0 dM ds C .D0 M/xMo C txT C .A0 T/xTo C m D 0 where T and M are the internal force and internal moment vectors, U and are the displacement and rotation vectors, p and m are the distributed force and moment vectors, respectively. A\u2019 and D\u2019 matrices comprise the cross-sectional rigidities for composite bars [9]. Equation 4 may be put in the form of S.\u2122/ D F.\u2122/S.0/ (5) where S (0) is the state vector at \u2122 D 0. To be able to consider the variable radius of cylinder and sections, the element transfer matrix F is obtained by solving the following differential equation 12 times for 12 different initial conditions based on the complementary functions method [7, 8]. dF .m/=d\u2122 D DF .m/ .m D 1; 2; :::; 12/ (6) where F .m/ denotes the solution when mth element of the unknown vector equals unit as its other elements are all zero. These solutions compose the exact transfer matrix for cylindrical helical springs. The elements of the state vector at both ends 122 PI. Kacar and V. Yildirim for an element can be expressed by the element end displacements, di and dj, and the element end forces, pi and pj, as follows (Fig. 1): Si D fd1; d2; d3; d4; d5; d6; p1; p2; p3; p4; p5; p6gT D fdipi gT Sj D fd7; d8; d9; d10; d11; d12; p7; p8; p9; p10; p11; p12gT D fdjpj gT (7) Using the above definitions, Eq. 5 can be rearranged for an element as S.\u2122j / D F.\u2122j \u2122i /S.\u2122i / (8) The element transfer matrix can be expressed in global coordinates as Fijk.\u2122j \u2122i / D T 1.\u2122j /Ftnb.\u2122j \u2122i /T.\u2122i / (9) where T D 2 664 B 0 0 0 0 B 0 0 0 0 B 0 0 0 0 B 3 775 and B D 2 4 cos\u2019 sin \u2122 cos\u2019 cos \u2122 sin \u2019 cos \u2122 sin \u2122 0 sin \u2019 sin \u2122 sin \u2019 cos \u2122 cos\u2019 3 5 (10) In the element equation, p D kd or fpipj gT D kfdidj gT , elements of the element stiffness matrix, k, is obtained by solving Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000541_j.jappmathmech.2005.11.002-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000541_j.jappmathmech.2005.11.002-Figure6-1.png", "caption": "Fig. 6", "texts": [ " A L L O W A N C E F O R \" T A N G E N T I A L \" P L I A B I L I T Y Of course, normal pliability, when there is a tangential load, cannot correspond to the mechanical properties of all materials from which brake shoes can be manufactured. It is most likely that this is one of the limiting properties of materials. Obviously, at the other limit, there is the property of tangential pliability, to simulate which we shall also borrow the scheme proposed earlier in [4], having slightly corrected it. Suppose now (Fig. 6) that a body C of small mass m can only move along a tangent to the wheel. Between the brake shoe and the body C we place a viscoelastic spring with stiffness and viscosity k and h. The equation of motion of the body C, the position of which relative to the brake shoe we define by the coordinate x, has the form m\u2022 = - kx - h2 + T (5.1) The equation for the equilibrium of the brake shoe is a(kx + hYc) = bN (5.2) As long ask < o~R, we have T= fN and Eq. (5.1) takes the form ms = ( f b - 1 ) ( k x +hA) (5" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000743_j.talanta.2004.08.043-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000743_j.talanta.2004.08.043-Figure3-1.png", "caption": "Fig. 3. Calibration graphs for triiodide with membrane A using the following carrier solutions: (\u00a9) 0.1 mol l\u22121 sodium sulphate; ( ) 0.1 mol l\u22121 sodium sulphate, 5 \u00d7 10\u22126 mol l\u22121 sodium thiosulphate; ( ) 0.1 mol l\u22121 sodium sulphate, 5 \u00d7 10\u22125 mol l\u22121 sodium thiosulphate. The transient signals (from left to right) correspond to a concentration 5 \u00d7 10\u22125 mol l\u22121 triiodide using the three carrier solutions, respectively.", "texts": [ " It was thought that the thiosulhate would react with the triiodide present in the extremes f the sample plug and with the triiodide that had penetrated nto the membrane during exposure to the sample plug. The hiosulphate concentration must be kept very low and a FI ystem of low dispersion such as that proposed should be sed in order to keep unreacted most triiodide present in the sample plug. Two different sodium thiosulphate concentrations, 5 \u00d7 10\u22126 or 5 \u00d7 10\u22125 mol l\u22121 in 0.1 mol l\u22121 sodium sulphate were tested as carrier solutions. The corresponding calibrations are shown in Fig. 3 together with that obtained in the absence of thiosulphate in the carrier solution. The transient signals obtained for 5 \u00d7 10\u22125 mol l\u22121 triiodide are also included. As can be seen, the potential response and the linear response triiodide concentration range decreased as the thiosulphate concentration in the carrier increased. On the other hand, the time to return to the baseline decreased, thus improving the sampling frequency. As a compromise between sensitivity and sampling frequency a 5 \u00d7 10\u22126 mol l\u22121 thiosulphate concentration in the carrier solution composition was selected" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003625_aina.2010.49-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003625_aina.2010.49-Figure1-1.png", "caption": "Figure 1. The oscillation moving of sensor node A", "texts": [ " The optimal coefficient k under the constraint of T and Rs is obtained. 1 2ln s s Tk R R \u2212\u2264 (9) Oscillation moving is an unavoidable drawback when the molecule-like deploying model is used. It contributes less profit to determine the final location but wastes energy. Reducing the oscillation moving of a sensor node is crucial for saving the power consumption from mobility. Generally, when a group of sensor nodes are cast over a small region, the interior sensor nodes have higher oscillation moving probability than the exterior ones. As the example in figure 1, the sensor node A is an interior sensor node. In figure 1(a), sensor node A is pushed by sensor nodes E and F toward the direction of sensor nodes B and C. When the sensor node A is close to sensor nodes B and C shown as figure 1(b), it is pushed back toward its original position again. To tell whether a sensor node locates at the interior or exterior of a group, the Distance-Force Ratio (DFR) is introduced. In the example of figure 1, sensor A is surrounded by other nodes. Some of the repulsion forces received by sensor node A are canceled. Although sensor A receives more forces than others, its move distance is short. Therefore, the sensor node with large accumulated force but short move distance will have higher probability locating at the interior of a group. The ratio of the move distance, d, and the accumulated forces, fc, is called as the distance-force ratio (DFR) of a sensor node. T c d M f \u2265 (10) In (10), a moving threshold, MT, is used to reduce the oscillation moving" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002981_iemdc.2009.5075343-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002981_iemdc.2009.5075343-Figure2-1.png", "caption": "Fig. 2. Skew of the rotor bars bsk (unrolled view of the cage rotor)", "texts": [ " It proves to be a very useful approach to calculate the saturated zig-zag flux, and the corresponding tooth flux pulsation under load, along with the induced rotor cage harmonic currents. Hence it allows a satisfactory analytical prediction of the saturated stalled voltage-current characteristic and of the stray load losses. Weppler combined for induction machines with 2p poles the main flux saturation factor 0 < kh0 < 1, which determines \u03a6h(Im) and leads via hsw2 \u03a6\u03c0 NkfUh \u22c5= to the no-load curve Fig. 1b. Due to the skew bsk between stator and rotor slots (Fig. 2) the rotor fundamental m.m.f. is shifted to relative to the stator fundamental m.m.f. in axial direction by psk /\u03c4\u03c0b , leading to different resulting magnetizing m.m.f. values and corresponding iron saturation at different axial positions z along the iron stack. On the average along the iron stack, an increased main flux saturation occurs at load, which is included in the main flux saturation coefficient kh < kh0. The saturated main flux linkage is considered via the magnetizing inductance Xh in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000498_j.mechmachtheory.2005.09.001-Figure13-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000498_j.mechmachtheory.2005.09.001-Figure13-1.png", "caption": "Fig. 13. The definition of the internal nodes on the mechanism (x2, x3 and x4 represent the positions of the internal nodes).", "texts": [ " The error indicators based on total energy, tip displacement, tip rotation and midpoint strain are used to find the optimal positions of the internal nodes (Table 11). This section presents the implementation of the adaptive mesh on the four-bar mechanism in order to find the optimum positions of the internal nodes. Two elements of cubic shape functions on each link will be implemented in this section. We implement a non-dimensional position indicator of the internal node, with values ranging from 0 to 1. Test cases are formed for every 0.05 of the position indicator. Thus, 21 results are formed in order to find the optimum position of the internal node. Fig. 13 shows the definition of the internal nodes on the mechanism. Several cases are presented in the following subsections to show the performance of the adaptive mesh. 516 Y.L. Kuo et al. / Mechanism and Machine Theory 41 (2006) 505\u2013524 Table 12 shows the optimum positions of internal nodes on each link based on some error indicators, such as six global variables (/i, i = 2,3, . . ., 7), three midpoint strains (ei, i = 2,3 and 4, representing the input, coupler and output links, respectively) of each link and total energy (E) of the mechanism" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003809_s0025654411030113-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003809_s0025654411030113-Figure1-1.png", "caption": "Fig. 1.", "texts": [ " ,m) and h are constants and \u03be(t) is a Gaussian white noise. The behavior of the system will be analyzed by the statistical modeling method based on numerical solution of differential equations (the Runge\u2013Kutta method) combined with the numerical method for obtaining realizations of random stationary processes. To illustrate the proposed method, consider the transverse vibrations of a thin rectilinear viscoelastic plate hinged along all of its edges and subjected to a uniformly distributed load applied in the plate plane to two opposite edges (Fig. 1). It is assumed that in the course of the strain process the opposite plate edges can move in the direction of the x1- and x2-axes but remain parallel to each other. If the plate material is isotropic with Poisson ratio \u03bc constant in time, then, using the Kirchhoff\u2013Love hypothesis, we can write out the equations of motion of the plate in the case of finite deflections, which generalize the well-known Karman equations, in the following form: D(I \u2212 R)\u22074w \u2212 h(\u03a6,22w,11 \u2212 2\u03a6,12w,12 + \u03a6,11w,22) = \u2212\u03b3w\u0308 \u2212 kw\u0307, (3" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003148_2009-01-0049-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003148_2009-01-0049-Figure1-1.png", "caption": "Fig. 1 Steering and Vehicle Model", "texts": [ "(6)(7) This paper details the construction of an EPS control that compensates for the influence of vehicle dynamics on the steering system based on a novel approach. It also describes verification of the effectiveness of the control by simulations and actual vehicle tests using steering torque input, as well as by subjective evaluations from drivers of vehicles equipped with the control. SAE Int. J. Passeng. Cars - Mech. Syst. | Volume 2 | Issue 1 239 To observe the influence of vehicle dynamics on the steering system, the extremely simple steering and vehicle model shown in Fig. 1 was used. The equations of motion are shown in equations 1 to 7 below after Laplace transformation. )7( )6( )5(1 )4()( )3()( )2( )1()( 2 2 swsatcmpssw s s fsat r rr f ff rrffz rf TTTsCsI N N FT V rlKF V rl KF lFlFsrI FFrsVm =++\u22c5+\u22c5 = = +\u2212= \u2212\u2212= \u2212=\u22c5 +=+\u22c5 \u03b8\u03b8 \u03b8\u03b4 \u03be \u03b2 \u03b2\u03b4 \u03b2 Here, equations 8 to 12 were adopted to make the above equations easier to visualize. In addition vehicle yaw moment and cornering power were normalized, and IZN was set to 1 for simplicity. In the equations below, Tcmp is the torque that compensates for the influence of vehicle dynamics" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003364_978-94-007-5125-5_1-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003364_978-94-007-5125-5_1-Figure3-1.png", "caption": "Fig. 3 Kinematic scheme of wheel-legged robot suspension", "texts": [ " Its function is to move the wheel relative to the robot chassis in order to enable both driving (rolling and turn) and walking so that the robot can negotiate obstacles on its way as well as to enable the levelling of its chassis when the robot travels on a bumpy surface. The way in which walking and chassis levelling (keeping the chassis at a constant elevation from the ground) are executed is schematically shown in Fig. 2. The robot suspension structure was designed as part of a wheel-legged robot construction project being carried out in Wroc\u0142aw University of Technology. A kinematic scheme of the wheel suspension is shown in Fig. 3. The basic dimensions of the wheel suspension, shown in Table 1, were matched using a special method of geometrical synthesis and a genetic algorithm [4, 6]. During travel the mobile robot chassis moves vertically and horizontally relative to the travel path. In the suspension model the vertical movement of the chassis was obtained by placing the chassis on linear guides mounted on a base [1]. In order to run simulations a computational model of the suspension, shown in Figs. 4 and 5 was created in the LMS DADS dynamic analysis system [2]", " The suspension model will be used in dynamic and kinematic computer simulation studies of the levelling motions of the robot suspension system. The aim of the studies is to design a suitable control system ensuring the automatic maintenance of a constant elevation of the chassis above the ground while travelling on a bumpy surface. In the case of the computational robot model, the levelling function boils down to keeping a constant elevation of the robot chassis above the ground according to the schematic shown in Fig. 2b. This function can be effected solely by lifting the chassis by means of lift actuator qp (Fig. 3) while the other drives (protrusion, turn and rolling) remain fixed. First, simulation studies of chassis lifting were carried out. A schematic of the simulation is shown in Fig. 4. The aim of the studies was to determine the dependence between chassis elevation hk above the ground and lifting actuator extension qp. Location zA of suspension rotational couple A (Fig. 3) was assumed as chassis elevation hk. Figure 7 shows the obtained graphs of chassis elevation hk and the change in active force Fp in the lifting actuator versus lifting actuator extension qp. The range of change in actuator extension qp is determined by the basic suspension specifications shown in Table 1. For such a actuator its chassis elevation hk was found to range from 0.44 to 0.7 m. Only one lifting actuator qp is used to level the robot chassis (keep it at specified constant elevation hz k) during travel on a bumpy surface, which requires such excitation force Fp of the actuator which results in prescribed elevation hz k with the required accuracy" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003411_s11771-012-0984-7-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003411_s11771-012-0984-7-Figure4-1.png", "caption": "Fig. 4 Photograph of spindle", "texts": [ " Thus, it is expected that the device will reduce manufacturing and operating costs because it makes possible to remove hydraulic systems or electric equipments used in the conventional variable preload device. A spindle is fabricated for installing an automatic variable preload device and evaluating its performance. Figure 3 shows the conceptual drawing for fabricating a spindle in which four angular contact ball bearings are used. Two sets of bearings are installed at both the front and rear of the spindle, respectively. Auxiliary bearings, which support the shaft, and main bearings, are installed at the rear and front sides, respectively. Figure 4 shows the spindle fabricated in this work. Figure 5 shows the installation of an automatic variable preload device on the spindle. J. Cent. South Univ. (2012) 19: 150\u2212154 152 For the performance evaluation of the spindle, the vibration and noise according to the installation of an automatic variable preload device were compared and analyzed. A device used to measure the vibration of the system was Vibrometer (VL-8000) by HOFFMANN that presents a resolution of 0.1\u2212200 mm/s, a vibration range of 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003633_s10846-012-9726-1-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003633_s10846-012-9726-1-Figure5-1.png", "caption": "Fig. 5 The schematic showing the UAV in trim. a Level flight. b Climbing flight", "texts": [ " Deviation in heading can be neglected from the trim condition for the longitudinal dynamic analysis. A UAV is called in trim if the resultant moment is 0. Roughly, at steady level flight, we have Fthrust = Fdrag_T(Va, \u03b4e, CD, \u03b1) mg = Flift_T(Va, \u03b4e, CL, \u03b1) (4) where \u03b1 is the angle of attack, CD and CL are the drag and lift coefficients; Fdrag_T and Flift_T are the total drag and lift force, respectively, acting on the UAV\u2019s center of gravity (CG). At other attitude, the flight path angle, \u03b3 , should be considered, as shown in Fig. 5b, Fdrag_T + Flift_T sin(\u03b3 ) = Fthrust \u00b7 cos(\u03b3 ), Flift_T cos(\u03b3 ) + Fthrust \u00b7 sin(\u03b3 ) = mg. (5) There are usually multiple combinations of throttle percentage, \u03b4t, elevator deflection, \u03b4e, and the angle of attack, \u03b1, to keep the airplane flying at a desired unchanging attitude and airspeed. However, for steady level flight, there is a unique combination of \u03b4t, \u03b4e, and \u03b1, corresponding to the particular airspeed, calculated by Eq. 4. A brief trim condition is listed in Table 2. The trim conditions for different attitudes at various airspeeds can be represented in the form of (Va, \u03b4t, \u03b4e, \u03b1) pairs" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003633_s10846-012-9726-1-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003633_s10846-012-9726-1-Figure6-1.png", "caption": "Fig. 6 Interpretation of the related angles", "texts": [ " The trim conditions for different attitudes at various airspeeds can be represented in the form of (Va, \u03b4t, \u03b4e, \u03b1) pairs. 3.2 Longitudinal Dynamic Analysis State variables considered in longitudinal analysis involves Xl = [x z u w q \u03b8 ]T . Ailerons and rudder on Titan are only used for lateral attitude control, thus the control inputs for longitudinal analysis are solely \u03b4t and \u03b4e. (i) Longitudinal Equations of Motion The mounting angle of the wing on Titan is \u03b1\u0304 = 2\u25e6. Following aerodynamic convention the total angle of attack is \u03b1 = \u03b8 + \u03b1\u0304 \u2212 \u03b3 , which is interpreted in Fig. 6. Under steady level flight, the elevator at trim is defined as ZDP, \u03b4e = \u03b4e0, and the fuselage has no pitch, \u03b8 = 0\u25e6. The trim condition at Va = V\u2217 g requires \u03b4t = 40%. The analysis proceeds based on this nominal working point. From aerodynamics, lift force, drag force, and pitching moment are expressed below, Flift = 1 2 \u03c1V2SCL (6) Fdrag = 1 2 \u03c1V2SCD (7) M = 1 2 \u03c1V2ScCm (8) in which S is the wing area, c is the average chord length, Cm is the pitching coefficient and \u03c1 is the air density which usually ranges from 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000620_0094-114x(75)90059-2-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000620_0094-114x(75)90059-2-Figure1-1.png", "caption": "Figure 1. Angle-pairs.", "texts": [ " I - - F identify the individual relative positions, e.g. P3-P4 indicates that the third and fourth relative positions are finitely separated. A detailed presentation of the equations needed in computation is given in Appendices 1-3. Appendix 1 lists the general equations, Appendix 2 relates to the four precision-point problems and Appendix 3 to the five precision-point problems. In order to be consistent with the usual convention of complex numbers, counterclockwise angular quantities are assumed positive. In Fig. 1 the starting position of the four-bar linkage AoABBo is designated by the crank angles 6o and 60. Successive angular displacements of the two cranks AoA and BoB from the starting position are denoted by 62, 62; 63, 63; etc., 6J being coordinated with 6J. An angular displacement pair 6J, 0r will be called an \"angle-pair\" for short. The function generator is said to be synthesized with four precision-points if three angle-pairs are coordinated, and with five precision-points if the number of angle-pairs is four" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000035_robot.1999.772527-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000035_robot.1999.772527-Figure1-1.png", "caption": "Figure 1: The system under consideration.", "texts": [], "surrounding_texts": [ "For a robot t o work in distributed environment, it should consider only the status of the task and rely upon own sensors to estimate the future behavior. Therefore, it is desired to model a robot in contact with the object, and control all of its motions depending upon the motions of the object only. 3.1 Figure 2 shows the model of an i-th ME in contact with an object (not shown). As in micro systems, the motions are very slow, hence the dynamics of the system can be neglected. The end-effector/finger forces ffz can be simply given as Forces Generated by an ME f f z = [Fm, FstIT (1) where F, is the motor force in the active joint and F,, is the force due to spring-damper assembly in the passive joint. The forces along z-axis and y-axis of a universal frame CO are then known to be O F , , = Fmzcosq5, - Fs,sinq5, (2) O F y t = F,,,,sinq5, + Fs,cosq5, (3) where 4, = Ol0 + 0, is the angle of the push rod with respect t o the reference frame, while 8, is the absolute angle (in robot frame), and t9io is rotation of the robot frame in the world. 3.2 Forces on an ME Consider n MEs cooperating for an object manipulation as shown in Figure 3. As robots realize only point contacts with the object, hence no local moments will be generated a t the finger-tips. Therefore, the MEs only need t o apply forces but no moments. These forces, however, collectively generate the desired moments in the object. From this observation, it is clear that the control problem is reduced to determination of only two variables instead of three if the system is controlled in a distributed fashion . The finger-tip forces for each ME are known using equations (1) through (3). Consider that the i-th ME is fixed in position, the forces available at its finger-tip will be the resultant of the forces applied by all other MEs, i.e. I T f f 1 = [ 2% k 0 F y k ; IC # 2 . (4) k=1 k = l These will be the forces sensed by a force sensor if one is installed on the finger-tip. 3.3 Finger-Tip Trajectory Prediction For an ME to work independently, it should also predict its finger-tip trajectories. This is also important for the reason that an ME can know only two variables, i.e. its own F, and Fy, while three variables are desired corresponding to the three object motions. Therefore, it is needed to restrict the motions of an ME to its own controller only, instead of the object motions. However, an ME needs to consider the object's trajectories while predicting its own. Consider again an ME in contact with the object as seen by the reference world frame, shown in Figure 4. As the motions of the object are considered in the world frame, hence the finger-tip trajectories can be found easily, knowing the position of the finger-tip with respect to the object. This is true when there is no slip a t the contact-points. Let \"pa = [ O x a '%IT be the position of the center of mass of the object and & be the rotation of the object frame E, with respect to the world frame, i.e. object's orientation. Then the finger-tip trajectory of i-th ME can be found easily as Opfz = \"pa + \"%(?la> \" l f z ( 5 ) where 'I&(+,) is the rotation transformation matrix of C, through an angle w.r.t. C O , and \" 1 f Z is the position vector of the a-th finger-tip with respect to the object frame C,, it remains unchanged if there is no contact-point slip. Hence corresponding to a reference object posture [\"z, f an ME can predict its own reference finger-tip trajectory \"pf ,. The individual trajectories of all MEs will then define a resultant trajectory same as the object's desired trajectory. 4 Distributed Event-Based Control Now for control, an ME should exert the proper forces which may take part in the successful manipulation of the object in cooperation with the other MEs in the system, utilizing the constraints developed by them. There are three different types of forces which need to be determined. 4.1 To stop the undesired motions of the object expected due to the forces applied by all other MEs, given in equation (4), it is desired for the i-th ME to apply exactly similar but opposite forces as Forces to Balance the Object \" f f i = - f f i - (6) 4.2 Grasping Forces Developed by Tra- The finger-tip trajectories defined in equation ( 5 ) can only define the finger-tip position at the surface of the object. It is not possible for an ME to exert the forces on the object required to grasp it if forces are small enough in the start , for example. Therefore, a realistic approach is to modify the finger-tip trajectories such that the MEs can exert sufficient forces depending on the stiffness of the object. For this purpose a \"grasping factor ( I C g r o s p ) \" is defined which reduces the dimensions of the position vector ,1fz such that jectory Modification \" l f z (1 - k g r a s p ) \" Z f z . (7 ) This grasping factor can be a constant related to an initial desired magnitude of internallgrasping forces and the stiffness of the object. It can also be controlled depending upon variations in internal forces. \" I f z is specially selected for this purpose as it always links a finger-tip to the center of mass of the object. Hence any imbalance in the forces will a t the most change the magnitude of grasping forces but there will be a little chance of instability of the system or loss of grasp. The forces exerted by an ME on an object of stiffness pa which will contribute in developing a grasp with desired internal forces can be stated as f z , g r a s p = pa(,lfZ - \"If,>- (8 ) The modified trajectories will then define an object trajectory same as before but will adjust the grasping forces only. This process is illustrated in Figure 5 . 4.3 Forces for Motions of the Object Finally force components are required which can generate the desired motions in the object. These forces are simply computed by multiplying the fingertip trajectory errors AOpfz with some gain matrix K f , having the components of gain for both z and y motions, and it can be written as (9) A simple PID controller can be used for determination of these forces. 4.4 Desired Finger-Tip Forces ponents of desired actuator forces are known as Using equations (2), (3), (6), (8) and (9), the com- 4.5 Interpretation of an Event In this paper, a point-to-point task is interpreted as an event. Normally a complete manipulation task is divided into many small PTP tasks, as shown in Figure 6. For an event-based controller, this PTP task becomes a main command. Controller is supposed to move the robot towards minimizing the error in the reference event command and the actual robot status irrespective of whatever amount of time is spent for this job. 4.6 The Controller Figure 7 shows the block diagram of the distributed event-based controller for the proposed scheme. The main functional blocks are explained below. 0 Lookup Table: The proposed scheme utilizes a supervisory control type architecture in which the supervisor does nothing but manages the communication of necessary data from and to the individual MEs. Moreover, it also computes the current object posture and internal forces, and communicates this data back to the individual MEs. In this sense, it only acts as a lookup table for the robots from where they can access the control parameters and can write back the results of their own executions. 0 Finger-Tip Trajectory Generation: This block of an individual robot controller accesses the object posture and reference manipulation commands from t he L - . - Event Count Figure 6: An event. Reference Object Posture, Internal j Forces Current Posture of the Object Internal Forces Determination supervisor and computes the reference finger-tip trajectory utilizing equation (5). 0 Trajectory Modification for Internal Forces: This block modifies the desired finger-tip trajectories for development of proper grasping forces using the illustrations of Figure 5, as defined in equation (8 ) , if there arises a need of it. Normally, this block is needed only in start of manipulation when forces are not known exactly. Moreover, it is also needed if internal forces are to be controlled throughout the manipulation. This block computes \u201c l f 2 using equation (7) and finally computes the modified trajectories using equation (5), again. Finger-Tip Position: This block accesses data from the sensors, i.e. the length of push rod 1, and angle of rotation of robot\u2019s base 8, and geometrically computes the actual current finger-tip position. 0 Controller: This is an event-based PID controller which generates the forces for motion, using equation (9). At the output of this block, both forces for motion and grasping are available as it accepts already modified reference trajectories. 0 Finger Force Computation: If a force sensor is not available a t the finger-tip, equation (4) is used to compute the resultant opposing forces a t the finger-tip due to all other hfEs. 0 Gain: This block converts the desired finger-tip force components of equation (lo), computed at the node before it, to a resultant motor force command. Then this command is processed with factors dependent on motor/actuator parameters and consequently a motor torque command is achieved. The components inside the dashed enclosure define an individual robot\u2019s controller. All the robots (LIES) realize a same controller and then all these controllers are supervised by the lookup table. Each controller owns a communication/event tag, e.g. the event count, which tells the other controllers I Controller for ROBOT I d s h Computation 0 e e Controller for ROBOT i t Figure 7: Block diagram of the controller. and the supervisor that it has completed its job for the current event and is ready to communicate the results and access new event parameters. This tag is incremented at the completion of each event, i.e. a PTP motion, as commanded by the interpolated object trajectory." ] }, { "image_filename": "designv11_61_0002984_s00170-008-1573-7-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002984_s00170-008-1573-7-Figure3-1.png", "caption": "Fig. 3 Roll, pitch, yaw angles", "texts": [ " It consists of six identical extensible legs, connected to the mobile platform and fixed base by spherical joints at points Bi and Ai, where i=1,2, 3,\u20266. The legs are made up of two elements connected by a prismatic joint. Two Cartesian coordinate systems are attached to the mobile platform and fixed base. The position vector p of the centroid P and rotation matrix ARB of the mobile platform describe the transformation from the mobile platform to the fixed base. The rotation matrix is expressed in terms of roll \u2018\u03b8\u2019, pitch \u2018\u0444\u2019, and yaw \u2018= \u2019 angles, as shown in the Fig. 3. ARB \u00bc cosy cos f cosy sin f sin q siny cos q cosy sin f cos q \u00fe siny sin q siny cos f siny sin f sin q \u00fe cosy cos q siny sin f cos q cosy sin q sin f cos f sin q cos f cos q 2 4 \u00f01\u00de Vector loop equation for ith leg of the manipulator is as follows: ai \u00fe di \u00bc p\u00feA RBbi: \u00f02\u00de The length of the ith leg is obtained by taking the dot product of the vector di with itself. 3.1 Optimization criterion Surface finish and machining accuracy of a work piece mainly depends on the stiffness of the hexapod machine tool, which in turn depends on its configuration" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003632_iros.2013.6696529-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003632_iros.2013.6696529-Figure5-1.png", "caption": "Fig. 5. Omnicopter MAV prototype", "texts": [ " For on-board sensing and position control, we are currently using a sonar sensor (MB 1200, MaxSonar) to measure the distance from ground, and infrared sensors (GP2Y0A02YK0F, Sharp) to navigate the Omnicopter away from obstacles. The control board reads in the sonar and infrared sensors on its ADC port and also outputs PWM signals to the ESCs and servos. A GPS is incorporated for outdoor navigation. A component level breakdown of the major parts of the system and how they communicate with each other is depicted in Fig. 6. An Omnicopter MAV prototype has been constructed as shown in Fig. 5. The prototype, including a 2700 mAh battery, weights 1.4 kg and measures 45 cm from a ducted fan hub to another. Custom mounts for each of the ducted fans were 3D printed out of ABS plastic. For actuation, we are using two 920 Kv motors to drive the two central propellers of size 10x7, and three 55 mm AEO ducted fans. The servo input signals are of standard 50 Hz, while the ESCs of the five motors receive inputs with a rate of up to 490 Hz. In this section, the results of real-time experiments are presented" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000614_robot.2004.1307460-FigureI-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000614_robot.2004.1307460-FigureI-1.png", "caption": "Fig. I . 2 - W F planar arm and figured task", "texts": [], "surrounding_texts": [ "time varying adaptation deadzone such that the parameter is updated only when\nk= 1 k = l\n+ L ( A f + 6,1u(t)l)2. (32) K 111. n - T H O R D E R slso S Y S T E M S\nThis section briefly presents the generalization to higher order SISO systems. Consider the n-th order SISO system of the form\nX I = 2 2\n5,-, = X\" i\" = f(x) +g(x)u\nwhere x = 51 , x = [ X I , ..., xn I T E $3\". and U E $2. 2 ) Control Law: Consider a control law\n1 U = - (-f(x),+ zpl) - Ke) (33)\na x ) .where denotes the n-th &me derivative of x, K = [KI , K z , . . . , K,] is the feedback gain vector with K i > 0 and s ~ + C ~ = ~ K&' = 0 having roots in the left half plane, and e = [ e, d, . . . , e(\") I T is the tracking error vector with e = x - xd. When the perfect approximation is assumed such as A, = A, = 0 and \u20ac f , k = \u20ac g , k = 0, the tracking error dynamics can be expressed in the controllable canonical form of the state space representation as:\ne = A e + b ( - f f+(g-G)u)\n= Ae + b ( - @ T ( ~ j +e9u)) (34)\nwhere\nNote that A is Hurwitz. Define the sliding surface\nel = c e (36)\nwhere c = [AI, . . . , A n ] is chosen such that (A, b, c) is minimal (controllable and observable) and H ( s ) = c(s1 - A)-'b is strictly positive real. This filtered tracking error will be used in the tracking error-based parameter update [2] and the strictly positive real assumption will be necessary in the Lyapunov stability analysis.\nThe composite parameter adaptation law with projection for the n-th order SISO system becomes:\na ) Composite Parameter Update Law with Projection:\n6 k = Proj { P k % k ( C k e l . + W k e p k ) } (37)\ni ) k = X P k - W k P k % k ? T P k (38)\nNote that the complete proof with imperfect approximation requires that (37) and (38) be modified by appropriate deadzones to prevent parameter adaptation when the prediction and tracking errors become sufficiently small.\n3) Stubilitj Analysis: Consider the Lyapunov function\nl K 1 (39) V = -eTSe 2 + - 2 k = l QkprGjk.\nBy the Lefschetz-Khan-Yakubvich lemma [9], with the positive real assumption of ( A , b , c ) , there exist real symmetric positive definite matrices S and L, a real vector q. and P > O\nATS+SA = -qqT-pL (40) Sb = cT (41)\nSimilar to the derivation in Section II-D, the time derivative of V can be calculated as\nwhere 1 2 a = -Xmi.(qqT + pL) > 0.\nThis implies asymptotic convergence of the tracking error and the approximation error. With imperfect approximation, we need to treat the magnitude of U and the terms associated with the function approximation error for introducing a deadzone.\nIV. MIMO SYSTEMS\nThis section sketches a further generalization of the proposed learning adaptive controller to a class of MIMO systems:\nrn\nX = f(x) + G(x)u := f(x) + zg,(x)u, (43) *=1\nz = h(x) (44) where x E R\",z E W,u E Rm, f : Rn 4 R\",g, : 8\" - R\",G: R\"-WnX\", and h: R\" + Rp\nOur ideas presented in this paper can be generalized if the system is square (having as many inputs as outputs such that m = p), minimum phase, can be transformed into a decoupled linear system via static state feedback [lo], [ I l l , and the Lie derivative of the output function, h(x) is linearly parameterizable. Under similar conditions, OUT results also extend for m 2 p.\nV. EMPIRICAL EVALUATIONS For an emprical evaluation of the proposed learning adaptive controller, we chose an application for a planar 2-DOF arm as an example of a nonlinear MIMO system. The plant dynamics take the form of of a standard two-link planar manipulator:\nM(q)q + V(q, 4 + G(q) + DO= 7 (45)\nwhere q = [ el, O2 IT, T = [ sl, sZ I T is the torque input, M(q) is the inertia matrix, V(q, q) is the Colioridcentrifugal", "vector, G(q) is the gravity vector, and D denotes the viscous friction coefficient matrix.\nThe task is to draw a periodic figure-8 pattern in Cartesian space at a 1Hz base frequency under the influence of gravity as depicted in Figure 1. Both links of the arm are 1 meter long, and the figure-8 had a height of 0.8 meters and width of 0.5 meters. Due to these large dimensions and the double frequency in the horizonal direction that is needed to draw the figure-8, the nonlinea~ities of the dynamics of the 2-DOF arm are significantly noticable. Desired trajectories in joint space were computed from standard analytical inverse kinematics formulae.\nThe proposed learning adaptive control is implemented as follows: First, we rewrite the dynamics of the robot arm (45) in the forward model representation\n4 = f (e4) + G(q)T (46)\nwhere\nf(q,q) = -M-'(V + G + Dq), and G(q) = M-' (47)\nand we approximate f and G using locally linear models for f, and gi, where\nEssentially, this formulation requires two independent function approximators for\n4 = f1(9>4 +g11(sh +g12(4n 02 = f2h4) +921(q)TI + 9 2 2 ( 9 ) h\n(49) (50)\nwith input vector i k for each local model as %k = [ x k , Xk 71, xZT2 I T , and %k = - ck, (5 1) Then, we design a feedback linearizing controller with the estimates f and G [:I T -T\nwhere\ne = q - q d , and K, and Kd are PD gain matrices. Note that, in practice, to ensure the numerical stability of the matrix inversion associated with the feedback linearization (in this case G-I), we employ ridge regression [7] for matrix inversion. In this paper, we are concerned with the forward model representation of the plant dynamics for general control applications. However, in our future work, for the particular application of rigid body dynamics including robot arms, we will address an inverse model formulation which does not require matrix inversion and would be made suitable for such plant dynamics.\nFigure 2 illustrates the performance of the proposed learning adaptive controller in comparison to a low gain PD controller. In Figure 2, Xdes is the desired end-effector trajectory, and X p o is the result of PD control. X,,,,,,,. denotes tracking with the proposed controller after 60 seconds of learning on a slowly drifting figure-8 pattern around the desired trajectory. For our learning adaptive controller, learning staned from scratch with no local models, and new local models were added as necessary if a training point did not activate any local model by more than a threshold. The distance metric of the local model was also learned on-line. While the low gain PD controller has very large tracking errors that lead to a strongly distorted figure 8 pattern, the proposed controller achieved almost perfect tracking results after 60 seconds of learning. As a measure of the tracking performance, the La norm of the tracking errors in joint space is 4 . 8 0 ~ lo-' rad for the low gain PD controller and 1.37 x rad for the ro sed controller, which is defined by L2[e(t)] = 4- where T = t f - to. These simulation results demonstrate rapid learning of the system dynamics and convergence of the tracking error of the proposed learning adaptive controller.\n2651", "VI. CONCLUSION\nIn this paper, we presented a comprehensive development of a provably stable learning adaptive controller that covers a large class of nonlinear systems. The method employs a locally weighted learning framework [7], in which unknown functions are approximated by piecewise linear models and the learning parameters are updated by both tracking and function approximation errors. We extended our results from previous work to a class of first-order SISO systems with an unknown control input gain term g ( x ) . A parameter projection method [8] was included to avoid singularities during parameter adaptation. We also briefly presented a further generalization to a class of higher-order SISO systems and MlMO systems. Stability analyses and numerical simulations were provided to demonstrate the effectiveness of the proposed algorithm.\nFuture work will address developments of theoretically sound learning and control algorithms toward real-time highdimensional system control including humanoid robots. Our current locally weighted learning algorithm for function approximation with piecewise linear models (RFWR) will become computationally very expensive for learning in highdimensional input spaces. As a replacement, we consider using an advanced statistical learning algorithm, locally weighted projection regression (LWF\u2018R), proposed in 1121 which achieves low computational complexity and efficient learning in high-dimensional spaces. In this paper, from a control theoretic point of view, we considered a general forward dynamics representation. However, for the particular application to rigid body dynamics, we are more interested in an inverse model representation. We will adapt our framework to this special case of nonlinear. system and also compare it with biologically inspired internal model learning such as feedback error learning [131.\nACKNOWLEDGMENT\nWe would like to thank Chris Atkeson for his helpful comments. This work was supported in part by National Science Foundation grants ECS-0325383, IIS-03 12802, IIS-0082995, ECS-0326095, ANI-0224419, a NASA grant AC#98-516, an AFOSR grant-on Intelligent Control, the ERATO Kawato Dynamic Brain Project funded by the Japan Science and Technology Agency, Communications Research Laboratory (CRL), and the ATR Computational Neuroscience Laboratories.\nREFERENCES \u2019\n[ I ] F.-C. Chen and H. K. Khalil. \u201cAdaptive control of a class of nonlinear discrete-time systems using neural networks,\u201d IEEE Tranrocrions on Aulomoic Conlml, vol. 40, no. 5, pp. 791-801, May 1995. 121 1. Y. Choi and J. A. Farrell. \u201cNonlinear adaptive connol using nelworkks of piecewise linear approrimatiors,\u201d IEEE Tmonsocrions on Neural\n131 A. U. Levin and K. S . Narendra. \u201cControl of nonlinear dynamical systems using neural networks: Conmllability and stabilization:\u2019 IEEE Tmnsocrions on Neuml Networks. vol. 4, no. 2, pp. 192-206, Mar. 1993. [4] R. Sanner and J.-J. E. Slotine, \u201cGaussian networks for direct adaptive control,\u201d IEEE Tonsonions on Neural Networks, vol. 3, no. 6. pp. 837- 863, Nov. 1992. IS1 S . Ssrhagiri and H. K. Khalil, \u201cOutput feedback control of nonlinear systems using RBF neural networks,\u201d IEEE Tmmnsoerions on Neural NcrworLs, vol. 11, no. 1. pp. 69-79. Jan. 2000.\n.\u2019 Nenvork. vol. 11, no. 2. pp. 390-401, Mar. 2wO.\n161 J. Nakanishi. J. A. Farrell, and S. Schaal. \u201cA locally weighted learning composite adaptive controller with S ~ C ~ U R adaptation,\u201d in IEEWRSJ Inremotional Conference on lnrslligenr Robots and System. 2002, pp. 882-889. 171 S. S c h d and C. G. Atkeson. \u201cConstructive incremental learning from only local information,\u2019\u2019 Neural Compuralion, vol. IO, no. 8, pp. 2047- 2084. 1998. 181 M. KristiC. I. Kanellakopoulos, and P. Kokotovif. Nonlineor and Adoprive Contml Design. John Wiley & Sons, Inc., 1995. 19) G. Tao and P. A. loannou, \u201cNecessary and sufficient conditions for stictly positive 4 matrices:\u2019 IEE Pmceedin.qs G, Circuirs. Devices ond Sysrems, vol. 137, no. 5. pp. 360-366, 1990. 110) S. Sasvy and M. Bodson, Adaplive Conrml: Slabili@ Convergence, and Robusmess Prentice Hall. 1989. [I I ] S. Sastry and A. Isidori, \u201cAdaptive control oflinearimble systems,\u201d IEEE Tronsonions on Automatic Conrml. vol. 34, no. 11, pp. 1123-1131. 1989. 1121 S . Vijayakumar and S. Schaal. \u201cLocally weighted projection regression: An Mn) algorithm for incremental real time laming in high dimensional space:\u2019 in Inremorional Conference in Machine Leoming (ICMLJ. 2MX). pp. 1079-3086. I131 H. Gomi and M. Kawato, \u201dNeural network control for a closed-Imp system using feedback-emr-learning,\u201d Neuml Networkr. vol. 6. pp. 933- 946. 1993.\nAPPENDIX\nCONSTRAINED OPTIMIZATION USING LAGRANGE MULTIPLIER METHOD\nConsider the following constrained optimization problem:\n(54)\nwhere x E L\u201d, b E Ln is the parameter vector of the linear model, D is a positive definite distance metric, and f i is a constant which is representative of the local model boundary. Note that the h e a r model xTb does not include the bias for notational simplicity.\nminimize f(x) = xTb subject to the constraint g(x) = xTDx - fi = 0.\nDefine an objective function\nJ = xTb + X(xTDx - f i ) (55)\nwhere 1 is the Lagrange multiplier. Then, at the stationary points of J , we have\nand solving this equation for x yields\nSubstituting (57) into (54), we have\n1 x = *-..miFG. 2Ji;\nThus, we obtain\n(57)\n(59)\nThe value of f(x) at (59) is given by\nXTb = *Jm, (60)\nThus, the minimum of xTb subject to the constraint (54) is -J-~atx=-*D-\u2018b." ] }, { "image_filename": "designv11_61_0002096_robio.2007.4522345-Figure9-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002096_robio.2007.4522345-Figure9-1.png", "caption": "Fig. 9. Trajectories of the J-type and other types", "texts": [ " In this experiment, we realized that the running courses are different among the J-type and the other methods; the J-type can run on smooth curves in a manner similar to human motion and can stop at a desired position, while the others alternately combine straight runs and rotation and are difficult to stop at a desired position. In the latter cases, it is important to detect the spot where the wheelchair should stop and change directions. For example, when running a pathway such as that shown in Fig. 9, an operator needs to stop at a position that is further than the stopping point of the J-type in order to run smoothly with fewer changes in directions. To solve this difficulty, we have proposed a system for detecting the position and direction using RF tags. This system comprises an array of RF tags on a floor and a sensing device attached to the bottom of the wheelchair. We then determined the optimal position for stopping the wheelchair. First, consider the stopping points A, B, and C, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000740_robot.2004.1308050-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000740_robot.2004.1308050-Figure3-1.png", "caption": "Figure 3: Non-holonomic vehicle.", "texts": [], "surrounding_texts": [ ":T = 3 \"ab) :T (1)\nwhere ',T and \"J are the given transformation matrices of i t > w.1.t. and of \"base kame\" w.r.t. , respectively;\nwhile Y(q) is instead the transformation matrix of w.r.t. <@, it also real-time evaluated via the knowledge of the current joint position vector q.\nThen, by letting\neT = [pT,d'] (2)\nbe the collection of the (projected on world frame <@) misalignment error vector e and distance error vector d, of h m e if> w.r.t ; and moreover by also letting (with a small abuse of notation due to the use of the ''.\" upper sign)\nbe the collections of (projected on ) angular and linear velocities, of frame and y> respectively (with the former it also given as a real time reference input), we can immediately see, hy considering the (tentative) Lyapunov h c t i o n\n(4) y = - e re 1 =) 3 = - e T ( x - f * )\n2 that an assignment to the joint velocity vector (if any) 4 capable of satisfying, at all tbe times, the condition\ni = f = y e + i ' = ~4 ; y > o (5 ) (whereJ(q) is the Jacobian matrix of w.r.t. with output velocities projected on ) would consequently drive asymptotically.\nIn order to fulfill (5), a kinematic control architecture is here proposed, where a task-space velocity reference generator module (the dotted square in fig. 2) is first used for computing x' fiom the second of (5), via the knowledge of transformation matrices Tg and :T, and the use of the so called \"versors lemma\" (see for instance [l]) for evaluating the current error vector e. Then a kinematic control module (the dashed square in fig. 2, receiving x' as reference input) is in turn used for finding out, via kinematic inversion, the\nproper joint velocity reference vector 4' maximally satisfying (see later for further details) the last condition in (5). In parallel, the already mentioned Direct-Geometry module is used for evaluating, other than the transformation matrix :T, also the associated non singular matrix S = d x 6 , representing the instantaneous rigid body velocity transformation matrix fiom frame to W e (input velocities projected on , output velocities projected on ) which is explicitly required for real-time computing the Jacobian matrix J in ( 5 ) via the product\nJ = S 7 (6) beingJ(q) the (so called \"basic\") Jacobian matrix of \nw.r.t. <@, with output velocities projected on ; it also real-time computed on the basis of the current joint position knowledge.\nFinally, the presence of an inner joint velocity dynamic control module ,(not shown in figure and possibly implemented on dedicated hardware) is also assumed, capable of precisely tracking the real-time generated joint velocity reference 4 .\nAs it concerns the maximal fullfilment of the last of ( 5 ) via kinematic inversion, it must be formerly recalled how this last should actually be performed via the use (at least) of a regularized pseudoinversion of the Jacobian mat& J (see [15]); thus preventing any component of the joint velocity references to grow toward infinity in the vicinity of any Jacobian singular configurations that could be encounterd during the motion.\nDespite such an undoubtely and necessary benefit, some related secondary drawbacks are however implied by the use of the regularization itself; simply consisting in motion perturbations occurring in the neighbourhood of singularities, or even possibilities of stuck in correspondence of some of such configurations (with related need of possibly complex manoeuvring for departing).\nIn order to also reduce such occurrences, a secondary task, aiming to maintain the arm far f iom singularities while however accomplishing the (now primary) task (4), can actually be introduced by just exploiting the assumed redundancy of the arm.", "To this end, by referring to the so called \"manipulability measure\" (MM) [15,23], that is the scalar quantity\n,U = det (JJ') t 0 (7) and considering its time derivative; that is (see [I41 for an easy technique of evaluation) :\nWe can immediately see (by assuming a starting posture far from singularities) that a vector q capable of maximally attempting to satisfy also the condition\n& A p = p q ; 1>0 (9)\nwould sensibly reduces the risk of any near-to-singularity occurrences during motion.\nThus, by combining tasks (5) and (9) in the mentioned . order, via the well known priority-based task composition\ntechnique described in [15], the following espression for the corresponding joint velocity .reference vector 1 is consequently obtained\n;= J'x' + h ( z - kx' ) (10) with\nand k + p J y (11)\nh t [ p ( l - J ' J ) ] # (12)\nThe adoption of (IO), though being effective for cases when the goal frame is strictly located inside the \"dexterous reachable workspace\" (DRIV) [23], cannot however prevent singularity occurrences whenever the goal h m e is instead located outside DRW. In this latter case, in fact, due to the assumed priority of the goal reaching task, the arm will totally stretch toward the unattaineable goal, thus unavoidably reaching the associated singular configuration.\nThen, in order to also prevent sucb possibility of singularity occurrence, a nice solution has been recently proposed (tirst in [I41 and then [24,25]) simply consisting in smoothly changing the priority of the two tasks whenever, during motion, MM is recognised tending to become close @om above) to an a-priori established \"safety\".threshold po. This is done by simply and automatically adding, to the external . Cartesian velocity reference $, an internally generated Cartesian signal i , accomplishing the function of compensating for those component directions of $ that, otherwise, would further reduce the value of MM. Such a needed additional internal signal i (see [24,25] for specific details) results to be of the form\ni = a k # ( P - k q (13)\nwhere is a continuous bell-shaped, scalar 'function of p (unitary for #fi and tending to zero within a finite support,\nfor p >& ), used for automatically driving the task priority inversion, whenever needed, in a smooth way.\nAccordingly with the addition of such internal reference termi , the resulting output Cartesian velocity now attains the general form X =*+ i = &'z+ (I - &*k); (14) not any more coinciding with the external reference%, whenever in the vicinity of the UM threshold ,U@ In such Occurences the opposite of i obviously assumes also the meaning of a resulting output Cartesian velocity error b ; i.e.\ni ; - X = - i =-&'E+ a(k'k)$ (15)\nIn. VELC\u20acIl\" CONTROL OF A NON-HOLONOMIC VEHICLE A 3D non-holonomic vehicle is now considered as a standalone system. Such system is assumed to be non-holonomic in the sense that it allows a linear velocity vector x only directed along the principal vehicle axis, and an angular velocity vector - Ronly lying on the plane passing through a known point of such principal axis and orthogonal to it. A fixed body frame is then attached in correspondence of such point, with one of the axis just directed as the principal one; while a \"vehicle end-effector M e \" Q> is it also assumed to be rigidly attached to the vehicle, as depicted in fig.4.\nMoreover, in order to maximize the similarity with the previously considered arm system, a \"vehicle base h e \" is also considered, now supposed to be'rigidly attached to . Finally, for such system, only absolute velocity tasks referred to <(> will be considered.\nIn this framework first let\nA T = [ w * , u ]\nbe the three dimensional vector resulting from the collection of the two non null components w of Qand the sole non null component U of 1: whenever both Q and v are projected on M e . Moreover also let be the absolute velocity vector (angular and linear parts both projected on world frame", ") induced by 9 on frame v> and, consequently, also keep into account the linear relationship relating the two; that is\nX = J 8 (17)\nwith the Jacobian matrix J E 31':' now taking on the form\nJ = SHQ (18)\nwhere the nonsingular matrix S E 316x6 represents the rigid body velocity transformation from frame to , output velocities projected on world frame ) to be evaluated as for the arm case. Similarly, H E x''~ is a constant nonsingular matrix espressing the rigid body velocity transformation from to (input velocities projected on , output velocities projected on world frame ); while Q E 31'\" is simply a full rank selection matrix suitably composed by 0 and 1 elements.\nThen, by keeping into account the defective structure of the above Jacobian matrix (directly induced by the nonholonomic characteristics of the vehicle) it immediately follows that any tool fiame absolute velocity task, represented by a corresponding externally assigned velocity reference signal x' , can always be solved in a least square sense only; that is only in the form\nbeing f the left pseudoinverse of matrix J; thus generally leading to the following output Cartesian velocity error B\ni I . i - X = 1 - I JJ +j+ (20)\nAlso for this case, the corresponding control scheme results in a form which is structurally identical to the kinematic part of the one seen in the previous section (see fig 5); with the Kinematic-Inversion block now represented by the implementation of (19), and the Direct-Geometry block still performing the same real-time evaluations as for the ann case; the sole difference being now that 'Jis computed via integration of odometric data (andlor any other more robust telemetric or visual means, possibly fused with the odometric ones).\nw. ARM-VEHICLE COMPOSETON AND SELFCOORDINATION TECHNIQUE\nThe case of a redundant arm mounted on a 3D nonholonomic vehicle (see fig.5) is now considered. The overall system is assumed to be assembled in such a way that the vehicle end-effector frame coincides with the arm base frame; i.e. < e W O p , where subscripts \"a\" and .'fl will be hereafter used (with an obvious meaning), whenever useful for avoiding possible ambiguities. The assigned (now common) task still consist in having the arm tool frame .\nI2 =\nt I--\n#,&+\nF i g w 5: Arm supported by a vehicle.\nTo better describe the benefits in terms of modularity and ease composability offered by the self-coordination technique proposed in the following, it is first usefu'starting with a description of the information that must be exchangeb by two subsystems. To this end, a graphical schematization of the kinematic control module of both the subsystems, is depicted in fig. 6, where (differing from 'fig. 2 and fig. -4) only the input/output signals have been represented.\nSuch symbolic representation actually allows to make clearer the description of the proposed composition technique. This is done by making explicit reference to the scheme in fig.7, evidencing all the signals, and relative connections, required for coordination. Such connections have been divided into two separate sets, represented by dashed and solid lines, respectively.\nThe dashed lines represent the signals used by the subsystems for exchanging geometric data (Transformation matrices) just reflecting the considered subsystems assembly; without being related to any specific motion task to be accomplished. More specifically:" ] }, { "image_filename": "designv11_61_0003445_9781118361146.ch7-Figure7.31-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003445_9781118361146.ch7-Figure7.31-1.png", "caption": "Figure 7.31 Diagrams to show how a rotating magnetic field is produced within an induction motor", "texts": [ "31. Three coils are wound right around the outer part of the motor, known as the stator, as shown in the top of Figure 7.30. The rotor usually consists of copper or aluminium rods, all electrically linked (short-circuited) at the end, forming a kind of cage, as also shown in Figure 7.30. Although shown hollow, the interior of this cage rotor will usually be filled with laminated iron. The three windings are arranged so that a positive current produces a magnetic field in the direction shown in Figure 7.31. If these three coils are fed with a three-phase alternating current, as in Figure 7.23, the resultant magnetic field rotates anti-clockwise, as shown at the bottom of Figure 7.31. This rotating field passes through the conductors on the rotor, generating an electric current. A force is produced on these conductors carrying an electric current, which turns the rotor. It tends to \u2018chase\u2019 the rotating magnetic field. If the rotor were to go at the same speed as the magnetic field, there would be no relative velocity between the rotating field and the conductors, and so no induced current and no torque. The result is that the torque/speed graph for an induction motor has the characteristic shape shown in Figure 7" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001823_978-1-4020-6405-0_16-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001823_978-1-4020-6405-0_16-Figure1-1.png", "caption": "Fig. 1. \u201cFlat-on-cylinder\u201d contact geometry (tribometer test configuration).", "texts": [], "surrounding_texts": [ "In the present paper, the dependence of the dynamic friction coefficient value with the contact pressure in rubber-aluminium cylindrical contact geometries is analysed in detail. In this \u201cflat-on-cylinder\u201d configuration, which is also selected for tribometer testing with reciprocating motion, contact pressure distribution is namely not uniform along the whole area of contact if it is compared to a classical \u201cflat-on-flat\u201d tribometer test configuration. Since contact pressure plays a substantial role in rubber friction as dependency of the friction coefficient, to obtain values of the friction coefficient calculated directly as the ratio between the tangential and vertical forces for a given average value of the contact pressure along the area of contact, without detailed postprocessing of the experimental results, may lead to severe errors. As it will be shown later, contact pressure distribution changes significantly along the area of contact, showing high scatter with respect to the mentioned average value. Due to the previous, the friction coefficient value vs. contact pressure has to be calculated indirectly from friction force results obtained in tribotesting by means of a mathematical method. 258 The Influence of Contact Pressure on the Dynamic Friction Coefficient In addition, the apparent area of contact is also not constant when the vertical load increases. Thus, the robustness of the method has to be improved by comparison of experimental measurements of the apparent area of contact with FEM results of the tribotesting, combining them with the adjustment of the rubber material model. As it will be explained in this paper, the method consists on combining FEM simulations of the tribotesting to obtain contact pressure distributions along the cylindrical area of contact for different vertical loadings, and then on developing a mathematical procedure for obtaining a final analytical expression for the dynamic friction coefficient as dependent with the contact pressure. The main objectives of the present paper are the following: \u2022 to analyse the dependency of the friction coefficient with contact pressure on rubber-metal sliding contacts; \u2022 to compare the contact pressure distribution in \u201cflat-on-flat\u201d as well as in \u201cflaton-cylinder\u201d configurations. To study the friction coefficient dependency with contact pressure in cylindrical metallic counterparts. The main expected result is \u2022 to develop \u00b5 = \u00b5(p) experimental laws from tribometer tests results on \u2018flat-oncylinder\u201d configurations. According to the previous objectives, the next approach has been followed: \u2022 planning and execution of tribometer testing on \u201cflat-on-cylinder\u201d configurations using rubber as sliding sample; \u2022 FE analysis of area of contact and contact pressure distribution vs. applied vertical load; \u2022 mathematical procedure for obtaining real \u00b5 = \u00b5(p) experimental laws." ] }, { "image_filename": "designv11_61_0000278_1-84628-179-2_7-Figure7.2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000278_1-84628-179-2_7-Figure7.2-1.png", "caption": "Fig. 7.2. Boeing Heliwing vehicle.", "texts": [ " With the pilot replaced by modern control systems it should be possible to realize the original promise of the tail-sitter configuration. The tail-sitter aircraft considered in this chapter differs substantially from its earlier counterparts and is most similar in configuration to the Boeing Heliwing vehicle of the early 1990s. This vehicle had a 1450 lb maximum take-off weight (MTOW) with a 200 lb payload, 5 hour endurance and 180 kts maximum speed and used twin rotors powered by a single 240 SHP turbine engine [113]. A picture of the Heliwing is shown in Figure 7.2. Although conflicts over the last 20 years have demonstrated the importance of military UAV systems in facilitating real-time intelligence gathering, it is clear that most current systems still do not possess the operational flexibility that is desired by force commanders. One of the reasons for this is that most UAVs have adopted relatively conventional aircraft configurations. This leads directly to operational limitations because it either necessitates take-off and landing from large fixed runways, or the use of specialized launch and recovery methods such catapults, rockets, nets, parachutes and airbags" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000414_physreve.69.041707-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000414_physreve.69.041707-Figure3-1.png", "caption": "FIG. 3. Definition and geometry of the contact line vector j for anisotropic contact lines. Schematic of a segment of the contact line, the contact line vector j and its components, the principal anisotropy frame (po ,bo), and the normal plane ~NP! to the unit tangent t. The projection of the line vector j on the NP defines the principal anisotropy frame (po ,bo).", "texts": [ " For anisotropic contact lines, there is a principal anisotropy coordinate frame (po ,bo), and the rotation of the unit line tangent t around bo produces the maximum increase in surface energy. The principal anisotropy frame (po ,bo) on the plane normal to t is selected by the main anisotropic axes of the surface. Anisotropic lines can change line energy by dilation and by rotation. Figure 2 shows an element of length Lo5Lo\u2022to and line unit tangent 04170 to of the contact line that undergoes an expansion to L and rotation to t. Since the line energy density is a function of t, x~t!, the total line energy X of the contact line can be increased by expansion and by rotation of t. Figure 3 shows a segment of the contact line, the components of the contact line vector j, their magnitudes, the principal anisotropy frame (po ,bo), and the normal plane ~NP! to the unit tangent t. The principal anisotropy frame (po ,bo) defines the NP. Figure 4 shows a schematic of the capillarity vectors j and 2j, the normal 2j' and tangential 2ji components of 2j, acting on a point of the contact line, the NP, and the unit tangent vector t. The vector 2j represents the line force acting on the line vector L5Lt tending to rotate (2j') and 7-3 shrink (2ji) the contact line" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002992_cso.2009.244-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002992_cso.2009.244-Figure2-1.png", "caption": "Figure 2: The relationship of AUV ,obstacle and sea flow", "texts": [ " So moving-to-goal behavior can be represent by compute the direction vector, as shown in equation (1): \u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 \u2212 \u2212 \u2212+\u2212 =\u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 = cg cg cgcgMTG MTG MTG yy xx yyxxy x V 22 )()( 1 (1) Here MTGV is the moving-to-goal vector, ),( gg yx is target position; ),( cc yx is the current position of AUV. The vector of obstacle avoiding behavior can be computed by equation (2): \u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 \u23a2 \u23a3 \u23a1 \u00d7\u23a5 \u23a6 \u23a4 \u0394 \u0394\u2212 \u0394 \u0394 =\u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 = d d OA OA OA y x y x V )cos( )sin( )sin( )cos( \u03b8 \u03b8 \u03b8 \u03b8 (2) Here MTGV is collision avoiding vector, ),( dd yx is the velocity vector of AUV,which shows the current direction of AUV, \u03b8\u0394 is the angle that the AUV should rotate. The relationship of AUV, obstacle and sea flow can be demonstrated as figure 2. P is the position of AUV, Vc is AUV\u2019s current velocity, Vf is the velocity brought by sea flow, and is the angle between Vc and Vr. So if AUV want to keep the direction of Vc, it should have the velocity of Vr to resist the sea flow. It\u2019s easy to think the Vr is subtraction of Vc with Vf, as shown in equation (3)or (4): fcR VVV \u2212= (3) Or \u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 \u23a2 \u23a3 \u23a1 \u2212\u23a5 \u23a6 \u23a4 =\u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 = f f d d R R R y x y x y x V (4) Here ),( dd yx is the vector form of Vc, and ),( ff yx is the vector form of Vf. But how the sea flow influence the AUV is a very complex problem, it is far from what we list here" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000414_physreve.69.041707-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000414_physreve.69.041707-Figure2-1.png", "caption": "FIG. 2. Schematic of a contact line element of length Lo 5Lo\u2022to and line unit tangent to that undergoes an expansion to L and a rotation to t. Since the line energy density x~t! is a function of t, the total line energy X of the contact line can be increased by the stretching of its length Lo and by rotation of its unit tangent to .", "texts": [ " The selected normal component of the line tension vector j is the one that maximizes the increase of surface energy with rotation, and hence j'5j\u2022Ic5S dx du D max po , ~13! where po is the unit normal vector along which dx/du has the maximum rate of increase. For anisotropic contact lines, there is a principal anisotropy coordinate frame (po ,bo), and the rotation of the unit line tangent t around bo produces the maximum increase in surface energy. The principal anisotropy frame (po ,bo) on the plane normal to t is selected by the main anisotropic axes of the surface. Anisotropic lines can change line energy by dilation and by rotation. Figure 2 shows an element of length Lo5Lo\u2022to and line unit tangent 04170 to of the contact line that undergoes an expansion to L and rotation to t. Since the line energy density is a function of t, x~t!, the total line energy X of the contact line can be increased by expansion and by rotation of t. Figure 3 shows a segment of the contact line, the components of the contact line vector j, their magnitudes, the principal anisotropy frame (po ,bo), and the normal plane ~NP! to the unit tangent t. The principal anisotropy frame (po ,bo) defines the NP" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002291_bfb0119397-Figure10-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002291_bfb0119397-Figure10-1.png", "caption": "Figure 10. A modified operator joystick.", "texts": [ " The slew drive has a nested loop structure with feedback of axis rotation angle and also bucket swing angle rate. The result of a swing motion with and without swing stabilization are compared in Figure 9. Due to inter-cycle variability it is more difficult to compare actual performance with and without swing stabilization. 4. U s e r i n t e r f a c e The automation system 'drives' the dragline by physically moving the control joysticks and pedals; like the cruise control in a car. To achieve this we fitted brushless servo motors to each of the control devices, see Figure 10. Servoing these controls facilitates the smooth transfer of system set points in the transitions between automatic and manual control modes. Sensing on the drag and hoist joystick servo motors allows the automation system to sense if the operator is opposing the motion of the joysticks - - in which case it smoothly transfers control back to the operator. Thus the servoed operator controls impose on the control computer the same safety interlocks and overrides that apply to an operator. The servoed joystick could also be used to provide a kinesthetic interface to the operator" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000089_rspa.2002.1105-Figure9-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000089_rspa.2002.1105-Figure9-1.png", "caption": "Figure 9. The Liverpool splice.", "texts": [ " The procedure first assumes a direction cosine or an average direction, which yields the maximum and minimum radial position (rmax and rmin) of the subropes; this then enables the calculation of the constant C and this gives an estimate of ds/dz, an improved value for the direction cosine. The estimation of the subrope radial positions and the direction cosines depends upon the type of assembly. The structure of the component subropes arising from assumptions of the softness and hardness of the subropes will result in different assembly algorithms. In each case the maximum and minimum radial points lead to the mean radius rm and the radial travel \u2206r. Proc. R. Soc. Lond. A (2003) 1654 C. M. Leech The geometry of the Liverpool splice is shown in figure 9; the rope is at the bottom and the splice is evolved by progressing up the figure. The shaded region at the bottom is an R subrope, and this is twisted in the clockwise (Z) direction and progressing into the splice. As it encounters the splice S subropes, it is twisted against its own S subrope in the S direction to form a two-component twisted \u2018strand\u2019, so that the twist of the R\u2013S subrope assembly is in the opposite direction to the direction that the \u2018strand\u2019 twists about the rope/splice core" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003445_9781118361146.ch7-Figure7.29-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003445_9781118361146.ch7-Figure7.29-1.png", "caption": "Figure 7.29 The rotor and stator from an SRM. (Photograph reproduced by kind permission of SR Drives Ltd.)", "texts": [ " Correct turning on and off of the currents in each coil clearly needs good information about the position of the rotor. This is usually provided by sensors, but modern control systems can do without these. The position of the rotor is inferred from the voltage and current patterns in the coils. This clearly requires some very rapid and complex analysis of the voltage and current waveforms, and is achieved using a special type of microprocessor called a digital signal processor.5 An example of a rotor and stator from an SRM is shown in Figure 7.29. In this example the rotor has eight salient poles. The stator of an SRM is similar to that in both the induction and BLDC motor. The control electronics are also similar \u2013 a microprocessor and some electronic switches, along the lines of Figure 7.21. However, the rotor is significantly simpler, and so cheaper and more rugged. Also, when using a core of high magnetic permeability the torque that can be produced within a given volume exceeds that produced in induction motors (magnetic action on current) and BLDC motors (magnetic action on permanent magnets) (Kenjo, 1991, p" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002236_2007-01-3740-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002236_2007-01-3740-Figure1-1.png", "caption": "Figure 1: Full toroidal Variator schematic", "texts": [ " Full toroidal traction drive transmissions are normally associated with high torque, longitudinal RWD / 4x4 applications due to previous demonstrations of a \u2018geared neutral\u2019 Infinitely Variable Transmissions (IVT) arrangement in a fleet of V8 powered Sport Utility Vehicles (SUVs). At the beginning of 2007 the full toroidal technology achieved commercial launch in the USA in the Outdoor Power Equipment (OPE) market for domestic and commercial ride-on lawnmowers and garden tractors. The experience generated from both the OPE and Automotive market segments have been married together to generate a range of new, cost effective, front wheel drive (FWD) transmission concepts. The heart of the Torotrak IVT is the full toroidal traction drive Variator. The schematic in Figure 1 explains the operating principle of the Variator. The engine drives the input discs (1) and power is transmitted via the rollers (2) to the output discs (3). When the rotational velocities of the input and / or output discs change, the rollers automatically alter their inclination in order to adjust to the new operating conditions (4). Power transmission is achieved by traction, i.e. by shearing an extremely thin, elasto-hydrodynamic fluid film (traction fluid [1]) and not through metal-to-metal friction" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002412_icems.2009.5382984-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002412_icems.2009.5382984-Figure6-1.png", "caption": "Fig. 6. Nodal force", "texts": [ " The conditions for each analysis are explained. A PWM carrier frequency has little effect on the electromagnetic vibration, because the largest resonance frequency was found around 3000Hz, the PWM carrier frequency is higher than it. Therefore the sine wave current is added to the magnetic analysis. The front cover of the motor is tightly attached to the motor mount, the surface is selected as the constraint in the structural analysis. The surface is shown in Fig. 5. The electromagnetic force distribution is shown in Fig. 6. This is concentrated on the surface of the stator and is facing to the air-gap. The radial component of the electromagnetic force on the tip of the stator teeth greatly effects on the electromagnetic vibrations. The mechanical characteristics such as young\u2019s modules and mass density are identified by comparing the measured natural frequency for each component with the analysis results. The identified mechanical characteristics are shown in Table II. The laminated steel of both rotor and stator are modeled as the anisotropic material" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001173_20060912-3-de-2911.00061-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001173_20060912-3-de-2911.00061-Figure7-1.png", "caption": "Fig. 7. Cartesian Robot", "texts": [ " Especially during the conceptual design phase there exists a high demand for models to describe the design concept with respect to the given requirements. The following section demonstrates the above experiences from an industrial case study. Typical applications for robots are pick and place operations, screw driving, welding etc. A robot consists of two main parts, a positioning system, which locates the functional part, sometimes called an \u201cend effector\u201d. A manipulator with prismatic first three joints is known as a cartesian manipulator (see Figure 7). For the cartesian manipulator the joint variables are the cartesian coordinates of the endeffector with respect to the base. As might be expected the kinematic description of this manipulator is the simplest of all manipulators. Cartesian manipulators are useful for table-top assembly applications and for transfer of material. We only consider the robot without the end-effector. In order to fulfill the design requirements, the movement of the three axes, the determination of the x,y,z-position of the end-effector and the control of the three drives have to be considered" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000014_memsys.2002.984233-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000014_memsys.2002.984233-Figure2-1.png", "caption": "Figure 2: 3Dperspective view of inclined microswitch design", "texts": [ " Subsequently the substrate is tilted at some desired inclination allowing the SU-8 to settle and planarize parallel to the horizon but inclined relative to the substrate surface, as shown in Figure lb. The SU-8 is then cured to form the final inclined state. Finally the trough walls are removed to complete the ramp structure, as shown in Figure IC. Although the primary application explored in this paper is the fabrication of magnetic switches, many alternative uses for continuously-varying, three-dimensional SU-8 structures can also be considered using this process. Such applications include microoptics, microfluidics, and mechanically-interlocking structures. Figure 2 shows a 3D model of a magnetically actuated switch that exploits the inclined ramp structure. The underlying incline guarantees that when a near-uniform magnetic field approximately normal to the substrate is 3-7803-7185-2/02/$10.00 02002 IEEE 176 applied, the actuation will unambiguously occur in the downward direction. In the presence of an external magnetic field, a permalloy material will develop a net magnetization M. This magnetization, interacting with the applied magnetic field will cause a net torque on the permalloy material: Tjeld= M x H = M H sin a where a is the angle between and E" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003470_iccasm.2010.5619227-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003470_iccasm.2010.5619227-Figure2-1.png", "caption": "Fig. 2 calculate the model layer", "texts": [ " B Material The material is commercially available, gas atomized stainless steel 316L powder, the particle size distribution is shown in table 1 and table 2, and the steel has the following chemical composition: 16.9Cr, 1O.SNi, 0.02C, O.SSi, 2.l3Mo, 1.3Mn, 0.060, 0.20N, O.OOSS, Fe balance (%wt). The steel powder density is 4.S glcc. C Scan algorithm This paper will put forward the helix scan strategy based on Voronoi diagram. The algorithm and its implementation will be presented. Firstly, before every layer of any model is processed, STL file model is imported into the software (Fig. 2 a), and the slice of STL format file waiting for processing is calculated according to the height of the layer (Fig. 2 b, c). When the STL model is slicing, a fast error-tolerant slicing algorithm is adopted. Based on hole-tracking technology and the thought of reducing dimensions, the algorithm makes full use of three dimensional information in STL file including the topological structure and the contour for holes, which makes almost every STL file sliced correctly without any manual operation that can guarantee gaining the correct slice information of every layer of the model. According to Held's wavefront-propagation algorithm, Voronoi diagrams of planar areas bounded by straight lines and circular arcs are calculated, and the planar areas is the aforementioned slice" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002961_icelmach.2008.4799942-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002961_icelmach.2008.4799942-Figure1-1.png", "caption": "Fig. 1. Cross section of experimental rotor.", "texts": [ "78-1-4244-1736-0/08/$25.00 \u00a92008 IEEE 1 I. INTRODUCTION This analysis focuses on a single-phase capacitor-run permanent magnet (PM) motor whose rotor consists of an aluminum rotor cage with arc-shaped interior Neodymium-Boron-Iron (NdBFe) magnets, Fig.1. The capacitor-run PM motor [1] is the single-phase version of a three-phase line-start permanent motor (LSPM) [2], [3]. This motor is suited for application in home appliances, such as refrigerator compressors [1]. However, in single-phase motors, where the auxiliary winding is supplied through a capacitor, the operation is further complicated by the imbalance between the main and auxiliary winding voltages. Because of this, the negative sequence field occurs and increases the pulsating torque, losses, vibrations, and noises compared to those of the three-phase LSPMs" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002543_20080706-5-kr-1001.00139-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002543_20080706-5-kr-1001.00139-Figure1-1.png", "caption": "Fig. 1. Experimental prototype.", "texts": [ " By joining the takeoff and landing capabilities of the helicopter with the forward flight efficiencies of fixedwing aircraft in such a simple way, the tail-sitter promises a unique blend of capabilities at lower cost than other UAV configurations. During the hover-flight regime, most single-rotor aircrafts employ the propeller to provide thrust and air slipstream to control surfaces (elevator, ailerons, rudder), as a result, the dynamics stabilization is totally based on the aerodynamic torques. In the present paper we are interested in the stabilization of a single-rotor tail-sitter CUAV (see Fig. 1) at hover flight. The proposed configuration is based on a reduced mechanics, which simplifies its maintenance and replacement. In addition, we have incorporated certain structural features to improve the stability of the aircraft with respect to external perturbations. The operation description and dynamic model are presented in section 2. The control algorithm based on separated saturation 978-3-902661-00-5/08/$20.00 \u00a9 2008 IFAC 809 10.3182/20080706-5-KR-1001.3884 functions is described in section 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003751_acc.2010.5531257-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003751_acc.2010.5531257-Figure1-1.png", "caption": "Fig. 1. Illustration of the geometric condition in Theorem 2 with an example non-convex star domain K. For \u0393 = I and any z \u2208 Rn \u00d7 \u2202K, zker \u2208 Rn \u00d7 ker(K), the projection of fp onto z \u2212 zker is always more negative (or less positive) than the projection of fn onto the same vector.", "texts": [ " Using the definition of fI\u2217 (6), this reduces to showing that x\u0303TNI\u2217 ( NT I\u2217\u0393NI\u2217 )\u22121 NT I\u2217fc \u2265 0, if 1 \u2264 |I\u2217| \u2264 q. By assumption, K is a star domain defined by the functions hi, i \u2208 I2m. It follows from Lemma 3 that x\u0303TNIsat \u2265 0. Since I\u2217 \u2282 Isat, we have x\u0303TNI\u2217 \u2265 0, which shows that x\u0303TNI\u2217 is a row vector of non-negative numbers. By Lemma 2 and our choice of I\u2217, we have( NT I\u2217\u0393NI\u2217 )\u22121 NT I\u2217fc as a column vector of non-negative numbers. The sums and products of non-negative numbers must be non-negative. The conclusion follows. The geometric condition stated in Theorem 2 is illustrated in Fig. 1 with an example non-convex star domain for the unsaturated region. If the control objective is to regulate the system state about the origin, we can set zker = 0, provided ker(K) contains the origin of Rq . When the output equation of the nominal controller depends only on the controller state, then exact controller stateoutput consistency is achieved for the GPAW compensated controller when appropriately initialized. When the nominal controller does not possess this structure, an arbitrarily close approximating controller can be constructed that has the required structure" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003632_iros.2013.6696529-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003632_iros.2013.6696529-Figure1-1.png", "caption": "Fig. 1. Schematic and free-body diagram for Omnicopter MAV", "texts": [ " Most of the famous tilt-rotor type UAVs for military use, like Bell Boeing V-22 Osprey [1], are mechanically complex since they employ a swashplate and differential rotor tilting to control pitch and yaw, respectively. Several research groups have also developed some tilt-rotor/wing type MAVs with simpler tilting and actuation designs [2], [3], [4], [5], [6]. In this paper, we introduce an original MAV configuration, named the Omnicopter MAV, composed of five rotors and three servo motors (see Fig. 1). The main characteristic of this configuration is that the attitude and translation dynamics are decoupled, such that we can design controllers for the two subsystems individually and fully control its 6 degrees of freedom (DOF) for more agility. For example, it can maintain zero roll and pitch attitude during lateral translation or arbitrarily orient the fuselage during hover. Comparing with some other over-actuated multicopters in the literature, the Omnicopter has some potential advantages. In [7] and [8], the eight-rotor UAV\u2019s control inputs are linearly related to its motor input signals", " Section IV presents some simulation results to illustrate the performance of the proposed control and allocation tech niques. The platform setup and experiments are described in Section V, and finally concluding remarks based on all the presented work are given in Section VI. In this section, we apply the Newtonian mechanics to model the Omnicopter. Let I = I x , I y , I z denote the inertial frame, and B = B x , B y , B z the aircraft body frame, with the z axes pointing downwards, as shown in Fig. 1. Using the Newton-Euler approach [12], we can derive the dynamics of a rigid body under external forces and torques applied to the rigid body 978-1-4673-6358-7/13/$31.00 \u00a92013 IEEE 1380 ml3x3 0 \u03bd\u0392 \u03c9 \u03c7 mvB _ / 0 J \u03ce + UJXJUJ ~ \u03c4 (1) (2) where \u03bd\u0392 = [\u03bd\u03c7 vy vz]T and \u03c9 = [\u03c9\u03c7 \u03c9\u03c5 \u03c9\u03b6]\u03c4 are the linear and angular velocities in the body frame, J = diag(Ixx, Iyy, Izz) is the inertial matrix and m mass, / = [fx fy fz]T and r = [TX ry \u03c4\u03b6]\u03c4 are the force and torque vectors in the body frame. We can expand (1) to obtain 6 independent equations of motion as the following m(vx \u2014 vyujz + vzujy + gs9) = fx m(vy \u2014 \u03bd\u03b6\u03c9\u03c7 + \u03bd\u03c7\u03c9\u03b6 \u2014 gc9s(f>) = fy m(vz \u2014 vxujy + vyujx \u2014 gc9c(f>) = fz \u0399\u03c7\u03c7\u03ce\u03c7 \u2014 {Iyy \u2014 Izz)U}yU}z = \u03a4\u03c7 Iyyujy - {\u0399\u03c7\u03b6 - IXX)UXUZ = Ty LzUz - (Ixx - Iyy)uJxUJy = TZ where \u03c6 and \u0398 stand for roll and pitch angles" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001982_iet-smt:20060039-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001982_iet-smt:20060039-Figure4-1.png", "caption": "Fig. 4 Equipotential lines for electric scalar potential at crosssection y \u00bc 0 in V3D, as well as at u \u00bc 0 and u \u00bc p in Vcyl. Dotted rectangles indicate boundaries of 2D model parts; rest of cross-section represents 3D model part", "texts": [ " 2d, the insulation sheets are shown in an enlarged view. As the transformer is assumed to be placed inside a grounded metallic tank, the electric potential is set to 0 V at boundaries not representing symmetry planes. Fixed potentials are applied at the transformer leg and the transverse yoke (0 V) and at the high-voltage and low-voltage windings (900 V and 0 V, respectively). The field distributions in the (x, z) cross-section at y \u00bc 0 of V3D and in the (r, z) cross-section at u \u00bc 0 and u \u00bc p of Vcyl are shown concurrently in Fig. 4. At the cross-section u \u00bc 0, the equipotential lines expand widely into V3D in the z-direction, as no metallic parts are present above the windings. Towards the boundary in the x(r)-direction, the potential decays smoothly to zero at the metallic wall of the grounded transformer tank. However, at u \u00bc p in the left part of Fig. 4, the equipotential lines are compressed in the z-direction, owing to the presence of the transverse yoke at zero potential. In the negative x(r)-direction, the almost horizontal equipotential lines reflect the proximity to the second parallel high-voltage winding, which is at the same potential. The transition at the interface is continuous, indicating correct coupling between the adjacent domains. Evaluation of the solution on the circular interface between V3D and Vcyl shows the azimuthal dependence of the potential (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003889_j.phpro.2013.11.056-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003889_j.phpro.2013.11.056-Figure8-1.png", "caption": "Fig. 8. Distribution of Austenite (volume fraction (a) ideal case at step 55; (b) the case considering casting at step 55", "texts": [], "surrounding_texts": [ "The stress distribution of the Step Part after casting is shown in Figure 6. The stress of the other area was in a uniform and low level other than the large stress around the sprue gate and the feeder head. It is concluded that the stress distribution of the ideal condition is similar to that considering the influence of the casting, and the stress values of the two states have no differences, see Figure 7. Comparing Figure 6 and Figure 7, the heat treatment produces great stress, and significantly heightens the stress concentration. Besides, the stress concentration around the sprue gate and the feeder head generated by casting could be eliminated by heat treatment. 4.3. Analysis of the phase transition Comparing figure (a) with figure (b) in Figure 9~10 separately, the stress generated during casting cannot change the nature of the organization transformation. Furthermore, the distribution trends of the two cases have little difference. The stress generated by casting cannot influence the organization transformation obviously." ] }, { "image_filename": "designv11_61_0000285_cdc.2003.1271669-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000285_cdc.2003.1271669-Figure5-1.png", "caption": "Fig. 5. Control objective", "texts": [ " To make the COG of the center stick approach the origin, it is better to keep a relationship between 8 and qi as constant, that is 8 - qi = constant. So we have to control: 1) distance from the origin to the COG: T 2) difference between the argument and the attitude angle: 3) angular velocity: $ 8 - 4 to realize an enduring rotary motion of the devil stick. To begin with, we try to make a radius r converge to a desired value r d , that is to make the COG of the center stick follow a circle where its radius is r d (See Fig. 5). Then we try to make (8 - 4) converge to some consfant value a. Finally we don't control the angular velocity 4 and leave it as zero dynamics. In.other words, we control only two state r, 8, and analyze q5 as zero dynamics. IV. SIMULATION A. Simulation of the rotary motion by output zeroing Based on SIII, we set the output function as Though this output function doesn't contain $, we will derive a condition about stable 4 by analyzing zero dynamics. By linearizing (13), we can obtain the following linear state equation" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002570_j.tsf.2008.10.058-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002570_j.tsf.2008.10.058-Figure2-1.png", "caption": "Fig. 2.HAADF-STEMand EELS characterization of Si\u2013Ti\u2013SiOx substrate. a)HAADF-STEM image o in the energy range of Si\u2013L2,3, C\u2013K, Ti\u2013L2,3 and O\u2013K edges from the regions indicated in image", "texts": [ " The silicon oxide film on top of the titanium film is visible, and appears to be of roughly uniform thickness (of the order of 5 nm) and to smoothly follow the titanium surface. Since the titanium film is polycrystalline, its surface is not entirely flat, but is rather punctuated by crystallites that rise out of the surface, as in the centre of Fig. 1. (Images of other areas suggest that these features may be more common than shown in Fig. 1.) A high-angle annular dark field (HAADF) scanning transmission electron microscopy (STEM) image of the surface is shown in Fig. 2a. Electron energy-loss spectroscopy (EELS) measurements were performed in the various regions, i.e. the titanium substrate, the silicon oxide film and the glue used for the TEM sample preparation. The different spectra for the titanium substrate, the silicon oxide film and for the glue are shown in the energy region between 90 and 600 eV in Fig. 2c,d,e, and confirm the chemical identity of the deposited films. The Si\u2013L2,3 electron energy-loss near-edge structure of the SiOx film corrected for the background is shown in Fig. 2b. This spectrum contains two initial sharp peaks that are followed by a sharp and a broad peak and have been attributed to electron excitation into molecular orbitals associated with a silicon atom that is tetrahedrally coordinated [26,27]. Our intentionwas to investigate a silicon oxide, and not a titanium or titanium oxide surface; it is thus essential that the titanium layer should be completely or almost completely covered by the silicon oxide layer over the whole of the substrate. This is shown to be the case by the X-ray photoelectron spectrum (XPS) of the surface (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000267_papcon.2002.1015139-Figure28-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000267_papcon.2002.1015139-Figure28-1.png", "caption": "Figure 28: WPI (450 HP Open Dripproof 5000 Frame, 8-Pole Motor)", "texts": [], "surrounding_texts": [ "INTRODUCTION The effects of variations in voltage, load, speed, starting, ambient temperature, service factor and altitude will each be reviewed. Examples of how these factors influence the motor performance and life are provided. Reasonable operating limits for these variations are provided and suggested rules of thumb are presented that will enhance the motor reliability and longevity. Only horizontal motors will be addressed in this paper. Many of the principles apply to vertical types of motors.\nFigures 1A. 1B. and 1C show a typical applications for horizontal motor intended for use in the pulp and paper industry.\nTEMPERATURE CONSIDERATIONS The motor has three main temperature related elements. They are the stator, rotor and bearing system. Each of these elements must be considered individually and with regard to how they influence each other.\nEach of these elements must be balanced against the motor cooling system which is made up of conduction (the frame, lamination and shaft), radiation (end turns, rotor end rings and frame) and convection (internal and external fans and cooling surfaces of the motor).\nThe base sources of heat generation are the 12R losses to the rotor and stator, the core loss of the electrical laminations, the stray load loss associated with the air gap and tooth surfaces of the rotor and stator and the bearing system losses.\nCombining these three elements, the motor is made up of its heat source within the stator, rotor and bearing system, and the cooling system which is made up of the conduction, convection and radiation elements.\nFigure 1A: Typical Severe Duty TEFC Motor,\nFig(\n0-7803-7446-0/02/$17.00 02002 IEEE - 1 1 5 -", "INTAKE INTAKE INTAKE EXHAUS1 COOLING AIR COOLING AIR COOLING AIR HOT AIR\nFigure 1C: Weather Protected II (WPII)\nAll of these must be balanced so the heat generated by the motor load and ambient can be satisfactorily disposed of through the motor cooling system while maintaining an equilibrium and without damaging the.stator, rotor or bearings.\nFigure 2A graphically i l lustrates the basic motor elements, heat sources and cooling system for a typical TEFC horizontal motor.\nFigure 2B and Figure 2C for WPI and WPll motors.\nFigure 2C: WPll Motor (8000 Frame WPII, 6-Pole Motor)\nTHERMAL AGING PROCESS The thermal aging process as shown in Table 1 is always present and is occurring even when the motor is not running. At one extreme it is aging at the rate caused by the ambient temperature to which the winding is exposed. The other extreme is when the motor is operating under service factor conditions (limited to 155\u00b0C Class F insulation) average winding temperature when the motor is running.\nOther stresses present while the motor is running include dielectric, mechanical and environmental stresses (which may also be present when the motor is not running). At some point the thermal aging renders the winding insulation vulnerable to these stresses and the system begins to \"short out\" between turns, or to ground at which time the insulation system has failed by definition.\nFigure 3 provides temperature l ife curves for the standard motor insulation systems that are available today. These curves assume that the insulation life doubles for a 10\u00b0C decrease in temperature.\nTABLE 1 THERMAL AGING PROCESS\n1. Oxidation 2. Loss of Volatile Product 3~ Molecular Polvmerization 4. Reaction to Moisture 5. Chemical Breakdown 6. Vulnerable to Other Stresses\n-116-", "1.wo WO\nClass of Insulation System A B F H (see MG-1-1.65) Time Rating (shall be continuous or any short-time rating given in MG-1-10.36) Temperature Rise (based on a maximum ambient temperature of 40\"C), Degrees C 1. Windings, by resistance method\nfactor other than those given in items 1.c and 1.d.\nhigher service factor\na. Motors with 1.0 service 60 80\nb. All motors with 1.15 or 70 90 5 1,000 Y 3\n-\n105 125\n115\n-\n100 60 80 100 120 140 160 180 200 220 240\nTOTAL WINDING EMPERATLIRE . Degrees C\nFigure 3: Temperature Vs. Life Curves for Insulation System (per IEEE 117 & 101)\nNEMAllEEE INSULATION CLASSIFICATIONS AND TEMPERATURE RISE\nThe following section comes directly from the newly released IEEE 841-2001 Standard. 5.4 Insulation system and temperature rise a) Insulation system\n1. The motor shall have a non-hygroscopic, chemical and humidity-resistant insulation system. The thermal rating of the insulation system shall be a minimum of Class F as defined in section 1.66 of NEMA MG 1-1998. A lead wire having a temperature rating that is more than 5'C (9\u00b0F) less than the temperature rating of the insulation system in which it is connected shall be separated from the windings by a barrier or envelope of a material compatible with the system. The temperature rating of the lead wire shall not be less than 125'C. b) Temperature rise - When operated at rated voltage, frequency, and power, the average temperature rise of any phase of the stator winding shall not exceed 80\u00b0C by winding resistance.\nNEMA Temperature Limits: Note that in Table 2. NEMA allows for a Class F system with a 1.15 service factor to have a temperature rise of 115\u00b0C. whereas, the iEEE 841 standard limits the temperature rise to 80%. The result is a theoretical increase in the thermal life of the winding from 20,000 hours to over a million hours when operated at full load.\n- 117-" ] }, { "image_filename": "designv11_61_0002004_s1007-0214(07)70003-2-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002004_s1007-0214(07)70003-2-Figure3-1.png", "caption": "Fig. 3 Two-link manipulator", "texts": [ " (19) is replaced by a saturation function of the form: ( ) ( )T T T T T sgn , if ; sat , if i i i i i \u03b7 \u03b7 \u03b7 \u23a7 > \u23aa\u23aa= \u23a8 \u23aa \u23aa\u23a9 E Pb E Pb E Pb E Pb E Pb \u2264 where 0\u03b7 > is a small constant required in order to remedy control chattering. Figure 2 illustrates the overall scheme of the robust adaptive neural networks controller presented in this paper. Tsinghua Science and Technology, February 2007, 12(1): 14-21 20 3 Simulation Example Since the dynamics of robot manipulators is highly nonlinear and may contain uncertain elements such as friction, we can use a two-rigid-link-robot manipulator to verify the effectiveness of the proposed control scheme. The two-rigid-link-robot manipulator is shown in Fig. 3. Its dynamic model[12] is as follows. ( ) ( , ) + ( )+ ( )+ ( )= ( ),m t t+M q q V q q q G q F q T \u03c4 2 2 2 2 2 1 1 2 1 2 2 2 2 2 1 2 2 2 2 2 2 2 1 2 2 2 2 2 ( )+2 cos cos ( ) , cos l m l m m l l m q l m l l m q l m l l m q l m \u23a1 \u23a4+ + + = \u23a2 \u23a5 +\u23a2 \u23a5\u23a3 \u23a6 M q , ) 2 1 2 2 1 2 22 2 1 2 2 21 2 ( 0 5 )sin ( ) sin m l l m q q . q q , l l m q q \u2212 +\u23a1 \u23a4 = \u23a2 \u23a5 \u23a3 \u23a6 V q q q 1 1 2 1 2 2 1 2 2 2 1 2 ( ) cos + cos ( ( )= , cos ( ) l m m g q l m q q l m g q q + +\u23a1 \u23a4 \u23a2 \u23a5+\u23a3 \u23a6 G q [ ]T 1 1 2( )= 12 0 5sgn ( ) 12 0 5sgn ( ) ,q " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002258_b978-1-4831-9821-7.50021-1-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002258_b978-1-4831-9821-7.50021-1-Figure1-1.png", "caption": "FIG. 1. Radially constrained ring under inertial load.", "texts": [ " It has been shown that, using this improved scheme, a converging solution can be obtained for the threshold dynamic-buckling load of a spherical cap using a reasonably small number of mesh-points. The general numerical method has also been extended to the analysis of the dynamic response of unbonded concentric rings [5] where a simple relation is assumed for the collision between the rings. The present paper is concerned with an extension of the general method to the study of dynamic buckling of a ring which is constrained in a rigid circular surface (Fig. 1) and is subjected to a transiently applied inertial loading. In this case a relationship accounting for the collision between the ring and the rigid boundary is required. A method has been developed to calculate the deformation and the snap buckling behavior of such a ring under statically applied inertial loading [6]. Because of the use of large deformation equations an iteration procedure was necessary in this analysis. A special feature of the present problem is that, under varying loading amplitudes, the angle *, which defines the point of separation between the ring and the rigid surface, also varies continuously. T H E O R E \u0389 C A L F O R M U L A T I O N To investigate the dynamic buckling behavior of a thin ring, it is again necessary to consider the large-deformation behavior. For thin rings, however, the effects of rotary inertia and transverse shear deformation may be neglected. 285 Governing Equations Figure 1 illustrates the internal and external forces acting upon a deformed element of length, ds, of a two-dimensional structure. The differential equations of dynamic equilibrium of this structural element in the y and \u03b6 directions, respectively, are (d/ds) (N cos \u0398) - (d/ds) ( \u03c1 sin \u00d6) + - m F = 0 (d/ds){N sin \u0398) + (d/ds){Q cos e) + F,-mW = 0 (1) (2) where m is the mass per unit length, \u03b8 is the slope of the structure (\u00d6 = sin \\dW/ds)l and all other quantities are defined in Fig. 1. Dynamic buckling of a circular ring 287 The equation of moment equihbrium about an axis perpendicular to the yz plane is dM/ds-Q = 0 (3) where rotary inertia has been neglected. The finite-difference version of the above equations is : iVf+1 cos 0,. +1 - Ni cos Oi -Qi+i sin 0,\u00b7 + ^ + Qi sin + F,..[(As, + As,^,)/2]-m,F;- = 0 iVj+i s i n \u00f6 f + i - i V j . s i n \u00f6 i - l - \u00f6 i + i c o s \u00f6 i + i - \u00f6 i C O S \u00d6 f + F J ( A s , + A5 ,^J /2 ] -m,W^. = 0 \u039c , - \u039c , _ \u03b9 - \u03c1 , \u0391 5 , = 0 (6) where the subscript i denotes either the ith spanwise station or the ith element which is between the ( i - l)th and the ith stations" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002424_j.sysconle.2008.09.007-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002424_j.sysconle.2008.09.007-Figure1-1.png", "caption": "Fig. 1.", "texts": [ " The function g1(x)f2(x) \u2212 g2(x)f1(x) is called as the criterion function for global controllability, denoted by C(x) [7]. We remark that Theorem 2.1 is also valid for the case that \u2212g(x) is locally asymptotically stable to its equilibrium point 0. Now let us give an intuitive interpretation for the above theorem, whose rigorous and detailed proof is given in Section 3. By the uniqueness of solutions of ordinary differential equation, the plane R2 is partitioned by all the control curves corresponding to all initial conditions in R2. This results in a foliation which can be viewed on the whole as Fig. 1, where \u0393i (i = 0, 1, 2, . . .) are some regular control curves of (1) and \u03b3i (i = 1, 2, 3, . . .) are some 1 The two ends of a curve \u0393 (t), t \u2208 R extending to infinite means that: \u2016\u0393 (t)\u2016 \u2192 +\u221e, when t \u2192+\u221e and\u2212\u221e. control curves whose positive semi-trajectories tend to the origin. According to the Jordan curve and Jordan curve-like Theorem [7,16], each regular control curve separates the plane into two disjoint components. If the criterion function C(x) changes its sign over the regular control curve \u03932, then there exist two points x1, x2 on \u03932 such that for any control u, the positive semi-trajectory of the control systems with initial points x1 and x2 will go into different sides of the curve \u03932, as shown in Fig. 1. Hence, for example, at the point x1, there exists a neighborhood U(x1), such that the positive semitrajectory with any initial point in U(x1) will go into Side-B under any control function u. Let the control function u be sufficiently large on a tubular neighborhood of \u03932 which contains both x0 and x1 with the direction of vector field from x0 to x1. Then under this control, the positive semi-trajectory of the system (1) with initial point x0 will reachU(x1) at some finite time, sowe can force the trajectory to go into Side-B of \u03932 in U(x1)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002100_jmes_jour_1969_011_071_02-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002100_jmes_jour_1969_011_071_02-Figure4-1.png", "caption": "Fig. 4. Involute overlap area", "texts": [ " Impossibility of radial assembly and disassembly J O U R N A L M E C H A N I C A L E N G I N E E R I N G SCIENCE ORIGIN OF TIP INTERFERENCE AND Usually, tip interference is explained kinematically (3)-(6) (8). The tips of meshing teeth may overtake each other outside the line of action when the gears are turned, Fig. 3. This consideration produces the condition given below, but does not give insight into the origin of the phenomenon itself. A geometric treatment (2) (7) is more instructive. Meshing tooth flanks are tooth flanks of which the involutes touch on the line of action. These involutes have unequal base circles. Hence, they have different curvatures and may intersect each other, Fig. 4. There is no tip interference if and only if the point of intersection of the tip circles lies outside the involute overlap area. The condition for non-interference a t any particular centre distance may be written: CONDITIONS FOR NON-INTERFERENCE T > O T = z,(+, -inv ul, +inv aal) i I:: -zZ(&-inv u,, finv uL,12) razz - ra12 - az cos = T b 2 - r b 1 cos a,', = ~ a rbl cos ual = - ra 1 cos U g 2 = - r a 1 rbZ a i Fig. 3 . Tips of meshing teeth Vol I1 No 6 1969 at NANYANG TECH UNIV LIBRARY on June 5, 2016jms" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002085_iros.2007.4399111-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002085_iros.2007.4399111-Figure1-1.png", "caption": "Fig. 1. Wheeled vehicle", "texts": [ " Therefore, it is expected that smaller number of subproblems are created than the standard method, since only a few collisions at most are caused for one predicted trajectory in many formation control problems. Numerical examples and experiments applying the proposed method to the formation control method[12] show that the proposed method drastically reduces computation time. We consider a group of n unicycles indexed by i = 1, \u00b7 \u00b7 \u00b7 ,n: x\u0307i = vi cos\u03b8i, y\u0307i = vi sin\u03b8i, \u03b8\u0307i = \u03c9i, (1) where vi and \u03c9i are the linear and angular velocities of the vehicle i respectively, and (xi,yi,\u03b8i) denotes the measurable coordinate with respect to a global frame (see Fig. 1). We also define a leader vehicle described as: x\u0307r = vr cos\u03b8r, y\u0307r = vr sin\u03b8r, \u03b8\u0307r = \u03c9r, (2) where vr and \u03c9r are the linear and angular velocities respectively, and (xr,yr,\u03b8r) denotes the measurable coordinate with respect to the global frame. The reference position of the vehicle i in (1) is given as a constant vector (ri, li) in a local frame on the leader vehicle in (2) (see Fig. 2). In other words, the reference trajectory for the vehicle i is given with respect to the global frame as zd i := [ xr + ri sin\u03b8r + li cos\u03b8r yr \u2212 ri cos\u03b8r + li sin\u03b8r ] " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002762_0010-406x(70)90076-9-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002762_0010-406x(70)90076-9-Figure1-1.png", "caption": "FIG. 1. Electrode with self-contained oil reservoir. The reservoir is raised and lowered by means of a hydraulic advance made from two syringes connected by a length of plastic tubing.", "texts": [ " In the nerve-muscle experiments, tension and a stimulus marker were recorded directly on a Grass polygraph. In the experiments on reflexly produced retraction (Section 3), the reflex was evoked by a brief, high-velocity mechanical stimulus applied to the foot or collar from a solenoid stimulator or a hand-held probe (Olivo, 1970a). The duration of the mechanical stimulus was usually less than 30 msec (rise time plus fall time). In the reflex experiments, nerve activity was recorded from intact nerves by means of Ag-AgC1 electrodes in self-contained oil reservoirs (Fig. 1) ; such electrodes permit recording nerve activity without flooding the bath with oil and without lifting nerves large distances from their courses within the foot musculature. Recordings were made simultaneously from as many as five nerves in the reflex pathway. After conventional amplification, nerve activity and tension were recorded on a 7-channel FM magnetic tape recorder. For analysis, selected channels were played back to a 4-channel polygraph. Because the high-frequency cut-off of the polygraph penwriter was too low for accurate reproduction of nerve activity, the tape was played back at reduced speed in order to bring more of the signal into the penwriter's passband" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003917_icsem.2010.22-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003917_icsem.2010.22-Figure3-1.png", "caption": "Fig. 3(a) shows the instantaneous contact line of tooth surface of original toroidal worm gearing. Fig. 3(b) shows the real structure of worm surface. The contact line in Fig. 3 is accordant with the theory. The work area concentrates in a narrow area which lies in the middle of the tooth surface of worm. The central of the tooth surface is a limit line. The limit line is not only the limit to distinguish work area and non-working area but also the weakness of meshing. The purpose of modification is to eliminate the limit line of worm and extend the work area, thus improves the carrying capacity of worm bearing.", "texts": [], "surrounding_texts": [ "About the early theory of modification, the former Russian scholar . . pointed in the book of gear geometry and applied theory[1]: the purpose of modification of toroidal worm is to make a specified amount of deviation. So, the worm drive can have low sensitivity to errors which is caused during manufacture and assemblage and the worm 978-0-7695-4223-2/10 $26.00 \u00a9 2010 IEEE DOI 10.1109/ICSEM.2010.22 6059 can easily run in and obtain high carrying capacity. This opinion, which was based on the modification theory of involute gear, was summed up mainly through practices and experiments, becomes the dominative knowledge about the modification of toroidal worm. With the appearance of more and more new types of toroidal worm and the advancement of the manufacture technology of toroidal worm, it is discovered that the traditional modification theory can not meet the technology of toroidal worm. The main problems are: (1) The traditional modification theory can\u2019t explain the essence of the modification of toroidal worm and can not break through the constraint of the modification theory of involute gear. It doesn\u2019t recognize the importance of teeth profile along the relative motion direction of toroidal worm. All of above are negative to improve the performance of meshed gear and the manufacturing quality. (2) If the principle of modification can\u2019t be made clear, the problem of modification can\u2019t be solved. Why the toroidal worm has high performance? To make its performance best, what can be done? The correct way of the optimum design of toroidal worm can\u2019t be found. For this reason, although plane envelope toroidal worm has been manufactured for more than 30 years, its serialization and standardization of geometric parameters haven\u2019t been completed until this research is finished. III. THE STUDY ON THE NEW THEORY OF TOROIDAL WORM 3.1 The study method of modification theory Due to the complexity of spacial curve of the tooth d s e a li curve. the pri hei pro m al of t concept of modification curve refers to the modification curve of reference circle if there is no special announcement. With this method, the relation between the modification and tooth surface of toroidal worm has been built and the research process has also been simplified. The modification rule of whole toroidal worm has been summed up. Finally, the principle of modification has been established. The performance of single plane enveloping toroidal worm is decided by the original factors. It belongs to the forming theory. Based on the double enveloping, the modification research focuses on the influence on the performance of the worm pair caused by the tooth surface of worm. The best method to study the enveloping process and meshing performance is the spacial meshing theory[2]. According to kinematics of spacial meshing theory, the meshing point between the surface of worm and worm wheel must meet the meshing equation : =n v=0 (1) n \u2014\u2014normal vector of meshing point \u2014\u2014the relative velocity vector of meshing point Practice has proved that the modification of toroidal worm is not random but has its rules. What is the rule then? The rule must be proved on the theory besides in practice (include running-in and experiment). From microscopic viewpoint, the modification is the value from one point to the standard tooth surface on the deviated direction. If the base is the original tooth surface, can the modification tooth surface deviate from it at random? What is the basis to judge the modification value is positive or negative if it is supposed that the reduced direction of tooth surface of worm is the surface of toroidal worm, it is difficult to study the whole curve. In order to study the meshing relation of tooth surface between worm wheel and worm, the worm should be divide into some equal parts along the tooth height and each part is a narrow helicoids. While the number of divided part increases to infinity, each part of the tooth surface will become a ring helical line and the tooth surface of worm consists of a set of ring helical lines. Thus, the study of th spacial surface transforms into the study of a set of ring helical lines during the research of tooth surface of worm. The tooth surface of worm wheel can be subdivided in the same way. Every ring helical line of worm corresponds to tooth surface curve of the worm wheel. Based on the helical ne of the primal worm, the deviation between helical line of the current worm and the primal worm is the modification object. In other words, based on the primal worm, expanding the helical line of the current worm shapes a modification Through the study on the meshing relation between modification curve and the worm, the modification nciple of toroidal worm can be studied. The tooth surface of worm along the direction of tooth ght consists of numerous modification curves. Fig. 1 shows that every modification curve represents the tooth file of worm on the special position. So, to study odification curve is actually to discuss the tooth profile ong the helical line. The reference circle lies on the middle he working tooth, so its modification curve is most representative and usually regarded as the representative of the tooth surface of worm. In the following sections, the positive direction? Fig. 2 shows that the point M and N all lie on the original tooth surface. To point N, there may be two directions after modifications which are de and \u2013de. Are both directions satisfied with the meshing conditions? This is what the modification theory will solve. principle of the toroidal worm. Because the derivation is based on the curvature radio, the principle is also named as \u201cthe curvature modification\u201d. The meshing equation of the kinematic method is the necessary condition of the meshing of the conjugate hyperboloids on theory. The mathematical expression of the curvature modification derives from the meshing eqution, so it is the necessary condition of the meshing of the toroidal worm. It shows the changing law of the toroidal worm\u2019s thickness along the direction of helix (the tooth length) and also decides the shape and the feature of the toroidal worm\u2019s tooth surface. So, the essence of curvature modification is to find the optimum tooth profile along the direction of the toroidal worm\u2019s tooth length. 3.3 The effect of curvature modification In (2), the variable is not concerned in the variable e, so: e= +c( >0) (3) is the normal curvature radius along the direction of the surface relative velocity of toroidal worm, and: 2 2 2 d 2 d d d 2 d d d 2 E u F udv G v L u M u v N v \u03c1 + += + + (4) E F G\u2014 the first fundamental variable of tooth surface L M N\u2014 the second fundamental variable Because the same modification curve is on a helix, u is a constant, so: du=0, dv=d 1, and the normal curvature radius along the direction of relative velocity is: G N \u03c1 = Through long and hard work, the curvature modification theory, which is a great progress during the modification research, is summarized. 3.2 The curvature modification theory To the tooth surface of toroidal worm, the curvature radio is decided by its shape and the parameter de is its inherent changing law. Based on the kinematic method of spatial meshing principle, the relation between the modification and the tooth surface\u2019s curvature radio, which proved the feasibility of modification, is derived. There has: d e= d ( >0) (2) de---the differentiation of the tooth surface\u2019s modification of worm; d ---the differentiation of normal curvature radius along the direction of the surface relative velocity of toroidal worm; (2) is the mathematical expression of the modification (5) According to the boundary condition of e, and c can be calculated. So, (3) only includes u and 1. To choose diferent value of u (if 22 2 brRu \u2212= , it lies on the pitch circle), the helix at certain height along the tooth height of worm will be got. And the modification value at different points can be obtained by choosing different 1. Drawing the modification value and the worm\u2019s rotation angle 1 correspondingly in coordinate system, shown as Fig. 4(b), the curve of normal curvature radius will be got. The curve of curvature radius of toroidal worm shows that the variety of curvature radius between the working tooth surface (the entrance angle i and the exit angle 0 ) is approximately monotonous, except for an extreme point near the exit. Due to the special changing law of the toroidal worm, it is possible that the modification technology make rational use of the law to improve the meshing capability. Usually, the curvature radius is thought as the modification reference during the research of modification theory. And it will not influence the scientificity of the study. According the above-mentioned calculations, the curve of curvature radius is drawn. This curve is the basis of modification of toroidal worm. To ensure the working performance of toroidal worm, the modification must be based on the changing rule of the curvature curve. So, the performance of worm can be improved by the modification. The modification of toroidal worm is to change the tooth thickness of worm. The change of the tooth thickness makes the toroidal worm shape a new tooth profile. So, the new tooth profile forms the tooth shape along the direction of worm\u2019s tooth length. To modify according to the curve of curvature radius, the tooth profile along the direction of tooth length is the curve of curvature radius. The Fig. 4(a) shows the contact line between worm and worm wheel after the modification. The contact lines almost cover the whole tooth surface. The related parameters of worm pair: the transmission ratio i0=40; the number of worm wheel tooth, which is surrounded by the worm, is 4. As shown in Fig. 4(a), there are contact lines on the tooth surface from the surrounded teeth (1~5). This shows that the worm always meshes with the worm wheel from the entrance to the exit. performance but also shape a properly tooth profile. Compared with other traditional modification, the changing rule of parabola modification is closer to the curvature modification, therefore it has better performance. More exactly, the parabola modification is the modification of conic fitting of curvature curve. Through long research on many toroidal worms, it is proved: the closer the modification curve to the curvature curve, the better the performance. And during the study on the second plane envelope toroidal worm, it is discovered: the modification curve with bigger plane angle is very close to the curvature curve and the curve with smaller is on the contrary. So, the performance of plane envelope toroidal worm with bigger is higher than tat of the worm with smaller . The study on the curvature modification is focused on the deviation from the modification tooth surface to the original surface and does not relate to the method of modification and the form of main parameters. So, the rule of curvature modification is suitable to all types of toroidal worm. The study to the plane envelope toroidal worm also proves that the rule of curvature modification is suitable to it. IV. CONCLUSION (1) The principle of modification solves the key theory and mechanism of the modification of toroidal worm. A reasonable answer is found both theoretically and practically to the long-term pending problem of the modification of toroidal. Curvature modification can guarantee that toroidal worm has the greatest mating tooth number, the biggest working area and the best working performance. Experiment also proves: compared with the traditional modification, the modification based on the principle of curvature modification can bring higher bearing ability and transmission efficiency. The principle of curvature modification, which does not aim at a certain kind of toroidal worm but studies all toroidal worms as a whole, is a common law and suits to all toroidal worms. (2) Including as a technological measure to improve transmission stability, the modification principle of toroidal worm is a theory and technology to study the basic structure of toroidal worm. Confirming the essence of the modification of toroidal worm is to construct the tooth profile of toroidal worm breaks through the conventional theory of mending tooth shape and opens up a new area for the theory and practice of toroidal worm." ] }, { "image_filename": "designv11_61_0003650_iccpct.2013.6528935-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003650_iccpct.2013.6528935-Figure3-1.png", "caption": "Fig. 3 Eight possible switching configuration of the VSI", "texts": [ " 2. The VSI has eight possible voltage vectors, in those there are six active voltage vectors (U,-U6) and two zero voltage vectors (U7,U8), which can be achieved using the combination of the inverter switching states Sa, Sb, and Sc' When the upper part of switches are turned on, that indicating switching value is '1' and when the lower part of switches are turned on that indicating switching value is '0'. The eight possible voltage vector (U\\-U8) with inverter switching configuration is shown in Fig. 3 [13-15]. The vector control principle on AC machines takes advantages of transforming the variables from the physical three phase 'abc' system to two phase stationary frame 'dq' system. The stator phase voltages and associated stator voltage vectors are shown in the following matrix equation (1): [u,nj [2 -I -I][S ] \ufffdbn = U ;c -I 2 -I S: u en -I -I 2 Se (1) Where Uan, Ubn, and Uen indicate the phase voltages, Sa' Sb ' and Sc indicate the voltage source inverter switching states. The inverter output voltages Uan , Ubn , and Ucn are converted to u\ufffd and U\ufffds by following equations (2), and (3): \ufffd UOC( ) (2) Uds =-3- 2Sa -Sb -Sc -p 1 Uqs= " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001773_bf03546396-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001773_bf03546396-Figure1-1.png", "caption": "FIG. 1. Centralized Control Strategy.", "texts": [ " The formation flying concept relies on accurate control of the relative positions and orientations between participating spacecraft. Typically, the spacecraft fleet pools computational and navigational resources. Therefore, for ac curate control, command and navigation data must be exchanged between space craft. Command and navigation cross-link requirements vary depending on the formation control strategy. Control strategies range from centralized to fully dis tributed with many hybrid variations. Centralized and decentralized control, how ever, represent the extremes when discussing cross-link requirements [3]. Figure 1 illustrates the major characteristics of centralized control. Vehicle 1 is the supervisory spacecraft and is responsible for all estimation and control com putations. Navigational data are sent from each of the subordinate spacecraft, Vehicles 2 and 3. Upon receiving all of the data, the supervisor computes the entire formation state and transmits control commands back to the subordinate spacecraft. This strategy yields very good performance, but it has large communication re quirements and a significant computation load for supervisory spacecraft" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003445_9781118361146.ch7-Figure7.2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003445_9781118361146.ch7-Figure7.2-1.png", "caption": "Figure 7.2 (a) Cross-section through a four-pole DC motor. The dashed lines show the magnetic flux. The motor torque is clockwise. (b) Shows the convention used to indicate the direction of current flow in wires drawn in cross-section", "texts": [ " The most important of these are as follows: \u2022 The rotating wire coil, often called the armature, is wound round a piece of iron, so that the magnetic field of the magnets does not have to cross a large air gap, which would weaken the magnetic field. \u2022 More than one coil will be used, so that a current-carrying wire is near the magnets for a higher proportion of the time. This means that the commutator does not consist of two half rings (as in Figure 7.1) but several segments, two segments for each coil. \u2022 Each coil will consist of several wires, so that the torque is increased (more wires, more force). \u2022 More than one pair of magnets may be used, to increase the turning force further. Figure 7.2a is the cross-section diagram of a DC motor several steps nearer reality than that of Figure 7.1. Since we are in cross-section, the electric current is flowing in the wires either up out of the page, or down into the page. Figure 7.2b shows the convention used when using such diagrams. It can be seen that most of the wires are both carrying a current and in a magnetic field. Furthermore, all the wires are turning the motor in the same direction. If a wire in an electric motor has a length l metres, carries a current I amperes and is in a magnetic field of strength B webers per square metre, then the force on the wire is F = BIl (7.1) If the radius of the coil is r and the armature consists of n turns, then the motor torque T is given by the equation T = 2nrBIl (7.2) The term 2Blr = B \u00d7 area can be replaced by , the total flux passing through the coil. This gives T = n I (7.3) However, this is the peak torque, when the coil is fully in the flux, which is perfectly radial. In practice this will not always be so. Also, it does not take into account the fact that there may be more than one pair of magnetic poles, as in Figure 7.2. So, we use a constant Km, known as the motor constant, to connect the average torque with the current and the magnetic flux. The value of Km clearly depends on the number of turns in each coil, but also on the number of pole pairs and other aspects of motor design. Thus we have T = Km I (7.4) We thus see that the motor torque is directly proportional to the rotor (also called armature) current I . However, what controls this current? Clearly it depends on the supply voltage to the motor, Es. It will also depend on the electrical resistance of the armature coil Ra", " These copper losses are probably the most straightforward to understand and, especially in smaller motors, they are the largest cause of inefficiency. The second major source of losses is the iron losses, because they are caused by magnetic effects in the iron of the motor, particularly in the rotor. There are two main causes of these iron losses, but to understand both it must be understood that the magnetic field in the rotor is continually changing. Imagine a small ant clinging to the edge of the rotor of Figure 7.2. If the rotor turns round one turn then this ant will pass a north pole, then a south pole, and then a north pole, and so on. As the rotor rotates, the magnetic field supplied by the magnets may be unchanged, but that seen by the turning rotor (or the ant clinging to it) is always changing. Any one piece of iron on the rotor is thus effectively in an ever-changing magnetic field. This causes two types of loss. The first is called \u2018hysteresis\u2019 loss, and is the energy required to magnetise and demagnetise the iron continually, aligning and realigning the magnetic dipoles of the iron" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002607_978-0-387-77747-4_6-Figure6.7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002607_978-0-387-77747-4_6-Figure6.7-1.png", "caption": "Fig. 6.7 Flow field model for hydrostatic force (left) and hydrodynamic force (right) [19]", "texts": [ " The model introduced below is capable of dealing with all the elements of, (1) micro-devices, (2) dynamic response characteristics of axial flow hydrostatic journal bearings, (3) evaluation of stiffness, natural frequency and damping, (4) evaluation of instability boundaries, and (5) evaluation of effects of imbalance and bearing anisotropy. Some of these aspects have been studied in depth by a number of researchers (for example Smith [30], Larson and Richardson [15], Fuller [10], Childs [4], San Andres and Wilde [27], Arghir and Frene [2], Orr [23]). The key feature of this model is that it is able to analytically deal with all of these aspects together. 5 The effect of unsteady flow on the whirl ratio will be discussed in Section 6.5.2. Consider the simple journal bearing geometry depicted in Fig. 6.7. Assuming that the rotor is stationary and that the center of the rotor is offset from the center of the journal as depicted in Fig. 6.7 on the left, the static pressure drop across the gap sets up a Poiseuille type flow in the axial direction for each circumferential segment C(\u03b8 )\u00b7Rd\u03b8 . The local journal clearance is denoted by C(\u03b8 ) = C0\u2013e cos\u03b8= C0\u00b7 (1 \u2013\u03b5cos\u03b8 ), where e and R are the eccentricity and the mean journal radius respectively, and \u03b5 = e/C0 is the eccentricity normalized by the bearing clearance C0. The local axial flow rate is obtained by integrating the Poiseuille flow profile in the radial direction and yields qx(\u03b8 ) = \u2212C (\u03b8)3 12\u03bc \u2202p \u2202x (6", " The rotating motion of the rotor drags the flow along the circumference, and under the assumption of incompressible, fully developed viscous flow, the combination of a Couette flow and a Poiseuille flow is locally set up. For this type of flow the local tangential flow rate is q\u03b8 = RC (\u03b8) 2 \u2212 C (\u03b8)3 12\u03bc \u2202\u03b4p R\u2202\u03b8 . (6.4) The first term is due to the Couette flow and proportional to the local journal clearance. R is the rotor surface speed. When the rotor moves closer to the wall, the fluid will be forced to flow also in the axial direction and thus builds up pressure as depicted in Fig. 6.7 on the right. Assuming again Poiseuille flow in the axial direction (see Eq. 6.2), the hydrodynamic pressure distribution in the journal can be obtained from continuity \u2202qx \u2202x + 1 R \u2202q\u03b8 \u2202\u03b8 = 0, or \u2212 C (\u03b8)3 12\u03bc \u22022\u03b4p \u2202x2 + 2 \u2202C \u2202\u03b8 \u2212 1 R2 \u2202 \u2202\u03b8 ( C (\u03b8)3 12\u03bc \u2202\u03b4p \u2202\u03b8 ) = 0. (6.5) Since L/D is much smaller than 1, the third term in the above equation is negligible compared to the other two terms. Assuming \u03b4p(0,\u03b8 ) = \u03b4p(L,\u03b8 ) = 0 as boundary conditions, the hydrodynamic pressure distribution then yields after integration \u03b4p (x, \u03b8) = 3\u03bc C (\u03b8)3 \u2202C \u2202\u03b8 \u00b7 x \u00b7 (x \u2212 L) " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003003_1.34662-Figure11-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003003_1.34662-Figure11-1.png", "caption": "Fig. 11 Excitation of a SDOF system by a rotating unbalance.", "texts": [ " An unbalance mass rotating at a constant speed generates a harmonic excitation of constant frequency. However, when the speed varies, the amplitude of the cosine part of the force also varies, and sine terms arise due to the tangential acceleration of the unbalance mass. In this section, the complications arising from unbalance forces are addressed for the case of sinusoidally varying frequency. D ow nl oa de d by U N IV E R SI T Y O F C A M B R ID G E o n Ju ne 1 3, 2 01 6 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /1 .3 46 62 An elementary analysis based on themechanical model of Fig. 11, assuming the cosine argument is defined by Eq. (8) and mu M, shows that the EOM is y00 2 y0 y g ph cosp 2 cos g h sinp p2h sinp sin g h sinp (25) where y Mx=mur. The forcing function contains a sine part, and both the cosine and sine parts have time-dependent amplitudes. The time dependence of the amplitudes is sinusoidal, and so one could expand the forcing function into a series of constant-amplitude sines and cosines as was done between Eqs. (9) and (12). An exact solution could be straightforwardly written down but would be extremely cumbersome" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003734_12.2008329-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003734_12.2008329-Figure1-1.png", "caption": "Figure 1. Schematic drawing of the portable setup. For clarity only the main elements are shown. The red arrows represent the direct light from the LED light source whereas the blue arrows indicate the light scattered by the sample.", "texts": [ " Due to the weak absorption cross-section, a small amount of cyt c exhibits only a poor absorption, limiting the sensitivity of the method. In order to increase cyt c absorption, Suarez et al. suggested a multiscattering approach enclosing the cyt c in a highly scattering media made of polystyrene beads (PS) (dielectric spherical particles with diameter of 500 nm) 9. In this way the absorption of cyt c increases due to increased optical path length through the scatter. To implement the above mentioned principle, we use a dark-field configuration and a white light LED as the light source (Fig. 1). Consequently, only the scattered light stemming from the sample and no direct light is detected. Since we aim to Proc. of SPIE Vol. 8572 857218-2 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 05/19/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx work with small amounts of cyt c, directly transmitted light carrying no information with respect to cyt c has to be suppressed (in a bright-field configuration the absorption signal is completely hidden by the direct transmitted light)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002114_11762320801999746-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002114_11762320801999746-Figure1-1.png", "caption": "Figure 1 AUV model with pectoral fins.", "texts": [ " These results show that the adaptive control system accomplishes precise yaw-angle trajectory control in spite of the parameter uncertainties, and the yaw-angle trajectory remains close to the discrete reference trajectory between the sampling periods. The organisation of this article is as follows. The AUV model and the problem formulation are presented in the \u2018AUV model and control problem\u2019 section. An adaptive law for yaw-angle control is derived in the section \u2018adaptive control law\u2019, and the section \u2018simulation results for yaw manoeuvres\u2019 presents the simulation results. Figure 1 shows the schematic diagram of a typical AUV. Two fins resembling the pectoral fins of fish are symmetrically attached to the vehicle. The vehicle moves in the yaw plane (XI \u2212 YI plane), where OI XIYI is an inertial coordinate system. OB XBYB is body-fixed coordinate system with its origin at the centre of buoyancy. XB is in the forward direction. Each fin oscillates harmonically and produces unsteady forces. Use of two fins instead of one fin gives larger control force without increasing the size of the fin" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003890_icccyb.2013.6617610-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003890_icccyb.2013.6617610-Figure1-1.png", "caption": "Figure 1. Approximate ground contact model of the Szabad(ka)-II hexapod robot", "texts": [ " The implementation and measurements of a controller in the digital micro processor of the real robot is not easy to optimize. The model should also include a reliable functioning of the driving of motors as well as the arms [5]. The dynamic model of Szabad(ka)-II robot was implemented in Simulink environment [3], for more details see Chapter 2. B. Rigid body dynamics The Szabad(ka)-II robot is built from rigid perunal (special aluminum) alloy parts, which connected to each other with rotational links, 3DOF per leg [2]. The ends of the feet are equipped with a spongy material (Fig. 1), which enables the robot to walk both on even and uneven ground. This feet solution is suitable to walking on rough terrain similarly at other hexapod robots [26, 27]. The implemented dynamic model passes over the negligible deformation of rigid bodies, i.e. it uses the methods of rigid body dynamics [10, 15, 25]. Only the modeling of ground contact (connection between ground and feet) remains critical. A simple approximate solution has been chosen (known as compliant contact model or soft contact [10]) as in other robot models [14, 15, 25]: punctual geometry of the feet and a spring-damper absorber have have been implemented as substitutes for the elastic impact of the spongy feet (Fig. 1). C. Fixed-step solver The mathematical model only approximates the real system, as there are numerous tiny non-understandable and uncontrollable factors that influence the calculations. The current discrete simulation model also only approximates the more realistic analytical model. The two main calculation solver type methods for the discrete simulation within the Simulink are the fixed-step and variable-step methods. The simpler fixed-step calculation method was chosen to avoid study of the various variable-step methods and their un-examinable working and criteria", "5kg, so the entire weight of the robot is mrobot=6\u00b70.5+2.5=5.5kg. Usually the robot walks on two or three arms, so one arm gets mwalk=3kg of maximum weight (simple m in further equations). If the collisions of rigid bodies do not have such spongy parts the negligible deflection of the spring-damper can be taken as one thousandth part of the length of the robot\u2019s arm: DZ-ERROR=0.1/1000=10 -4 m. A similar magnitude of error tolerance is defined in [14, 16]. Conversely this robot arm equipped with a 5mm thick sponge (Fig. 1) which can deflect about DZ=0.003m under the robot mass. This is a substitute for the separate approximation springdamper model of ground with a 10 -4 m deflection. Then the DZ=0.003m length was the reference for defining the adequate sample time. When the robot steps on the ground with its foot, all its kinetic energy is transferred to the ground and/or the spongy feet, i.e. to the spring-damper approximate model. The ortoghonal speed of the robot\u2019s foot is determined by the walking algorithm. In case of fast walk the typical value of the impact velocity on the ground is 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003643_ictee.2012.6208635-Figure22-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003643_ictee.2012.6208635-Figure22-1.png", "caption": "Fig 22-Auto CAD model for explaining the view after the solid is sectioned", "texts": [ " 19 shows application of vsfaceopacity command Which helps us to make the model partially transparent so that internal detail, centre of gravity can be seen. I. Modern Approach Used for Explaining a Question on Section of Solids Question: A cube of 35 mm long edges is resting on the H.P. on one of its faces with a vertical face inclined at 30\u00ba to the V.P. It is cut by a vertical section plane parallel to the V.P. and 9 mm away from the axis and further away from V.P. Draw its sectional front view and the top view. Explanation tools and their application \u2022 Auto CAD model shown in fig. 20 , fig. 21 & fig. 22 are used to explain three dimensional concept \u2022 Drafting video explains the construction process \u2022 Animated power point presentation is used to display all writing work live. Fig. 21-Auto CAD model for explaining the position of cutting plane A square prism edge of base 30 mm and height 50 mm is given .The prism is having its base in H.P. with two edges of base perpendicular to V.P. .Draw the development of lateral surface of prism. Explanation tools and their application \u2022 Auto CAD model shown in fig18,fig 19, fig 20,fig 21&fig 22 are used to explain three dimensional concepts \u2022 Drafting video explains the construction method \u2022 Animated power point presentation is used to display all writing work live Fig. 23-Auto CAD model shows the square prism whose surface has to be developed reference planes make visualization very clear cult to show reference planes and projections Drafting videos can be played at desired speed and at desired point Making drawing on board by instruments is time consuming and laborious Animated power point presentations increase the legibility and saves time" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003957_icra.2012.6224664-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003957_icra.2012.6224664-Figure1-1.png", "caption": "Fig. 1: (a) An environment with two possible paths towards the goal. (b) Monotonical velocity field of (a). (c) Arrival time field with a uniform velocity function. (d) Arrival time field with a monotonic velocity function. Darker color means longer arrival time. By using the monotonic velocity function, taking path B is better according to (d)", "texts": [ " We can exploit the speed function by inserting information about several possibilities that the robot may face in the environment, such as rough terrain, variational distance of obstacles, etc. For example on a plain environment with obstacle, we can determine that the robot is better to move slower at the narrow space or corridor or area next to the obstacle. We then set a smaller speed on the area near the obstacle (for example, using monotonic velocity function which gives more value of speed at an area far from the obstacle), then a shorter travel time will be through an area far from the obstacle (see Fig. 1). It means the safety factor is also taken into account in the arrival time field calculation. For that purpose, we implement a monotonic function denoted by Vi1,j1 = { n\u2016xi1,j1 \u2212xi2,j2\u2016 for xi1,j1 \u2208 F, xi2,j2 \u2208 O 1 otherwise (5) where Vi1,j1 is the speed on the point xi1,j1 , xi2,j2 is the nearest point of obstacle to xi1,j1 , and n is a constant for adjusting the monotonic function\u2019s value, to give more differences on each cell. We have described the arrival time field and have given the example in 2D environment" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003564_iccda.2010.5541299-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003564_iccda.2010.5541299-Figure2-1.png", "caption": "Figure 2. Computation zone", "texts": [ " COMPUTE MODEL A. Model simply For simplification of calculate, several hypothesis were taken as foIlow [ 1]: \u2022 Both the rotator and stator are rigidity body \u2022 Gas flow model is laminar flow \u2022 The temperature of gas film is isothermal \u2022 The medium is ideal gas \u2022 Flow field is axial-symmetry In working condition, flow field between seal faces is an orbicular slice with bulge (correspond with spiral groove), because the flow field is axial-symmetry, may calculate only one zone of the flow field (Fig 2). Operation parameters: w=1038.2 radis, Po=2MPa, Pj=OMPa, B=0.82, Where B is the balance coefficient. Geometrical parameters of the spiral grooves: rj=72mm, ro=91.5mm, rg=81.7mm, y=l, Lr=0.25,hg=5Jlm, Ng=12, where rg is the end radius of groove, y is the ratio of groove width to the width of whole ring, hg is the groove depth, Ng is the number of grooves. Medium parameters: Ideal gas at 300K, with dynamic viscosity of 1.663e05Pa.s, Other characteristics are general. Gas film pressure, which provides the face load and moments, is directly related to the gas film lubrication process", " With the assuming of ideal gas and isothermal conditions, the equation is given by: \ufffd(ph3 ap)+\ufffd(ph3 ap) ax 12,u ax Oy 12,u Oy =\ufffd(phU )+ a(ph) ax 2 at (1) Where h is the thickness of gas film between faces. ,u is the gas viscosity. The left items in (I) denote the change of film pressure along with the coordinate x, y, and the right items in (1) present different dynamic effects of film pressure. Flow inlet is Pressure-inlet, flow outlet is Pressure-outlet, two sides interface are periodic, and the end faces set as waIl (Fig 2). Use unstructured grids, the grids of computation zone are shown in Fig 3. Because the size in Z-axis is so smaIl, the model was divided into two main zones (detailed in Fig 3) to get higher grid quality, and near the area of conjunction, finer grid is needed. In the calculation process, the X-axis direction nodes from 60 to 90, he Y-axis direction nodes from 50 to 60, and the Z-axis direction nodes from 30 to 60, ensure the V5-228 Volume 5 difference of opening force less than 1 %, meshing quantity is about 220000" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002392_ma802825u-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002392_ma802825u-Figure7-1.png", "caption": "Figure 7. (a) Reciprocal lattice that satisfies the observed diffraction pattern (closed circles) and (b) corresponding real lattice in fiber specimen. The diffraction geometry in (a) is based on the azimuthal -scanning data of Figure 6c. The angle of 7\u00b0 between the fiber axis and c-axis in real lattice was determined from the azimuthal angle of 27\u00b0 for the (001) reflection with a spacing of 19.4 \u00c5 (refer to Figure 6c).", "texts": [ " The outer halo indicates a liquid-like association of mesogens in their lateral direction and its location on the equator means that the long axes of mesogens are parallel to the fiber axis. On the other hand, the inner reflections cannot be explained by a simple layer structure, but by the type of frustrated structure with twodimensional lattice.16,17,21,24,25,26 The spacings of these reflections are listed in Table 2. All these reflections can be interpreted as a two-dimensional lattice with a ) 12.9 \u00c5, c ) 20.6 \u00c5, and ) 70.0\u00b0; c is the layer periodicity and a the density-modulated periodicity along the layer. Figure 7a illustrates the orientation of the reciprocal lattice that satisfies the observed diffraction pattern with respect to both the spacing and diffraction geometry. Figure 7b illustrates the corresponding orientation of real lattice in the fiber specimen. Here, the tilt angle (7\u00b0) of c-axis to the fiber axis was determined from the azimuthal angle (27\u00b0) of (001) reflection from the meridian (refer to Figure 6c). Considering that the lateral spacing between neighboring mesogens is around 4.5 \u00c5, one can assume that three mesogens are included within a unit lattice length of a ) 12.9 \u00c5 (along the layer plane). On the other hand, length of c ) 20.6 \u00c5 is fairly larger than 15" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002648_1.2908909-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002648_1.2908909-Figure1-1.png", "caption": "Fig. 1 Schematic diagram of cold rolling", "texts": [ " Results of present investiations are compared to experimental results provided in Ref. 30 . Significant reductions in minimum film thickness and maxium film temperature rise are observed in the presence of starvaion, i.e., scarcity of oil. Mathematical Formulation of the Problem The coupled solution of Reynolds and energy equations incororating lubricant\u2019s rheology has been achieved for thermohydroynamic lubrication of starved line contact inlet zone of roll and trip interface . A schematic diagram of the roll-strip interface long with coordinate system is shown in Fig. 1. In order to reuce the computational complexity, second order Legendre polyomial temperature profile across the film thickness in energy quation has been assumed. Detailed investigations related to temerature profile assumptions across the film thickness are provided n Ref. 31 . It is pertinent to mention here that this analysis treats he bounding solids in the inlet zone as rigid. In this investigation, obatto quadrature technique 32 is used for analysis. References 33,34 also provide application of Lobatto quadrature methodolgy. It is essential to mention here that in this section, only final orms of governing equations are presented in order to compress he paper. Related to derivation of governing equations, Ref. 34 ay also be referred. 2.1 Generalized Reynolds Equation. The derivation of genralized Reynolds equation has been obtained by taking the diverence of lineal mass flux across the film thickness refer to Fig. 1 . he lineal mass flux expression m\u0307= \u2212h/2 +h/2 udz is obtained by the ntegration of velocity expression across the film thickness. The nal form of expression is as follows 34 : m\u0307 = ur + us h 2 \u2212 h 3 \u03041A \u2212 h 3 \u03040 + 2 5 \u03042 B 1 ivergence of lineal mass flux m\u0307 / =0 leads to generalized hermal Reynolds equation as follows 34 : 24503-2 / Vol. 130, APRIL 2008 om: http://tribology.asmedigitalcollection.asme.org/ on 08/07/2017 Terms \u00b7 \u0304ph3 p = 6 us + ur h \u2212 2 \u03041 \u03040 h us \u2212 ur 2 where \u0304p = \u03040 + 0.4 2 \u2212 \u03041 2 3\u03040 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002089_1464419jmbd49-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002089_1464419jmbd49-Figure3-1.png", "caption": "Fig. 3 Four-bar mechanism of the rear suspension", "texts": [ " Multi-body Dynamics JMBD49 # IMechE 2007 at The University of Iowa Libraries on March 16, 2015pik.sagepub.comDownloaded from The front suspension system is of the telescopic type and is assumed to have linear stiffness and damping, and hence does not pose any modelling difficulty. The rear suspension system consists of a monoshock spring and damping combination that acts between a point on the swing arm and a point connected to the output link of a four-bar mechanism with the swing arm as the input link. The layout of the mechanism is shown in Fig. 3. This mechanism results in non-linear suspension characteristics, even if the spring stiffness and damping are linear. The points A and D are connected to the rear frame, the triangle ABF is part of the swing arm with point A at the hinged connection with the rear frame and F at the centre of the wheel, BC is the coupler, and the triangle CDE is the output link, which can rotate at D with respect to the rear frame. The monoshock is connected between the points B and E. The kinematic analysis of a four-bar mechanism can be found in reference [19]; a slightly different mechanism was analysed in reference [12]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000816_icita.2005.307-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000816_icita.2005.307-Figure7-1.png", "caption": "Fig. 7 The Virtual Force Field in Telegame", "texts": [ " danger d is the Proceedings of the Third International Conference on Information Technology and Applications (ICITA\u201905) 0-7695-2316-1/05 $20.00 \u00a9 2005 IEEE boundary between the alert section and the dangerous one. In telegame, alert d is set to 80cm and danger d is set to 10cm. The direction of the force is determined by the vector i j PP obtained by OCB approach. Then map the vector i j PP on the x and y plane, a line segment , , i jP P will be obtained as shown in Fig.5. The direction of the force is along the line segment , , i jP P on the plane. The simulated virtual force field can be seen in Fig. 7. In this system a local Cartesian coordinates is set at the center of the turntable and the working plane of the joystick is mapped to the turntable. Therefore when the user moves the joystick to send position command to the robot, the end effecter will move in space above the turntable and the coordinate of x and y is confined to the turntable. After calculating the key points on the vector of the two robots using the VRFF, the direction and magnitude of the virtual force that will apply on the joystick can be obtained real-timely" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001442_135065005x34080-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001442_135065005x34080-Figure2-1.png", "caption": "Fig. 2 Spring\u2013damper contact model of the ball bearing", "texts": [ " Engineering Tribology JET53 # IMechE 2005 at University of Ulster Library on March 24, 2015pij.sagepub.comDownloaded from Figure 1 shows a schematic diagram of a ball bearing containing balls, inner race, outer race, and cage. Points \u2018A\u2019 and \u2018B\u2019 are the centres of inner and outer races, respectively, under loaded condition. The total energy of this system is considered to be the sum of kinetic energy, potential energy, and strain energy of the springs representing contact and dissipation energy due to contact damping. Figure 2 shows the contact model of the ball on races represented by non-linear springs and dampers. Kinetic energy of the system is the sum of individual kinetic energies of each element and can be formulated separately. The total kinetic energy of balls is Te \u00bc XN i\u00bc1 Ti \u00bc XN i\u00bc1 1 2 mi _ ri \u00fe _ Ra _ ri \u00fe _ Ra \u00fe 1 2 Ii _f 2 i (1) where N is the number of balls in the bearing and fi is the angular displacement of the ball about its centre; then, the displacements in vectorial form are ri \u00bc (ri cos ui)i\u0302 \u00fe (ri sin ui)j\u0302 (2) and Ra \u00bc xa i\u0302 \u00fe ya j\u0302 (3) Considering pure rolling at ball inner race contact as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001424_jmes_jour_1973_015_066_02-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001424_jmes_jour_1973_015_066_02-Figure5-1.png", "caption": "Fig. 5. Castor with wheel permitted to move laterally", "texts": [ " 4, an imaginary cycle of oscillation is performed between f O o with the net area in a clockwise sense giving the energy output per cycle. The area here is negative, showing hysteresis damping. The hysteresis band is narrowed, i.e. stability diminished by either increasing forward speed v or reducing trail t, a fact easily demonstrated practically. 4 CASE OF WHEEL PERMITTED TO MOVE LATERALLY Various oscillating castors were observed on high speed film and instability was found to increase with increasing \u2018sideslop\u2019 of the wheel within its supporting fork (Fig. 5). The model used by Pacejka includes a parameter called \u2018wheel-bearing play\u2019, but this does not permit any relative movement of wheel and fork. One must now distinguish between the inclination of the wheel (and fork) Ow, and the angular position of the contact patch 8, (Fig. 5) since energy is fed into the system through this contact patch. Strictly speaking, c is now 3 but at the change of sign point Oc leads Ow in phase, and the wheel is at or near the limit of slop (for free slop it is at this limit, as seen from the high-speed film), t ow+; 8, Isw+; t 4.1 t ow+; 8, so 8, is close to 8, and c can be taken as - Also, with a sufficiently large slop, Oc may pass over the ee = 0 line before c changes its sign because of the phase lead, giving rise to an energy diagram such as Fig", " The same instability can occur in practical situations when the wheel has insufficient lateral stiffness within its supporting fork, e.g. a motorcycle front wheel with weak lower fork legs, or legs in which differential suspension movement is possible. The lateral stiffness used by Pacejka carries both wheel and fork, instead of the relative movement of wheel and fork. In his model, as with that of Temple, the only way in which Bc can lead 0, is for the tyre to deflect; hence the need to have a deformable tyre in all explanations of instability. By way of simple illustration, consider Fig. 5 with springs of stiffness k on the wheel axle, so that lateral movement is sprung rather than free. For co-ordinates choose Bu, and the wheel deflection 6 from the centre of the fork. Let M , = wheel mass, I , = wheel inertia about steering axis, Z, = inertia of supporting fork about this axis. The equations of motion are then, IJW = tk6 ZWB,+tMJ = -tk6-pRtr These are best rewritten as kt B , = - - s 1, and solved numerically, as in Fig. 7, for typical vehicle magnitudes. The unstable oscillation is quite clear" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001604_robot.1987.1087799-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001604_robot.1987.1087799-Figure4-1.png", "caption": "Figure 4", "texts": [ " This assumption is made for the sake of convenience and does not imply that the joints are capable of moving freely in both the CW and CCW directions. In fact, to go from -180\" (resp. 180') to 180\" (resp. - 180') the joint has to rotate counter-clockwise (resp. clockwise) by 360\". We shall first show that ~ A J must be a simple curve. Let q\" = ( q ; , q i ) be an arbitrary point on ~ A J and suppose that the manipulator is a t configuration q'. Then the second link is in contact with A. As shown in Figure 4, there are only six different ways in which the second link can make contact with A. Consider the case (a) in Figure 4. In order for the second link to maintain contact with A we can either (i) rotate joint 1 counter-clockwise and joint 2 clockwise, i.e. (J1 , Jz ) = (CCW, CW) for short, or (ii) ( J l , J z ) = (CW, CCW). Similarly, for the remaining five cases we have: case (b) I (CW, CCW) or (CCW,CW) (Jl, J2) I ::: 1 .(CCW,CW) or (CW,CW) 1 (d) (CCW,CCW) or (CW,CCW) (f) (CCW, CCW) or (CCW, CW) (CW,CCW) or (CW,CW) Hence, only two kinds of motion are possible in order for the second link to maintain contact with A" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002208_epe.2007.4417657-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002208_epe.2007.4417657-Figure1-1.png", "caption": "Fig. 1: Practical considerations regarding torque-angle relation in PMSM", "texts": [ " The incremental encoders can provide absolute position information through the use of so-called Z pulse which is generated only once per revolution, and due to this reason it is practically impossible to start the PMSM motor using this signal. In the present paper, a new method that use the information read from an incremental encoder is presented, which can work well even when a hard stop and/or constant load torque is being applied at the motor shaft In order to explain the relation between the electromagnetic torque and the rotor position, let us consider a simple PMSM machine, with one stator pair of poles (A, B, and C phases) and a non-salient pole type of rotor. Fig. 1a, shows the angular relation between the rotor position of the PMSM and the well known stator coordinates (usually the stator A-phase is taken as reference point in vector control). The D-Q and d-q coordinates are the stationary and rotating reference frame respectively. The mathematical model of a PMSM at standstill expressed in stationary reference frame is given by (1): [ ]D D D Q Q Q v i i = R + s L v i i [ ] cos 2 sin 2 sin 2 cos 2 1 2 e 2 e 2 e 1 2 e L L L L L L L \u03b8 \u03b8 \u03b8 \u03b8 + = + ( ) ( )d q d q1 2L = L +L /2, L = L -L /2 (1) (2) (3) Where vD , vQ are D- respectively Q- axis armature voltage, iD , iQ are D- respectively Q- axis armature current, R is armature resistance, \u03b8e rotational electrical angle of the d-axis and s is the differential operator. The rotor dynamic is described in (4): m e L mJ\u03b8 = T - T - B\u03b8 e m\u03b8 = p\u03b8 e Q eT = Ki sin\u03b8 (4) (5) (6) Where J, B, K are the moment of inertia, viscous damping coefficient and torque constant, respectively. Te, TL, p, \u03b8m and \u03b2 represents electromagnetic generated torque, load torque, number of pole pairs, angular position of motor shaft and motor internal angle. Fig. 1b, shows the situation when the motor shaft is preloaded with a constant load represented here by TL. Suppose that the motor develops an electromagnetic torque Te, equal to the load torque TL. As the angular characteristic Te = f(\u03b2) shows (Fig. 1b), there are then two possible values of \u03b2: \u03b20 \u2013 a value smaller than \u03c0/2 (the internal angle corresponding to maximal electromagnetic torque) and the complementary value, bigger than \u03c0/2. From the static stability criteria, the motor operation is stable if \u03b20 < \u03c0/2. When this condition is fulfilled, the rotor will start braking when the load torque increases suddenly. The internal angle will therefore increase and the electromagnetic torque will became larger, according to (6), until it equals the new load torque" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000687_cl.2005.1682-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000687_cl.2005.1682-Figure2-1.png", "caption": "Figure 2. Equivalent electric network of the electrochemical interface. Rs, electrolyte resistance; Cdl, double-layer capacitance; Ret, electron-transfer resistance; Zw, Warburg impedance.", "texts": [ " Next the electrode was incubated in 10mg/mL BSA for 1 h at 37 C to block non-specific site. The finished electrode was stored at 4 C when not in use. The schematic diagram of the biosensor and the structure of the electrode coating are shown in Figure 1. It is well known that electrochemical impedance spectroscopy (EIS) is an effective tool for studying the interface properties of surface-modified electrodes. The typical electrochemical interface can be represented as an electrical circuit as shown in Figure 2. Ret, which equals the semicircle diameter at higher frequencies in the Nyquist plot of impedance spectroscopy, controls the interfacial electron-transfer rate of the redox probe between the solution and the electrode. Thus, Ret can be used to describe the interface properties of the electrode. Its value varies when different substances are adsorbed onto the electrode surface. Figures 3 and 4 show the dependence of impedance spectroscopy with the electro-deposition time and absorption time. Figure 3 reveals a gradual increase of covering and Ret" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002893_dscc2008-2112-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002893_dscc2008-2112-Figure6-1.png", "caption": "Figure 6: Fish\u2013shaped part three\u2013dimensional scan using VPFRC with powder flow rate reference directly calculated from motion system speed.", "texts": [ " The reference powder flow rate and its first and second derivatives are shown in Figure 4 for one layer. The derivative peaks are due to acceleration and deceleration at the sharp corners and when transitioning from a linear to the circular segment and from the circular segment to a linear segment. The tracking performance, control signal, and disturbance estimate are shown in Figure 5. The average powder flow rate error absolute value and its standard deviation are 4.77\u00b710\u20132 g/min and 0.281 g/min, respectively. The deposition results are shown in Figure 6. Using the directly computed powder flow rate reference, the VPFRC greatly reduces the height variations for the points where the motion system transitions back and forth from linear and circular segments, and throughout the circular segment itself. The controller performance is good at these locations since the speed derivatives are relatively low. However, the height variation at the sharp corners is not improved due to the bandwidth limitations of the powder feeder system and the fact that the process dynamics significantly change when the motion system slows down and the laser delivers more energy" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002564_s10015-008-0560-2-Figure11-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002564_s10015-008-0560-2-Figure11-1.png", "caption": "Fig. 11. Mechanism of manipulator", "texts": [ " In Figure 8 and 9, it is shown that torques increase at 0.6\u223c0.7 s, and the energy consumption of the motors increase too. 4 Experimental results In this section, the results of fundamental experiment are shown to examine the effectiveness of modeling for the simulations. Under the condition that the torque of joint 3 is zero, the angular acceleration of joint 1 and 2 are given in iterative dynamic programming method, and the acceleration of link 3 and torques of joint 1, 2 are calculated simultaneously. Figure 11 shows a mechanism of manipulator whose mechanism is a closed type. The parameters of the system are shown in Table 2. The motors 1 and 2 (rated 24 V, 60 W) are on the frame, and sampling time of the control is 0.002 s. The feedback gain for angular displacement is 50 (V/rad), the feedback gain for angular velocity is 0.5 V s/rad. Figure 12 shows the experimental response, under the condition that initial position is (\u03b81i = \u03b82i = \u2212\u03c0/4, \u03b83i = 0), the fi nal position is (\u03b81f = \u03c0/2, \u03b82f = \u2212 \u03c0/4, \u03b83f = \u03c0/4), and the working time is 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002026_icma.2007.4303538-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002026_icma.2007.4303538-Figure7-1.png", "caption": "Figure 7. Actuator layout", "texts": [ " 5, when the internal bladder is pressurized, the highly pressurized air pushes against its inner surface and against the external shell, tending to increase its volume. Due to the non-extensibility of the threads in the braided mesh shell, the actuator shortens according to its volume increase and/or produces a load if it is coupled to a mechanical load. About 35% contraction can be expected with no load, and more than 20% for a load of 20 kg. Referring to muscle of human used for walking shown in Fig.6, actuator layout for the active walker is decided as shown in Fig.7. Table 1 describes number of actuator and length. We have decided actuator layout by empirical method and/or trial and error method, and are not sure that it is optimum or not. It is arguable but as the first step of this research, since the hart walker is employed all over the world so far and just by using simple attachment frame, McKibben artificial muscles are possible to mount to the hart walker, we apply this layout. Length(mm) 280 300 330 200 In case of HW, after attaching orthosis to the user, helper has to lift him/her for mounting to the stem" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001056_11429555_5-Figure5.4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001056_11429555_5-Figure5.4-1.png", "caption": "Fig. 5.4. 2D example construction of regions Wn(G, ). (A) A frictionless example grasp G consisting of five contact wrenches. (B) CHorig is the convex hull of these contact wrenches. Halfspace boundaries are also shown. (C) CHeps is constructed from CHorig by moving each halfspace boundary h to distance h from the origin. The region of acceptable wrenches w\u0302(cn) is an intersection of exterior halfspaces of CHeps. (D) Each such region can be mapped to a set of contact points on an object (Section 5.3.3) to obtain target contact regions Wn(G, ). Grasps in the family W (G, ) have one contact in each of these regions.", "texts": [ " The kinematics of the mechanism actually determine whether force closure and other properties of a grasp are achievable [29]. An overview of grasping research, including work in grasp synthesis, can be found in Bicchi [28]. The goal of our grasp synthesis algorithm is to create a family of grasps such that all grasps in the family have certain properties, which are specified at design time by the user. The algorithm we use is deterministic (i.e., does not involve search) and it involves constructing this family around a given example grasp. Figure 5.4 provides a 2D illustration of the construction process. Suppose we are given an example grasp G having N contacts g1, . . . , gN . Wrenches available at each contact are represented as a linear combination of L extremes. For example, a hard finger contact with friction may be represented using L samples bounding the friction cone at that contact point. The example, then, can be represented as the collection of NL extreme wrenches w\u0302l(gn). G = {w\u03021(g1), . . . , w\u0302L(gN)} (5.1) We assume that the w\u0302l(gn) span R6, although they may not positively span R6", ", halfspaces indexed by \u03c1n,l) after those halfspaces have been adjusted along their normals to distances h: some unit wrench available at contact cn must fall within this intersection. Sections 5.3.1 and 5.3.2 show how this particular definition makes it possible to control closure and quality properties of all grasps in W (G, ). In the frictionless case, L = 1 and subscript l is not needed, resulting in the following expression for contact region Wn, which is much cleaner and is illustrated in Figure 5.4D. Wn(G, ) = {cn : w\u0302(cn)c\u0307n\u0302h h \u2200h \u2208 \u03c1n} (frictionless case only) (5.8) Given this construction technique, what can we say about grasps in W? In [223], we show the following: Proposition 1. Suppose we are given grasp C having contacts {c1, . . . , cN } and grasp family W (G, ) constructed as in Equation 5.4. From Equation 5.2, CHorig(G) is the convex hull of the unit wrench extremes of G: CHorig(G) = ConvexHull{w\u03021(g1), . . . , w\u0302L(gN)} = [n\u03021 d1] T , . . . , [n\u0302H dH ]T (5.9) Define CHnew(C) as the convex hull of the unit wrench extremes of C: CHnew(C) = ConvexHull {w\u03021(c1), . . . , w\u0302L(cN )} (5.10) Let CHeps(G, ) be the intersection of all halfspaces having normals n\u0302h and distances h: CHeps(G, ) = { [n\u03021 1] T , . . . , [n\u0302H H ]T } (5.11) Then if grasp C is in grasp family W (G, ), CHnew(C) contains CHeps(G, ): C \u2208 W (G, ) \u2212\u2192 CHnew(C) \u2287 CHeps(G, ) (5.12) This Proposition describes the volume CHeps(G, ) in R6 (possibly empty) that is contained within the convex hull of contact wrenches of any grasp in the family W (G, ). This volume is shown for a 2D example in Figure 5.4C. Because CHeps in Equation 5.11 is determined by the values of h, properties of grasps in W (G, ) can be controlled through careful selection of these parameters. Grasps in W will be force-closure if the convex hull of the contact wrenches available from any grasp in W contains the origin in its interior. By Proposition 1, it is sufficient that CHeps in Equation 5.11 contain the origin in its interior. This goal is achieved by setting the following constraint: h > 0 h = 1, . . . , H (5.13) Any small number can be used for all h to ensure that force closure is possible for all grasps in W (G, )", "6, the problem of determining whether cn \u2208 Wn can be specified as follows: find Lx1 vector \u03b1l and parameter \u03bdn,l to maximize \u03bdn,l such that (5.19) (Y (cn)\u03b1l)c\u0307n\u0302h h\u03bdn,l \u2200h \u2208 \u03c1n,l (5.20) \u03b1l 0 (5.21) ||\u03b1l||L1 = 1 (5.22) Then from Equation 5.5: cn \u2208 Wn \u21d0\u21d2 L min l=1 \u03bdn,l 1 (5.23) This problem description states that there must be some unit contact wrench available at cn (i.e., some valid value for \u03b1l) such that all halfspace constraints are met or exceeded by this wrench (i.e., \u03bdn,l 1 for all l = 1, . . . L). In the frictionless case, cn \u2208 Wn can be determined more easily. From Equation 5.8 and Figure 5.4: cn \u2208 Wn \u21d0\u21d2 w\u0302(cn)c\u0307nh h \u2200h \u2208 \u03c1n (5.24) In this case, at each sample point on the frictionless surface the contact wrench w\u0302(cn) is formed and tested for inclusion in Wn by checking just a few 6D dot products. Referring back to Figure 5.3, one feature of this technique is that properties of the example grasp are retained. In some circumstances, such as that shown in the figure, some of these properties may be incidental and not especially desirable. Figure 5.5 shows the difficulty. Figure 5.5A shows the \u03c4 = 0 slice of the 3D force / torque space for the example in Figure 5", "5C and D show the differences after convex hull CHeps (Equation 5.11) is formed using h from Equation 5.15, with task wrenches in the fy and \u03c4 directions ([fx, fy, \u03c4 ]T = [0,\u00b11,\u00b10.3]T ). The symmetric example from Figure 5.3(Right) aligns much more closely to the set of forces and torques required for this task. As a result, given our constructive technique for grasp synthesis, the contact regions in Figure 5.3(Right) are larger and more symmetrical. It is also possible to see this phenomenon by examining Figure 5.4C. With reference to this figure, a poor alignment of halfspaces to task wrenches may produce CHeps with large interior and result in small regions w\u0302(ci), while a good alignment may produce smaller CHeps and tend toward larger w\u0302(ci). One solution to this problem is to adjust the frictionless contact wrenches of the example grasp so that they are aligned, where appropriate, with the task space or its nullspace. The assumption we are making is that if a wrench is nearly aligned with one of these spaces, then differences from exact alignment are incidental and do not reflect an important property of the example grasp" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003454_978-1-4614-0347-0_1-Figure12-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003454_978-1-4614-0347-0_1-Figure12-1.png", "caption": "Figure 12. In an exocytotic event based the mechanism kiss and run with a flickering fusion pore, multiple transient fusion pores are formed. Each fusion pore formation and neurotransmitter release event is interrupted by a temporary retrieval of the partially emptied vesicle.", "texts": [ " A special case of the kiss-and-run mechanism, a flickering fusion pore, has been observed in connection with DA exocytosis from small synaptic vesicles of rat ventral midbrain neurons.138 This mechanism further increases the efficiency of utilizing vesicles as well as controlling the amount of released neurotransmitters. The size of initial fusion pores of small vesicles undergoing exocytosis by the kiss-and-run mechanism has been determined to be of the order of magnitude of one ion channel as well as that the opening has to be a protein structure that is similar to an ion channel.119,139 Figure 12 shows a schematic overview of exocytosis based on the kiss-and-run mechanism with flickering fusion pore. In the case of the all-or-none mechanism, the number of released neurotransmitter molecules per quantum is equal to that of the vesicular content, which can be several millions in large vesicles. In small neuronal vesicles the number of neurotransmitter molecules can be as low as 3,000\u201330,000.127 If exocytosis follows the kissand-run mechanism with a flickering fusion pore, the vesicular content is not fully released in one exocytotic event; instead, only 25\u201330% is released" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002320_978-3-540-77457-0_2-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002320_978-3-540-77457-0_2-Figure5-1.png", "caption": "Fig. 5. Tangential deformation in the parallel distributed model", "texts": [ " Integrating potential energies caused by the perpendicular deformation of individual elastic elements in a parallel distributed model yields the potential energy of the fingertip as follows: Uperp(d, \u03b8p) = \u03c0Ed3 3 cos2 \u03b8p . (3) Note that the potential energy depends not only on the maximum displacement d but also on the relative orientation \u03b8p. As described in Section 2, tangential deformation should be introduced into the parallel distributed model so that a pinched object can rotate under the application of external force. Figure 5 shows the model of tangential deformation. Assume that the fingertip makes contact with the rigid object without tangential deformation of the fingertip, as illustrated in Figure 5-(a). In the parallel distributed model, point Qk on the fingertip surface moves to Pk, shrinking the elastic element of natural length QkRk to PkRk. Assuming that the rigid object moves tangentially by displacement dt as shown in Figure 5-(b), the point Pk moves to P\u2032 k. Then, the elastic element has tangential deformation determined by PkP\u2032 k. Given the position and the orientation of an object, we can calculate the perpendicular deformation QkPk and tangential deformation PkP\u2032 k of each elastic element. The tangential deformation determines the tangential force generated by the element. For the sake of simplicity, we assume that Young\u2019s modulus E characterizes the linear relationship between the tangential force and the tangential deformation" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003282_978-0-387-74244-1_11-Figure11.16-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003282_978-0-387-74244-1_11-Figure11.16-1.png", "caption": "FIGURE 11.16. A passing maneuver.", "texts": [ " Passing and lane-change maneuvers are two other standard tests to examine a vehicle\u2019s dynamic responses. Passing can be expressed by a halfsine or a sine-squared function for steering input. Two examples of such functions are \u03b4 (t) = \u23a7\u23a8\u23a9 \u03b40 sin\u03c9t t1 < t < \u03c0 \u03c9 0 \u03c0 \u03c9 < t < t1 rad (11.276) \u03b4 (t) = \u23a7\u23a8\u23a9 \u03b40 sin 2 \u03c9t t1 < t < \u03c0 \u03c9 0 \u03c0 \u03c9 < t < t1 rad (11.277) \u03c9 = \u03c0L vx . (11.278) where L is the moving length during the passing and vx is the forward speed of the vehicle. The path of a passing car would be similar to Figure 11.16. Let\u2019s examine a vehicle with the characteristics given in (11.226)-(11.229) and a change in half-sine steering input \u03b4 (t). \u03b4 (t) = \u23a7\u23a8\u23a9 0.2 sin \u03c0L vx t 0 < t < vx L 0 vx L < t < 0 rad (11.279) L = 100m (11.280) vx = 30m/ s. (11.281) The equations of motion for zero initial conditions are as given in (11.271)- (11.274). Figures 11.17 to 11.20 show the time responses of the vehicle for the steering function (11.279). Example 420 F Passing with a sine-square steer function. A good driver should change the steer angle as smoothly as possible to minimize undesired roll angle and roll fluctuation" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002926_pime_proc_1970_185_113_02-Figure17-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002926_pime_proc_1970_185_113_02-Figure17-1.png", "caption": "Fig. 17. Resultant ground force and its components", "texts": [ " The comparison between cross ply and radial ply illustrates the main value of slow speed dry testing in that differences in construction are displayed which demonstrate the handling and stability characteristics of tyres. However, in wet conditions, the value of machine testing is in developing new tread patterns to obtain maximum wet grip properties at high speeds. The progress made over the last 15 years in this field is shown in Fig. 16. The driven tyre A driven tyre is one which is subjected to braking and driving torques. The forces acting on a driven tyre are compared with those on a rolling tyre in Fig. 17. Proc lnstn Mech Engrs 1970-71 Vol 185 74/71 at UNIV OF VIRGINIA on June 5, 2016pme.sagepub.comDownloaded from ENGINEERING ASPECTS OF TYRE TESTING 7 0 - 8 0 - 9 0 - 00- 110- 2 0 - 30- 4 0 - 6 0 - 80- 1011 + 5 u a 4 :: 0 w 2 e vl 2 D i r e c t i o n o f m o t i o n B r a k e d wheel !-(I- O f f s e t 7 D i r e c t i o n o f mot ion - Resul tant Free r o l l i e g wheel , mrce _ _ , - - h e i g h t b on a rolling tyre with applied torques The vertical reaction of the ground to the tyre load is not through the axle but a little forward of it", " In fact, with rubber, the concept of a static friction force giving way to a lower sliding value does not operate; movement always occurs before the maximum friction is obtained. The practical result of this is that when a tyre is driven, its angular velocity or rev/min, for a given forward speed is greater than when rolling. In other words the tyre creeps or slips. Fig. 18 shows the variation of tyre slip with braking and driving coefficients, which represent the ratio of braking or driving efforts to tyre radial load, and illustrates the limiting conditions of sliding and wheelspin respectively. The effect of tyre slip is also illustrated in Fig. 17 by the distance 1 covered in one tyre revolution. Tyre power loss Tyre power loss is the difference between the power used to drive or brake a tyre and the power transmitted by the tyre to a surface. Tractive efficiency is defined as the ratio of output power to input power expressed as a percentage. In practice, tyre tractive efficiencies range between 75 per cent for thick treaded tyres to 98 per cent for radial ply car tyres. From reference (6), considering the work done by a driven tyre, the formula D = C/r -F is obtained, where Tabulated v a l u e I , I ; S l i d e 1 L D r a g c o e f f i c i e n t - - 4 I I < W h e e l s p i n '$ L-Slip c , D r i v e l r i v e c o e f f i c i e n t - 1", " Engagement and tyre noise, in our opinion, can be estimated on a machine having a drum with an inner rolling surface (rolling along the internal surface of a cylinder) imitating various hard road surfaces : cement concrete, asphalt concrete, stone and others with varying degrees of roughness. Tyre tests on machine using such surfaces will enable many important and interesting characteristics of tyre engagement and noise when rolling to be established. (15) FrictioEaf wear and rear of rubber 1964, 200 (\u2018Khimia\u2019, Moscow). B. D. A. Phillips Member I would like to query one or two points of detail in the paper. In Fig. 17, the offset is shown to move towards the front of the tyre under the action of a braking force, and to the rear under the action of a driving force, whereas it moves to the front under the action of a driving force and to the rear under the action of a braking force. In Fig. 18, the intersection of the tractive coefficient-slip curve is shown to be on the driving side of the free rolling axis whereas it should be on the braking side, this intersection being the rolling resistance of the tyre. I assume that a tyre is defined as having zero slip when there is no torque applied to the wheel axle", " In reply to D. Bastow, the range of maximum selfaligning torques for different size tyres varies in the dry from 2 kgf m (14 lb ft) for 13 in diameter tyres to well over 7 kgf m (501b ft) for 15in tyres. In the wet the available self-aligning torque is very much reduced. In reply to D. G. Powell, the effect of pattern depth on wet grip is similar to Fig. 16 where the year figures are replaced by pattern depths, the latest year being the deepest tread pattern. G. F. Morton (19) shows this relation in his Fig. 17. Currently, I am correlating pattern depth effects on the wet cornering force machine with glass plate photographs. Figs 17 and 18 were taken from reference (6) and are both theoretical diagrams. They are both thoroughly explained by V. E. Gough in his communication. The drag and drive coefficients of Fig. 18 are based on axle torque C, whereas the braking and driving coefficients of Fig. 19 are based on tractive effort F. Hence the free rolling case of Fig. 19 is on the braking side of the vertical axis, at a coefficient of 0", " In fact, one of the purposes of the paper is to provide a broad look at tyre testing covering his division (2) without involving the many formulae that would appear in division (3). REFERENCES (IS) AMES, W. F. \u2018Waves in tires\u2019, Text . Res. J. 1970 (June), 498. (19) MORTON, G. F. \u2018Factors influencing tyre development\u2019, J. I.R.I. 1970 (August) 4, 145. Corrigenda Page 1006 Last paragraph should read: For meaningful test results and true comparisons of types of tyre, machine conditions must be correctly set. Page 1009 Fig. 13. Caption should read: Effect of speed on rolling resistance (machine test temperatures). Page 1011 Fig. 17. Driving wheel illustration. Drum Force F2 should be Driving Force F2. Proc lnstn Mech Engrs 1970-71 Vol 185 74/71 at UNIV OF VIRGINIA on June 5, 2016pme.sagepub.comDownloaded from" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002412_icems.2009.5382984-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002412_icems.2009.5382984-Figure1-1.png", "caption": "Fig. 1. Overview of permanent magnet motor", "texts": [ " The electromagnetic vibrations under the running conditions are obtained using the accurate electromagnetic force distribution. Third, since it is difficult to accurately model the coil, we investigated the influences modeling the coil on the electromagnetic vibration. At the last, the modeling method in order to estimate the electromagnetic vibration of stator case is clarified by comparing the measurements with the analysis results. A permanent magnet motor was manufactured as a benchmark machine to test our magnetic filed simulator. The overview of the motor is shown in Fig. 1. The specifications are shown in Table I. Modeling method of Vibration Analysis Model for Permanent Magnet Motor Using Finite Element Analysis Tetsuya Hattori*, Katsuyuki Narita*, Takashi Yamada*, Yoshiyuki Sakashita*, Kiyotaka Hanaoka**, Kan Akatsu** *Electromagnetic Engineering Department, Engineering Technology Division, JSOL Corporation, Japan **Department of Electrical Engineering, Shibaura Institute of Technology, Japan The natural frequencies are measured by the impulse responses by using an impact hammer" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001961_s11044-007-9054-6-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001961_s11044-007-9054-6-Figure1-1.png", "caption": "Fig. 1 A prototype of 2SPS+UPR parallel manipulator", "texts": [ " In order to remove the redundant motion and enlarge the workspace, this paper focuses on a 2SPS+UPR parallel manipulator and its kinematics, workspace, and active/constrained forces. 2.1 The 2SPS+UPR parallel manipulator and its dofs A 2SPS+UPR parallel manipulator is composed of a (moving) platform m, a (fixed) base B , and two SPS-type active legs ri (i = 1,3) with the linear actuators, and one UPR-type constrained active leg r2 with a linear actuator and a rotational actuator (motor), where, m is an equilateral ternary link b1b2b3 with 3 sides li = l,3 vertices bi , and a center point o (see Fig. 1). B is an equilateral ternary link B1B2B3 with 3 sides Li = L,3 vertices Bi , and a center point O . Let {m} be a coordinate system o-xyz fixed on m at o, {B} be a coordinate system O-XYZ fixed on B at O . Each of ri (i = 1,3) connects m to B by a spherical joint S at bi , an active leg ri with a prismatic joint P , and a spherical joint S at Bi . The UPR-type constrained active leg r2 connects m to B by a revolute joint R3 attached to m at a2, a constrained active leg r2 with a prismatic joint P , and a universal joint U attached to B at B2" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003789_ijtc2011-61146-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003789_ijtc2011-61146-Figure7-1.png", "caption": "Figure 7: Photo of rear air thrust bearing and rear radial foil bearing (Design II) inside the rear bearing housing", "texts": [ " To implement both radial foil bearings and thrust air bearings, rotor was redesigned (Figure 5(b)) with a thrust disc with 12 mm diameter with higher bending critical speed above 1,000,000rpm. Monolithic stator housing was designed to house both the motor stator and front radial foil bearing. Rear part of the stator housing was designed to accommodate front air thrust bearing (Figure 6) and thrust spacer. Separate rear housing contains both rear air thrust bearing and rear radial foil bearing (Figure 7). The air thrust bearings have Rayleigh steps of 10 \u00b5m high, formed through selective electroless Ni coating. Nominal clearance of the thrust bearing is also 10 \u00b5m in one side (total 20 \u00b5m axial play of the thrust disc). The inner and outer diameters of the thrust bearing are 6.8 and 12mm, respectively. A dummy disc with 12mm OD was assembled on to the front end of the rotor to mimic an impeller and to balance the weight distribution along the axial direction. Figure 9 shows cross section of the assembled micro motor with foil bearings made of Design II elastic foundations, and Figure 9 shows the micro motor mounted on a pedestal for test" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003184_6.2008-6332-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003184_6.2008-6332-Figure1-1.png", "caption": "Fig. 1 The 65-cm Vertigo.", "texts": [ " The Mini-Vertigo is two times smaller than the Vertigo, thus the two aircraft present a workable experimental setup for the investigation of American Institute of Aeronautics and Astronautics 4 size effects on flight dynamics and designing control laws for VTOL MAVs. The approach utilized in this study includes the simulation of flight dynamics and closed-loop control design. In a series of flight tests of the two aircraft, telemetry data on control actuation, altitude, and attitude of the aircraft were collected and analyzed and used in the validation of predicted designs. II. Autonomous VTOL MAV Designs and Specifications In the present study, the Vertigo (Fig. 1) and the Mini-Vertigo (Fig. 2) MAVs were evaluated. Geometrical and mass data for the MAVs and their components are presented in Tables 1 and 2. Both aircarft are of a tail-sitter configuration capable of vertical take-off, hover, transition, and level flight. The 31-cm-wing-span Mini-Vertigo is two times smaller than 65-cm-wing-span Vertigo and, therefore, is better qualified as an MAV category airplane. The Vertigo was designed in the mini-UAV size range and is able to carry a payload up to 200 g, while the MiniVertigo has 50-g payload capability" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001575_fuzzy.2006.1681757-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001575_fuzzy.2006.1681757-Figure1-1.png", "caption": "Fig. 1 Dead-zone model.", "texts": [ " PROBLEM FORMULATION Consider a class of the following uncertain nonlinear system with an unknown dead-zone of the form ( ) ( ) ( ) ( )( ) 1 2 2 3 1 n x x x x x f f g D v t y x = = = + \u2206 + = x x x (1) or equivalently ( ) ( ) ( ) ( ) ( )( )nx f f g D v t y x = + \u2206 + = x x x (2) where [ ] ( )1 1 2, , , , , , TT n n nx x x x x x R\u2212 = = \u2208 x is the system state vector which is assumed to be available for measurement, v R\u2208 and y R\u2208 are the input and output of the system, respectively, f and g are unknown nonlinear functions and f\u2206 is the unknown uncertainty. Without loss of generality, it is assumed that the sign of ( )g x is positive, and ( ) ( )f h\u2206 \u2264x x , where ( )h x is an unknown continuous function and can be estimated by an adaptive law in the latter. ( )( ) :D v t R R\u2192 is the nonlinear input function containing a dead-zone. To clarify the dead-zone nonlinear input function ( )D \u22c5 , the dead-zone with input ( )v t and output ( )w t , as shown in Fig. 1, is described by ( ) ( )( ) ( )( ) ( ) ( ) ( )( ) ( ) for , 0 for , for , r r r l r l l l m v t b v t b w t D v t b v t b m v t b v t b \u2212 \u2265 = = < < \u2212 \u2264 (3) where 0rb > , 0lb < and 0rm > , 0lm > are parameters and slopes of the dead-zone, respectively. In order to investigate the key features of the dead-zone in the control problems, we have the following assumptions: Assumption 1: The dead-zone output ( )w t is not available. Assumption 2: The dead-zone slopes are same, i.e. r lm m m= = . Assumption 3: There exist known constants minrb , maxrb , minlb , maxlb , minm , maxm such that the unknown dead-zone parameters rb , lb , and m are bounded, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000390_05698190500313478-Figure13-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000390_05698190500313478-Figure13-1.png", "caption": "Fig. 13\u2014Ball-raceway contact and angle defining direction of ball about its center.", "texts": [ " Therefore, mbx of test bearing A is larger analytically than that of test bearing B under certain values of \u00b5 and p. This qualitatively matches the measured results. In this section, an equation is derived for the running torque of test bearings A and B caused by friction between the balls and D ow nl oa de d by [ U ni ve rs ity o f O ta go ] at 0 4: 56 0 6 O ct ob er 2 01 4 the polymer lubricant. The ball-raceway contact and the angle \u03b2, defining the direction of the ball about its center in test bearings A and B under an axial load, are shown in Fig. 13. In Fig. 13, o is the ball center, Y is the ball axis of rotation, and the X axis is perpendicular to the Y axis. Although the Z axis is not seen in Fig. 13, it is the circumferential direction of the bearing. RLi is the inner raceway groove curvature radius, R is the distance from the bearing axis to the ball-inner raceway contact point, and \u03b1 is the contact angle. In operation, the balls roll about their centers in test bearings A and B, and a slip between the balls and the polymer lubricant occurs. The frictional moment about the Y axis due to the slip of an element of area ds between the ball and the polymer lubricant dmbY is given by dmbY = \u00b5p (d/2) \u221a (d/2)2 \u2212 Y2\u221a (d/2)2 \u2212 Y2 \u2212 Z2 dYdZ [10] From Eq", " The angle \u03b2 defining the ball rotation axis is given by \u03b2 = \u03b1 + tan\u22121 ( d sin \u03b1 2R ) [15] where R is the distance from the bearing axis to the ball-inner raceway contact point and is given by R = RLi + RRi (1 \u2212 cos \u03b1) [16] In Eq. [16], RRi is the inner groove radius. The contact angle \u03b1 can be calculated based on the load distribution in the ball bearing (Bra\u0308ndlein, et al. (15)). The calculated contact angles \u03b1 for the axial loads in the experiments are 13.0-14.1\u25e6. Based on Eq. [15], the angles \u03b2 (defining the ball rotation axis) are 16.0 to 17.7\u25e6. From the geometry (see Fig. 13) of test bearings, the running torque Mb caused by the friction between the balls and the polymer lubricant can be described as Mb = 2nmbY R/d = 2cnmbx R/d [17] where n is the number of balls; c is the ratio of mbY to mbx and is given by c = mbY/mbx [18] The calculated values of c are not significantly affected by axial loads. The values for test bearings A and B are 0.88 and 0.84, respectively. From Eq. [17], the calculated values of the running torque Mb for test bearings A and B are 0.080 and 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003789_ijtc2011-61146-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003789_ijtc2011-61146-Figure6-1.png", "caption": "Figure 6: Photo of front air thrust bearing inside the stator housing", "texts": [ " The original PM rotor with two ball bearings had the first critical speed of 198,840 rpm according to [8] (including an axial turbine). To implement both radial foil bearings and thrust air bearings, rotor was redesigned (Figure 5(b)) with a thrust disc with 12 mm diameter with higher bending critical speed above 1,000,000rpm. Monolithic stator housing was designed to house both the motor stator and front radial foil bearing. Rear part of the stator housing was designed to accommodate front air thrust bearing (Figure 6) and thrust spacer. Separate rear housing contains both rear air thrust bearing and rear radial foil bearing (Figure 7). The air thrust bearings have Rayleigh steps of 10 \u00b5m high, formed through selective electroless Ni coating. Nominal clearance of the thrust bearing is also 10 \u00b5m in one side (total 20 \u00b5m axial play of the thrust disc). The inner and outer diameters of the thrust bearing are 6.8 and 12mm, respectively. A dummy disc with 12mm OD was assembled on to the front end of the rotor to mimic an impeller and to balance the weight distribution along the axial direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003372_978-3-642-25486-4_28-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003372_978-3-642-25486-4_28-Figure3-1.png", "caption": "Fig. 3. Geometry of Parallel Indexing Cam Mechanism", "texts": [ " Tq )100000( 1 11 1 = \u00d7\u2212 = \u03c9 \u03c9 \u03be ; (2) Ta q )10000( 2 22 2 \u2212= \u00d7\u2212 = \u03c9 \u03c9 \u03be . 3) Where 1\u03c9 , 2\u03c9 are angular velocity of turret and cam; 1q , 2q are coordinates of points on cam axis and turret axis. Tq )000(1 = ; Taq )00(2 = According to meshing relationship between turret center and cam profile, the theoretic profile of indexing cam ),( 21 \u03b8\u03b8fA is )0( 1000 sincossinsinsincos 00cossin cossincossincoscos )0()(),( 222121 11 222121 \u02c61\u02c6 21 1122 f fc A a aa AeeA \u2212 +\u2212\u2212\u2212 = = \u2212 \u03b8\u03b8\u03b8\u03b8\u03b8\u03b8 \u03b8\u03b8 \u03b8\u03b8\u03b8\u03b8\u03b8\u03b8 \u03b8\u03b8 \u03b8\u03be\u03b8\u03be (4) Then the geometry model can be derived as Fig. 3. In one automated machinery indexing cam system is combined by motor, indexing cam, roller and turret and rotary table as Fig. 4. Contact between cam and rollers of turret is the key factor affecting the system performance and is dealt by Unilateral contact model. Unilateral contact model denotes a mechanical constraint which prevents penetration between two bodies. For parallel indexing cam mechanism, clearance between rollers and cam profile can be divided into three situations: Fig. 5 (a) presents that roller penetrated into cam profile, contact force is defined by complaint distance and curvature radius, roller radio, material property, oil film thick between roller and cam; (b) presents roller contract with roller justly" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003473_19346182.2010.540468-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003473_19346182.2010.540468-Figure1-1.png", "caption": "Figure 1. Geometry of the bounce of a tennis ball incident at angle u1 on a tennis court. F is the friction force acting at the bottom of the ball andN is the normal reaction force, which acts at a distance D ahead of the centre of mass.", "texts": [ " In order to measure the COF, the ball must be incident on the court surface at a sufficiently lowglancing angle that it slides on the court throughout the whole bounce. At higher angles of incidence, the ball will grip the court during the bounce process, in which case sliding friction gives way to static friction. The time-average value of the COF during the bounce will then be lower than that due to sliding friction. The effect is larger than one might expect since the static friction force reverses direction during the bounce (Cross, 2002). The geometry of the situation is shown in Figure 1. A ball of mass m, radius R is incident without spin at speed v1 and at an angle u1 to the court surface. The ball bounces at speed v2, angle u2 with angular velocity v2. If F is the horizontal friction force (acting in the negative x direction) and N is the normal reaction force (acting in the positive y direction) then F \u00bc -mdvx/dt andN \u00bc mdvy/dtwhere vx and vy are the horizontal and vertical components of the velocity of the centre of mass of the ball. N does not necessarily act through the ball centre of mass. In Figure 1, it is indicated that N can act along a line at a distance D ahead of the centre of mass. Such an effect can be expected because the front edge of the ball rotates into the court surface while the back edge rotates out of the surface, generating an asymmetry in the distribution of the normal reaction force over the contact region of the ball. A similar effect arises with a vehicle when the brakes are applied, the front end rotating downwards and increasing the normal reaction force on the front wheels" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000342_2004-01-1448-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000342_2004-01-1448-Figure4-1.png", "caption": "Figure 4. a) Schematic of a Metal Cam Roller System", "texts": [ " This will further stress metallic components leading to more applications requiring the use of silicon nitride components. Additional interest in evaluating the use of silicon nitride components in various other applications has started. To understand the benefits of using silicon nitride for cam roller applications, one must understand the application and more importantly understand why metal components fail. The metal cam roller system is composed of, 1) a steel cam roller , 2) a steel cam lobe and 3) a bronze pin (Figure 4A). The performance of this system has been extensively studied1. 1. The wear and the high friction at the pin to metal roller interface prevents free rotation of the metal roller which in turn causes wear at both the ID and the OD of the metal rollers and at the cam lobe. 2. Cam roller sliding caused by high contact stresses and frictional \u201cstick-slip\u201d conditions causes the area of maximum stress to move from below the surface toward the surface and may initiate contact fatigue failures on both the roller and the cam lobe (Figure 5)", " The bronze pin and the metal cam roller exhibit high friction with \u201cstick-slip\u201d under high levels of lubrication, and very high friction and severe \u201cstick-slip\u201d under starved lubrication conditions1. Due to this, typical metal cam rollers in high-pressure systems show the following: \u2022 Scuffing and pitting on the OD of the metal cam roller \u2022 Wear on the ID of the metal roller \u2022 Scuffing and wear of the OD of the pins The silicon nitride cam roller system is composed of, 1) silicon nitride cam roller and 2) steel cam lobe 3) steel pin (Figure 4B). The performance of this system has also been studied 1,12. 1. The steel pin is much harder than the corresponding bronze pin of the metal system, making it more wear resistant to dirty lubrication and other debris. 2. Silicon nitride and steel make an excellent tribological couple. Silicon nitride and steel exhibit low and constant friction at both flooded and starved lubrication conditions. 3. The wear of the ID of the silicon nitride rollers is lower than the corresponding metal rollers by approximately a factor of 3-10 (Figure 6)13,14", " The crown profile of a silicon nitride roller needs to be large enough to keep the line contact away from the edges of the ceramic roller. The contact stresses will generally be higher in a silicon nitride system due to a higher elastic modulus of the silicon nitride versus steel, for the same steel-to-steel profile, but an optimized profile for the ceramic can reduce this. The engine performance results provided by Ceradyne\u2019s customers is qualitative in nature and can be summarized as follows: \u2022 The wear of the ID of a silicon nitride roller is reduced by a factor of 10 relative to the ID of a metal roller (Figure 4). \u2022 The life of a cam lobe running against silicon nitride is longer than one running against a metal cam roller, even at higher contact stresses. This is supported by independent laboratory study13. \u2022 The reliability of the engines using silicon nitride components has increased. Warranty claims against the silicon nitride component systems have been significantly reduced or eliminated with the use of silicon nitride components. \u2022 The reliability of fuel pumps using silicon nitride has also increased and warranty claims have been eliminated" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000924_05698190500225284-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000924_05698190500225284-Figure1-1.png", "caption": "Fig. 1\u2014A radially loaded ball bearing.", "texts": [ " Nondimensional stiffness and damping coefficients of a single ball contact can be written in terms of dynamic pressure distribution as K\u0304 = K E\u2032 Rx = \u222b\u222b A p\u0304h dx\u0304dy\u0304 [7] C\u0304 = CU E\u2032 Rxr = \u222b\u222b A p\u0304 \u02d9\u0304h dx\u0304dy\u0304 D ow nl oa de d by [ U ni ve rs ity o f C on ne ct ic ut ] at 0 2: 52 0 8 O ct ob er 2 01 4 Overall stiffness and damping of a ball bearing is the combination of inner and outer race contact stiffness and damping of individual load-sharing balls. Stiffness and damping, of course, vary with contact geometry and load. A radially loaded ball bearing with radial clearance Pd is shown in Fig. 1(a); the inner ring makes contact under static and noload conditions. It is noticeable that the clearance at the load line is zero and increases with the angle \u03d5. Now application of load causes elastic deformation of balls over the arc 2\u03d5l , and there will be lubricant film to support this external load. Then total interference along the load line (\u03d5 = 0) is given by (Fig. 1(b)): \u03b40 = V0 c \u2212 h0 c where V0 c is the combined inner and outer race elastic deformation of the ball along the load line (\u03d5 = 0), and h0 c is the combined inner and outer race lubricant film thickness in the load line contact (\u03d5 = 0). Interference at any angular position from the load line can be represented in terms of total radial distance of inner ring or shaft from the concentric position \u03b4: \u03b4\u03d5 = ( \u03b4 cos \u03d5 \u2212 pd 2 ) where \u03b4 = \u03b40 + pd 2 [8] Using the Hertz elastic deformation and Hamrock and Dowson (16) film thickness empirical relations, one can write \u03b4\u03d5 = KHW2/3 \u03d5 \u2212 KEHLW\u22120" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002945_icsmc.2009.5345930-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002945_icsmc.2009.5345930-Figure2-1.png", "caption": "Figure 2. The fuzzy membership functions of ZO and NZ", "texts": [ " A fuzzy rule base is of the form If s is ZO, Then u is equu = (25) If s is NZ, Then u is heq uuu += (26) where ZO and NZ denote zero and nonzero fuzzy sets, respectively, and input variable s is given in (7). The control law of the fuzzy controller is )()( ])[()( ss uusus u NZZO heqNZeqZO \u03bc\u03bc \u03bc\u03bc + ++ = (27) where )(sZO\u03bc and )(sNZ\u03bc is the membership functions of fuzzy sets ZO and NZ, respectively. The membership functions of fuzzy sets ZO and NZ are selected to overlap and be symmetric to satisfy 1)()( =+ ss NZZO \u03bc\u03bc . If we choose the triangle membership functions as shown in Fig. 2 for the fuzzy sets ZO and NZ of s, the control law u will be continuously adjusted by the use of the fuzzy logic depending on \u201c ZO\u201d layer 1s . When holding the condition 1ss \u2265 , it can be seen that the control law is the same as the proposed SFCMAC. However, the amount of hitting control in region 1ss < is dominated by the grade of the membership function of NZ, that is, the hitting control could be attenuated by the grade of NZ. V. SIMULATION In this section, the proposed SMC-based FCMAC control system will be applied to control a Duffing forced oscillation system [14]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000209_68.5.441-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000209_68.5.441-Figure2-1.png", "caption": "FIG. 2. Typical phase portrait of i th bird\u2019s flight in a formation. The upper bound is given by (3.9) and the lower bounds #1 and #2 denote (3.14) and (3.15), respectively.", "texts": [ "14) The last term of the lower bound is a hyperbola in terms of ui , and its coefficient in braces is a positive constant. One can derive a tighter lower bound if u jF+ u j for all j . Replacing u jT O with u jF+ , one obtains another lower bound: \u03b1i u 2 i + \u03b2i Gii u \u22122 i \u2212 \u03c4\u0304i \u2212 \u03b4\u03c4i + {\u2211 j =i \u03b2 j G \u221e j i u\u22121 jF+ } u\u22121 i . (3.15) The acceleration u\u0307i is a single-valued explicit function with respect to ui , and hence the i th bird\u2019s phase portrait of formation flight must be located somewhere in between these bounds. Figure 2 shows a sample locus of the i th bird\u2019s flight as a solid line meandering in a region between the upper and lower bounds. In this figure the lower bounds #1 and #2 denote those in (3.14) and (3.15), respectively. Another critical velocity uiF\u2212 shown in Fig. 2 is defined as the intersection between the ui -axis and the lower bound #2: uiF\u2212 = \u22121 2 { \u221a p + v + \u221a p \u2212 v + 2q\u221a p + v } , (3.16) where v = a \u2212 bc 1 \u2212 c , a = 32pr + 9q2 24pr + 2p2 (\u22121 + d), b = 32pr + 9q2 24pr + 2p2 (\u22121 \u2212 d), c = { 1 + 96pr + 27q2 72pr + 27q2 \u2212 2p3 d }\u22121/3 , d = \u221a 1 \u2212 4 (12r + p2)(16r2 + 12p2r + 3pq2) (32pr + 9q2)2 , p = \u03c4\u0304i + \u03b4\u03c4i \u03b1i , q = \u2211 j =i \u03b2 j G\u221e j i u\u22121 j F+ \u03b1i , r = \u03b2 j Gii \u03b1i . Let us consider the global attractiveness of the state where u\u0307i = 0 and ui \u2208 (uiF\u2212 , uiF+)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001604_robot.1987.1087799-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001604_robot.1987.1087799-Figure2-1.png", "caption": "Figure 2: Obstaslca and Shadows.", "texts": [ " There are two major drawbacks to most of the existing work on work-path planning. First, in existing methods the forbidden space is defined as the space occupied by the obstacles and the free space is defined as the complement of the forbidden space. Using this definition of the free space, a work-path obtained can only guarantee that the end-effector of the manipulator is free of collision; the links of the manipulator may not be free of collision. As an example, consider the situation shown in Figure 2. The regions A1 and A2 are not part of the obstacle A, yet these two regions must be considered forbidden because they give rise to collisions between the second link and the obstacle A. Regions in the workspace with such a property are called shadows of obstacles [7:, The second drawback of existing work-path planning methods is the fact that optimality criterion of a work-path usually does not correspond to that of the free-space graph, and v ice ver sa . For example, suppose that an optimal work-path is defined to be one which has the shortest distance between the initial and final postures in the workspace", " As we mentioned earlier, a work-path which is free of collision does not guarantee that the links of the manipulator are also free of collision. Therefore, joint-space path planning must be treated as a separate problem, independent of the work-path planning problem. It appears that the first drawback of the work-path planning problem can be easily overcome by first finding all the shadows of the obstacles and then defining the forbidden space to be the union of the obstacles and their shadows. However, as can be seen from Figure 2 , an obstacle may give rise to disconnected shadows, and a shadow may not be connected to the associated obstacle. In fact, it is not difficult to imagine that, depending on the lengths of the links, the location, shape and size of an obstacle, there may be more than one shadow, or no disconnected shadows at all. As a result, it is very difficult to characterize shadows of obstacles. In contrast, it will be shown later in the paper that obstacles in the joint space can be characterized easily. The difficulty of defining an appropriate optimality criterion is not limited to work-path planning" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003094_iros.2009.5354487-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003094_iros.2009.5354487-Figure2-1.png", "caption": "Fig. 2. Rimless wheel model.", "texts": [ " We first formulate the energy-loss coefficient mathematically taking SLR into account as a function of four parameters; the half inter-leg angle, the location of center of mass, the mass ratio and the ratio of angular velocities. We numerically and analytically investigate how these parameters affect the value of energy-loss coefficient. Furthermore, we discuss the relation between the energy-loss coefficient and the eigenvalue of Jacobian matrix for Poincare\u0301 return map. This section outlines the discrete dynamics of a rimless wheel model shown in Fig. 2. This model consists of massless leg flames whose length is l [m] and the total mass, M [kg], is concentrated at the central position. The relative angle between the leg flames is \u03b1 [rad]. Given a suitable initial condition, the rimless wheel rolls down a slope, and the rolling pattern converges to 1-period 978-1-4244-3804-4/09/$25.00 \u00a92009 IEEE 3214 stable limit cycle if the next collision always occurs. The stability principle is explained as follows. Let i be the step number and K [J] be the kinetic energy" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002170_1.2787013-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002170_1.2787013-Figure2-1.png", "caption": "Fig. 2. Orientation of Wiffle ball for wind tunnel experiments.", "texts": [ " The forces normal and parallel to the direction of airflow were measured with a strain gaugebased force sting balance manufactured by Aerolab13 and with a student-designed instrument capable of measuring forces on a spinning ball.14,15 The ball was mounted in the test section of the wind tunnel, and its orientation is characterized by the angle between the oncoming airflow and the horizontal the conventional angle of attack , the angle between the horizontal and the axis of the perforation formation, and the angle between the oncoming airflow and the axis of the perforation formation, or + . This geometry is illustrated in Fig. 2. An angle of attack of =0\u00b0, and the holes positioned facing the oncoming airflow, corresponds to angle =0\u00b0; =90\u00b0 corresponds to a ball at zero angle of attack with the holes on top, as in the center image of Fig. 1. Dimensionless numbers are of much value in fluid dynamics. One of the most important dimensionless numbers is the Reynolds number, the ratio of the fluid inertia to its viscosity. The Reynolds number is defined as Re = Vd , 1 where is the fluid air density, V is the fluid free-stream velocity, d is the diameter of the Wiffle ball, and is the fluid air viscosity" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002229_s026357470700389x-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002229_s026357470700389x-Figure1-1.png", "caption": "Fig. 1. (a) MNM and (b) gradient method.", "texts": [ " Other parts of derivation are the same as in the previous section. In this case, the desired orientation is calculated based on modified Newton direction dn instead of negative gradient direction dng \u03b8d = arctan 2(sgn (x)dny, sgn (x)dnx) (9) Theorem 1: Consider a system of the form (2) and assume the existence of a dipolar inverse Lyapunov function (DILF)-generted potential field. Then the control law (Eqs. (3)\u2013(5),(9)) guarantees obstacle avoidance and global asymptotic convergence for the system. Proof: See the Appendix. In Fig. 1, we show the performance of gradient method and Newton\u2019s method in a typical obstacle avoidance task. To facilitate the discussion, here and elsewhere in this paper, we adopt the notion \u201cstepsize\u201d from optimization theory6 and define it as the product of the speed and the sampling interval. The solid dark line represents the trajectory of the robot, the target is represented by the symbol @ and the obstacle is represented by the symbol #. This figure illustrates that the MNM can achieve a much faster progress in presence of obstacles" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001559_095440505x32814-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001559_095440505x32814-Figure2-1.png", "caption": "Fig. 2 The length of the transducer embedded in the roll exceeds the contact length", "texts": [ " By analysing the conditions at the leading edge and rear edge of the insert, a downhill flow at all positions in the contact zone is identified as explained in the following section. The transducer for measuring the normal and frictional forces is mounted into the roll in a somewhat protruding manner. The length of the transducer is larger than the length of the contact zone. In this way, the leading edge will leave the deformation zone before the rear end enters. A gap between the transducer and the roll is needed to allow the transducer to move freely, as shown in Fig. 2. The design of the transducer consists of two pillars connected by stiff beams. The pillars are equipped with strain gauges, which enable individual measurements of the normal and friction forces. In Fig. 3, the deformation pattern for a tangential and a normal loading of the transducer are shown. The transducer measures the integrated normal force and frictional force at the contact area. A constant situation is expected when the transducer covers the full deformation zone. In Fig. 4, the four different phases of the measuring cycle is shown" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002885_wnwec.2009.5335810-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002885_wnwec.2009.5335810-Figure2-1.png", "caption": "Figure 2. 3D modeling", "texts": [ " It has an important significance for improving the performance of transmission system for wind turbine. II. GEOMETRIC PARAMETER Taking a high-speed gear pair in a MW-class wind turbine transmission system as an example, the contact and bending of the gear pair are analyzed by using FEM (Finite Element Method) considering bearing clearance and thermomechanical coupling in the following sections. The transmission schematic and 3D modeling of the gear pair including the shafts and bearings are shown in Fig. 1 and Fig.2, respectively. The gear contact ratio factor of this gear pair is 4.582; therefore, five gear teeth are modeled in this study. These gear teeth are numbered as from No.1 to No.5, respectively. The FEA modeling of gear pair is shown in Fig. 3. 978-1-4244-4702-2/09/$25.00 \u00a92009 IEEE III. ANALYSIS OF GEAR PAIR CONSIDERING BEARING CLEARANCE There are three type bearings used mounted at the shafts in the high-speed gear pair: NCF****, NJ****, 3****. The clearance of bearing NCF**** at the two sides of wheel in hollow shaft is 225 \u03bcm and the clearance of bearing NJ**** at the left side of pinion in output shaft is 70\u03bcm" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000557_b:joep.0000036510.38833.05-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000557_b:joep.0000036510.38833.05-Figure1-1.png", "caption": "Fig. 1. Scheme of heat exchange between the heated layers of a powder particle: 1) oxide film; 2) powder particle; 3) contact zone; 4) isotherms; 5) direction of propagation of heat.", "texts": [ " Let us introduce the following model assumptions: (a) heat in the contact zones of all powder particles is released uniformly; the processes of heat exchange in them are independent of each other; (b) in view of the rapidity of the process of EDS, there is virtually no heat exchange with the ambient medium; (c) heating of a powder particle is carried out owing to the heat sources that are the contact zones; (d) the zone of contact represents a hemisphere whose diameter is determined by Eq. (13); (e) the specific energy in the zone of contact expended in heating the particle is determined by the ratio q = CU0 2 4Nc.powd.f ; (31) (f) propagation of heat inside the particle itself is by the scheme of Fig. 1, according to which the first half of the heat released in the contact zone goes to heat one powder particle whereas the second half goes to heat the other particle; (g) heat exchange between the internal particle layers is determined by the Fourier law q = \u2212 \u03bb \u2206T \u2206x \u2206S\u2206t . (32) Assumptions (a)\u2013(g) given above, Eq. (31), and criterion (12) yield the following: (1) the value of the specific thermal power P0 of the contact zone of a powder particle over one period in EDS is P0 = 2\u03c0q \u03c9 , (33) where \u03c9 is computed from (4); (2) based on the area of a spherical segment through which the heat flux into the particle goes and which is at a certain distance x from the contact zone, the change in the temperature between the thickness element of the heated layer dx and the contact zone is determined with account for (31)\u2013(33) as dT = \u2212 P0 \u03bb\u03c0 dx x (x + \u221aD0 2 \u2212 x2 \u2212 D0) ; (34) (3) when the heat release in the contact zones of a powder particle in EDS is uniform, a temperature field is formed in it in such a manner that a point with a minimum value of the temperature is at the geometric center of the particle" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002564_s10015-008-0560-2-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002564_s10015-008-0560-2-Figure1-1.png", "caption": "Fig. 1. Mechanism of manipulator", "texts": [ " Considering the fi nal condition about displacement and velocity of the passive joint, the trajectories of velocity for energy saving are calculated by iterative dynamic programming. Initial searching region, which is parallel, is shifted to minimize the energy consumption of the motor. The dynamic characteristics of manipulator control based on the above mentioned trajectory are analyzed theoretically and investigated experimentally. 2 Modeling of the manipulator with passive joint The dynamic equations of the manipulator with three degrees of freedom as shown in Figure 1, which is able to move in a vertical plane, are as follows: A A A A A A A A A 11 12 13 21 22 23 31 32 33 1 2 3 \u23a1 \u23a3 \u23a2 \u23a2 \u23a4 \u23a6 \u23a5 \u23a5 \u23a1 \u23a3 \u23a2 \u23a2 \u23a2 \u23a4 \u23a6 \u23a5 \u23a5 \u23a5 = \u03b8 \u03b8 \u03b8 \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 1 2 14 2 3 24 3 34 \u2212 + \u2212 + + \u23a1 \u23a3 \u23a2 \u23a2 \u23a4 \u23a6 \u23a5 \u23a5 A A A (1) where A a a C a C A a C a C A a C11 1 4 2 6 23 12 4 2 6 23 13 6 23= + + = + =, , A a S a S a C14 4 1 2 2 2 6 1 2 3 2 23 7 1= +( ) + + +( ) \u2212 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 A a a C a C A a a C A a C21 2 4 2 5 3 22 2 5 3 23 5 3= + + = + =, , A a S a S a C24 4 1 2 2 5 1 2 3 2 3 8 12= \u2212 + + +( ) \u2212 \u03b8 \u03b8 \u03b8 \u03b8 A a a C a C A a a C A a31 3 5 3 6 23 32 3 5 3 33 3= + + = + =, , Key words Manipulator \u00b7 Trajectory \u00b7 Dynamic programming \u00b7 DC motor \u00b7 Minimum energy \u00b7 Passive joint 1 Introduction For the purpose of enlarging the work space, it is necessary to study the optimal control of the manipulator whose mechanism has a passive joint" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003046_1.3650515-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003046_1.3650515-Figure2-1.png", "caption": "Fig. 2 Infinitely wide stepped bearing with incompressible lubricant (a) Velocity-induced flow", "texts": [ " In a stepped pad in which the gap does not change with distance, dh/dx is zero everywhere except at the internal boundary. If, in addition, conditions are steady, dh/dl is also zero. Therefore, the right side of the Reynolds equation vanishes. The pressure rise must thus be caused by another effect, namely, the combination of the internal boundary and the relative tangential motion which acts like a flow source. If we consider the stepped pad to be composed of two regions separated by a reservoir, as shown in Fig. 2, the flow into this reservoir from the leading edge region will be Uhi/2, and the flow out of the reservoir through the trailing edge region will be Uhi/2, giving a net flow into the reservoir of U(hi \u2014 hi)/2 \u2014 UA/2. The reservoir is of finite size, and the flow out of the reservoir has to equal the flow into it. The flow out of the reservoir can only be through the bearing, which acts like a flow restrictor, the impedance of which is hi'/12fih + h1i/i2jj.li. The pressure necessary to force the flow out of the reservoir is thus V", " But an increase in load capacity can be obtained by use of a pocket-type bearing 2 2 2 / M A R C H 1 9 6 5 Transactions of the A S M E Downloaded From: http://fluidsengineering.asmedigitalcollection.asme.org/ on 01/19/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use Film Thickoeu in Trolling Edge Region (KJ V - inch Fig. Film Thicknau In Troiling Edge Region (hj) V \" Inch 19 (Continued) UMHJ e 6000 RPM - A 10500 RPM \u2022 24000 RPM Ugond O \u00a3000 RPM A 12000 RPM - \u2022 24000 RPM Scoring Choroclctillicl R = .5 inch \u2022 b - .5 Fig. 2 0 Numerical solut ions\u2014Rolls-Royce bearings Journal of Basic Engineering M A R C H 1 9 65 / 223 Downloaded From: http://fluidsengineering.asmedigitalcollection.asme.org/ on 01/19/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use with only a very small increase in bearing power and, depending on the method of manufacture, at no increase in bearing cost. Correlation Between Theory and Experiments The experimental data obtained are summarized in Tables 1 and 2. Table 1 lists all the important bearing parameters and gives test results obtained for q = 2", " \\ ; f - H \u2014 H - H \u2014 r - | - f - - H - 1 ^ \" \" H ~ H \" \" : : : ;;;; ; ; I i . i j i - N 1 ' \" T 1 - 0 . ' \u2022 M ! 1 i ! 1 : ! , : : : ! : i \u2022IISSi V V 1 1 U t . I l ! | M M ! \u2022 \\ \u2022 \u00ab . a : ! : ' 7- J J . - . . \u2022 \u2014 . 6 - - - \u2022 , | : : - \u2022 \u2014 5 - -:\u2022 '\u2022:\u2022 T l \u2022 : ' : . . ' . 4 ; \u2014 \\ * - - i \u2014 v \\ \\ r : r : : : : . | : .\u2022; i t l | M : . ; : ( i * m m . . : i : : : ! . . : : : : : ; \u2022 : - 0 T T \u2014 . 2 . 4 . 6 . 8 \u2022Si1 ... \\ \\ 1 1 1 f c f f l f c i i r T t - g + S i ^ - f f f l f f i M 1 .0 1 .2 1 .4 1 .6 \u2022 1 .8 2 . 0 : : C l e o r o n c e Ro t io h 2 . 1 Fig. 2 4 Design chart B\u2014Load ratio versus clearance ratio If this calculated power is less than the available power, proceed with the design. In many applications it is the available power that will determine the maximum bearing speed, size, and minimum gap. Journal of Basic Engineering \u2014 .7 I f e i l ' . 5 \u2022 - 1 - _ i \u2014 \"1- ' !!!!!!!! 1 1 1 : j\" * - 1 i i l l P I - . ! - 0 \u2014 20 | 1 = = s c a iij 4 g g i s n 8 \u2014 - 40 l 5 4 \" ^ = 60 80 led 6c . .... i \u00ab or ing N u i \u2022nber \u2022 - ~ ~ ~ r p , i , ,20 - P=h2 J I _ 4 0 \u2014 j - - _ ri-b2] Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003115_secon.2008.4494342-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003115_secon.2008.4494342-Figure2-1.png", "caption": "Fig 2. Flux distribution at different positions (i=16A, from unaligned position to aligned position).", "texts": [ " shows the 978-1-4244-1884-8/08/$25.00 \u00a92008 IEEE 480 meshed FE model and different material of SRM in ANSYS. The flux linkage ( ,i) can be determined by computing the magnetic vector potential A over the machine cross section. The 2-D magneto static problems with computing the vector potential A in Cartesian coordinates (x, y) are described by nonlinear Poisson\u2019s equation ( ) ( ) z zz J y AB yx AB x \u2212= \u2202 \u2202 \u2202 \u2202+ \u2202 \u2202 \u2202 \u2202 \u03bd\u03bd (1) Where (B) is the magnetic reluctivity and Jz is the source current density. As Fig 2. shows, the magnetic vector potential and flux distribution can be obtained by taking into consideration the saturation effects of the isotropic ferromagnetic stator and rotor material. the flux linkage in each phase is dVAJ i V = 1\u03c8 (2) this, after finite element discretization or meshing, becomes k n k k SA S Nl = = 1 \u03c8 (3) Where l is the axial length of machine, N is the number of turns per phase, and S is the area of phase winding. Magnetic co-energy Wco( , i) can be calculated on the basis of flux linkage ( ,i) as ( ) ( ) const i diiiW =\u03a8=\u2032 \u03b8\u03b8\u03b8 0 (4) Torque is calculated from the co-energy derivative with respect to angular position as ( ) ( ) consti iWiT =\u2202 \u2032\u2202= \u03b8 \u03b8\u03b8 (5) the static torque T(i,\u03b8)characteristics can be obtained by equation (4) and (5) from \u03c8(i,\u03b8) characteristics" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003386_978-3-642-39348-8_44-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003386_978-3-642-39348-8_44-Figure5-1.png", "caption": "Fig. 5 Strip gauge positioning at the root of a test gear tooth", "texts": [ " Since the total contact ratio of the test gear pairs is between 1 and 2, two teeth of one test gear will be instrumented to measure strain (access and recess phases) over simultaneously meshing tooth pairs. Moreover, to assess the effects of lead microgeometry modifications and imposed misalignment, multiple strain gauges will be placed along the face width. Strip gauges of type Vishay Micro-Measurement EA-06-031PJ-120 will be used. Each strip gauge has 22.7 9 4.8 mm dimensions and consists of a linear pattern of 10 strain gauges (Fig. 5) each having base length of 0.79 mm and width of 1.78 mm. A short base length is chosen to have a local measurement of deformation. Accurate local measurement however requires careful gauge positioning and alignment [16] because of the steep strain gradients at the tooth root. Two measurement methods will be used to obtain dynamic strain: a conventional quarterbridge configuration and a direct measurement of resistance (described in the next section). In both cases limitations may be caused by the data acquisition system and by the slip rings used to transfer the signals from the rotating components" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000928_gt2004-53708-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000928_gt2004-53708-Figure1-1.png", "caption": "Fig. 1. High speed bearing chamber test rig", "texts": [ " Symbol Unit Quantity A [m2] Surface b [m] Width of bearing chamber BC [-] Bearing chamber d [m] Diameter Dh [m] Hydraulic diameter g [m/s2] Acceleration of gravity h [m] Height k [-] Velocity ratio m [-] Parameter m\u0307 [kg/s] Mass flow M [Nm] Moment n [rpm] Rotational speed R [-] Radial position r [m] Radius Re [-] Reynolds-number T [K] Temperature u [m/s] Velocity U [m] Circumference V\u0307 [l/h] Volume flow rate y [m] Wall distance \u03b5r [-] Specific dielectric coefficient \u03d5 [\u25e6] Angle \u03bb [-] Friction factor \u03bd [m2/s] Kinematic viscosity \u2126 [s\u22121] Angular velocity of rotor \u03c1 [kg/m3] Density \u03c3 [-] Parameter \u03c4 [N/m2] Shear Subscripts f Film g Gas in Inlet I, II, III Number of configuration l Liquid r Rotor s Stator sh Shaft t Turbulent w Wall \u03c4 Shear 0 Surface + Dimensionless Copyright \u00a9 2004 by ASME ms of Use: http://www.asme.org/about-asme/terms-of-use Do A co-axial sectional view of the test rig is shown in Fig. 1. The rig was used first by Wittig et al. [2], who provided a detailed description of the test facility. Due to the compact and modular design, variations of bearing chamber geometries and roller bearings can be realized with relatively low financial and manufacturing effort. The rotor (nmax = 20000rpm) is supported by two bearings. A ball bearing is used to prevent axial and radial displacements of the rotor. Due to the small size of the ball bearing only little amounts of oil were supplied for lubrication" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000788_cira.2003.1222263-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000788_cira.2003.1222263-Figure2-1.png", "caption": "Figure 2: The model of human parking support", "texts": [ " - d= - d t - ucos~cns0 2 = vcosdsine (1) - d e - dt - -sm+ L Here, x and y are the positions of the four-wheeled vehicle, 0 is an angle that the progress direction of the z axis and the four-wheeled vehicle does, II is a speed of the car, 4 is an angle of the steering wheel that instructs in movement by turning radius R, L is a wheelbase of the vehicle. 0 is center of the turn circular arc. The precondition for the parking control in this paper is shown below. Preconditiuns 0 The shape of the parking lot is already-known. The vehicle runs in low speed A dynamic obstacle doesn't exist e Vehicle enter the parking lot from the rear side. 0-7803-7866-0/03/$17.00 2003 IEEE 682 3 Intelligent Parking Support System 3.1 Human's parking support is modeled as follows. The support knowledge of the guidance member in the parking lot like Fig. 2 is targeted. The guidance member understands surrounding circumstances, the situation of the vehicle and the situation of the operation (Detect). Next, the strategy is mapped out based on empirical knowledge they have, and the target is set based on it (Target setting). Next, it thinks about the best operation to reach the target by predicting the state after a few seconds (Decison of operation candidate). Once operation candidate is decided, the guidance member think about offer information according to the situation of the vehicle and the operation at that time (Decision of support language information)", " Support information is \u201dStop\u201d. Pmgress of dimtion Strategy Tactics Steering wheel straight Attention Stop Processing and supplying the visual information In \u201cProcessing of the support image\u201d, based on the state of vehicle and maintained map information, three dimension image is generated. That draws the position of the vehicle and the parking lot that is seen from guidance member\u2019s aspect. When parking, the guidance member outside the vehicle can put out an appropriate instruction as shown in Fig. 2. This is why he understands the situation such as the position of the obstacle and the vehicle, the speeds, and the steer corners easily. Then, a floor was made on a virtual space, and the parking lot and the vehicle of the same scale as the experimental environment were prepared. The garage and the state of the vehicle are displayed on the monitor by using 3DCG as visual support information(Fig. 6). 3.6 Hardware Configration To decide the position and the direction of the vehicle, the developed system detect the speed and steer angle using rotary encoder and potentiometer" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003722_atee.2013.6563405-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003722_atee.2013.6563405-Figure1-1.png", "caption": "Fig. 1. Multi-slice model of the ma", "texts": [ " inding distributed in 36 slots, n asynchronous motor. There h phase. For each stator slot th the cross section of 0.246 e stator is 134 mm, the inner the airgap is 0.35 mm and the es. Each pole is provided with inum that are short circuited rpolar space is also filled with ited to the same end rings. In hat the starting/damping cage bars (with different cross- degrees. The skewing of the r this category of motors in pples caused mainly by the d by the magnetically rotor [7], [8]. illustrated in Fig. 1. lt in the simulation program. axial slices properly shifted in chine geometry. 978-1-4673-5980-1/13/$31.00 \u00a92013 IEEE order to account for the skewing angle. T computations are performed separately on e effect is mutually considered [7], [10]. III. DETERMINATION OF THE DIRECT- AN AXIS SYNCHRONOUS INDU The problem of estimating the direct- an inductances for the synchronous reluctan common one [11], [12], [13]. In the fo method of approximating the two extreme FEM is introduced. The method consists in slowly varying th the resultant magnetic field and the relati rotor at a small slip (s \u2248 0) [13]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003612_s11044-010-9190-2-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003612_s11044-010-9190-2-Figure7-1.png", "caption": "Fig. 7 Vibration model of a soft mounted electrical machine", "texts": [ "2, the shaft centre point W is now forced to move on an orbit, already without considering additional elastic bending of the shaft and elasticity of the bearings and their substructures (Fig. 6). Due to the displacement of the magnetic centre M and the centre of rotor mass S from the axis of rotation, additional forces\u2014unbalance forces and magnetic forces\u2014occur. Considering the elasticity of the shaft, the bearings and their substructures, these forces influence the orbit of the shaft centre point W [7]. After the different kinds of eccentricity are described, the vibration model of a soft mounted electrical machine is derived (Fig. 7). The model is a simplified model, describing the vibration in a plane (plane y, z). It is based on the general model in [9], which has been adapted in this paper for especially focusing on the excitation by rotor eccentricities and to derive therefore the mathematical description of orbital movements. The model consists of two masses, rotor mass mw, concentrated at the shaft\u2014rotating with angular frequency \u03a9\u2014and stator mass ms, which has the inertia \u03b8sx and is concentrated at the centre of gravity S" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001875_6.2007-6195-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001875_6.2007-6195-Figure4-1.png", "caption": "Figure 4. Single Sided Qualification Parts", "texts": [ " These requirements are a set of minimum tensile strength values, a set of mean tensile strength values, a maximum coefficient of variation of the entire data set of tensile data, a minimum fatigue life for a given load level, a minimum fracture toughness value and chemistry tests to verify the composition of the final article. These requirements will be shown in the following section. To assist in this qualification process, Boeing has developed 3 demonstration part geometries. These parts contain geometric features typical of aerospace parts, such as curved flanges, multiple intersecting stiffeners and varying heights. Figures 3 and 4 show these three part geometries. The part shown in Figure 4 is a two sided part, with depositions on both sides of the substrate plate. American Institute of Aeronautics and Astronautics These part geometries also contain sufficient volume to extract numerous coupons for mechanical testing. CAD models of these parts are supplied to the companies, which also allows them to compare the geometric accuracy of their deposited surfaces to the engineering model. IV. Process Overview At the onset of this qualification effort, Sciaky undertook the Stage 1 testing as a means of prequalifying their system" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000342_2004-01-1448-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000342_2004-01-1448-Figure2-1.png", "caption": "Figure 2. DDC Series 60 rocker arm assembly with a Si3N4 cam roller follower (dark gray) in the fuel injector position, allowing fuel injection pressures of over 210 MPa (30 ksi)10.", "texts": [ "9 (off scale) >75 Electrical Volume Resistivity (\u2126\u00b7cm) 60\u00b710-6 10-20\u00b710-6 >1\u00b71013 Thermal Expansion Coefficient (10-6 \u00b0C \u20131) 10.9 6\u00b710-6 3.2 As the contact stresses on the metal rollers increased in the early 1990\u2019s (Figure 1), field failure of metal rollers increased, increasing the system cost, which eventually made certain grades of Si3N4 rollers affordable. In 1993, Detroit Diesel Corporation was the first US diesel engine manufacturer to introduce Si3N4 cam roller followers into their Series 60 engines (Figure 2). Since then, additional successful applications for Si3N4 rollers have developed (Table 2). In some cases (Detroit Diesel Corporation and Stanadyne), low levels of ceramic spalling were found with the grades of Si3N4 initially used in the applications. Both companies, after evaluation of other material grades, chose a gaspressure sintered reaction bonded silicon nitride (SRBSN) material, Ceralloy 147-31N produced by Ceradyne to eliminate the spalling problem. For fuel pump applications (Figure 3), the roller sees both rolling contact against the internal cam ring and sliding contact against the roller shoe under boundary lubrication conditions in a hot diesel fuel environment" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000732_0094-114x(78)90039-3-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000732_0094-114x(78)90039-3-Figure1-1.png", "caption": "Figure 1. WaR's straight-line linkage.", "texts": [ " The center C of the coupler A B = 2b describes a sextic c which has the shape of the digit 8, if 2b < LM and 2[a - b] < LM < 2(a + b). Each one of the two branches passing through the inflection node O in the middle of the fixed base L M can serve, in a certain neighbourhood of O, to approximate a segment of a straight line. Watt is said to have been prouder of this detail of his steam-engine than of the machine as a whole. \u2022 The simplest and most natural proposal for a good approximation consists in using a linkage which in its initial posi t ion-LAoBoM forms right angles at Ao and B0 (Fig. 1), as then the coupler curve c has a 5-point contact at Co = O with the ideal straight line AoBo (viz. Section 4). In principle this idea belongs to Burmester theory which asks for the closest possible contact between the ideal line and an approximating curve in a certain point. Now just in kinematics it is well-known that better results can be obtained by Chebyshev theory which aims to minimize the maximal deviation from the ideal line, by allowing the approximating curve to reach the maximal deviation as often as possible in a certain interval" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002236_2007-01-3740-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002236_2007-01-3740-Figure3-1.png", "caption": "Figure 3 explains this approach using a simplified single roller model in high regime. Applying a reaction force F to the roller causes a reaction torque (TA and TB) at the Variator discs and consequently an acceleration of the two inertias (engine side inertia A and vehicle side inertia B). This may change the speed of the engine and / or vehicle inertia resulting in a change of Variator ratio. The application of a castor angle to the roller carriages (Figure 4) enables the rollers to \u2018steer\u2019 to a new angle of inclination (ratio). This happens automatically \u2013 only the Variator disc speeds and reaction force are defined externally. In the Torotrak design, reaction force is applied hydraulically to individual roller carriage pistons.", "texts": [ " Typical premium automotive IVTs therefore have two operating regimes; low regime is a split-power system providing Geared Neutral, forward and reverse drive; high regime bypasses the epicyclic gear set and the power flow is directly from the engine via the Variator to the road wheels so extending the ratio of the transmission in forward drive to high overdrives (typically 60mph / 1000rpm). The IVT is torque controlled, which essentially means that the required system torque is set by hydraulic pressure and the Variator follows the ratio automatically [3]. It is mainly this control approach enabled by the full toroidal Variator that renders geared neutral a safe controllable system and eliminates the need for a starting device such as a torque converter or friction clutch [4]. Figure 3: Principle of Torque Control Castor Angle Tangential Force EndloadEndload Reaction Force Figure 4: Variator Force Balance To compete in the cost competitive OPE market, the full toroidal traction drive Variator required significant simplification in both design and operation. Recognising the low power and torque requirements of this market, the simplification was primarily achieved by utilizing a single cavity design, reducing the number of rollers in the cavity from three to two and introducing a simple \u2018yoke\u2019 style roller control mechanism (figure 5)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002175_robot.2007.363146-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002175_robot.2007.363146-Figure5-1.png", "caption": "Fig. 5. The first case of actuation singularity", "texts": [ " In another way, according to screw theory[15;16], $r2=[lr2, mr2, nr2, pr2, qr2, rr2]T (20) where lr2= (1/\u03bb) (x3-x2)/(x2y3- x3y2); mr2= (1/\u03bb) (y3-y2)/(x2y3- x3y2); nr2= 0; pr2= (1/\u03bb) (m4n5-m5n4)/(l4m5-l5m4); qr2= (1/\u03bb) (l5n4-l4n5) /(l4m5-l5m4); rr2= (1/\u03bb); 2 2 2 2 rr ml +=\u03bb ; Assume the unit direction of normal vector for plane P45 is n45=[n45x,n45y,n45z] = [m4n5-m5n4, l5n4-l4n5, l4m5-l5m4], then pr2= (1/\u03bb) n45x / n45z (21) qr2= (1/\u03bb) n45y / n45z (22) $ r2=(1/(\u03bbn45z))[n45zlr2,n45zmr2,n45znr2, n45x,n45y,n45z]T (23) Thus, wrench system $r of the movable platform consists of six wrenches, $ r2 of five limbs and $r1, T)5( 2 )4( 2 )3( 2 )2( 2 )1( 21 ],,,,,[ rrrrrrr $$$$$$$ = (24) where )( 2 i r$ denotes the $r2 of the ith limb, i=1,2,3,4,5. At general configuration, the rank of the wrench system in Eq.(18) is usually six and hence the movable platform can be controlled by five actuators. However, if wrench system degenerates at some special configuration, the actuation singularity occurs. There are three cases of actuation singularities for the manipulator 5-RRR(RR). 1) First case: As shown in Fig.5, rotation center locates in P23 of five limbs. Let origin O of reference frame locate at the rotation center, and Z-axis be perpendicular to the base plane. In this case, rotation center is on the intersection line of P23 and P45 since it is always in P45, namely the axis of $r2. Moreover, $r1 also passes through the rotation center. Then six constraint wrenches shown in Eq. (24) intersect at the rotation center. According to the Grassmann geometry[5], there are only three linear independence vectors in the spatial intersection case", " Obviously, rank of wrench system, $r is three. In this case, the movable platform can still rotate about the rotation center even after locking five actuators. In other words, there are three uncontrollable rotation DoF. This actuation singularity can be passed by choosing different joints as actuators. For example, six wrenches in Eq.(24) are not dependent if choosing three R1 and two R2 as actuators. 2) Second case: As shown in Fig.6, the plane P45 of five limbs parallel to the base plane. Let the reference frame be the same with Fig. 5. In this case, P45 of five limbs will be the same plane. And $r2 of five limbs will be also in the plane. According to Grassmann geometry[5], the rank of coplanar linear vectors (five $r2) is three. Thus, the number of linear independent constraint wrenches (five $r2 and $r1) is four. According to screw theory, from Eq.(23), the last three entries of the $ r2 are the direction cosine of the normal vector for P45. When the movable platform parallel to the base plane, the five P45 planes are the same", " Then, their normal vectors parallel to each other, namely T 222 )( 2 ]1,0,0,,,)[\u03bb/1( i r i r i r i r nml=$ (27) Then \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 = 111110 000000 000000 \u03bb 0 0 )\u03bb/1( 5 2 4 2 3 2 2 2 1 2 5 2 4 2 3 2 2 2 1 2 5 2 4 2 3 2 2 2 1 2 rrrrr rrrrr rrrrr r nnnnn mmmmm lllll $ (28) Obviously, rank of the wrench system, $r is four. In this case, there are two uncontrollable freedoms for the movable platform. The movable platform can instantaneously rotate around any axis in plane O-XY even after locking five actuators. One of input selections to avoid this singular configuration is choosing three R1 and two R4 as actuators. 3) Third case: Configurations of five limbs are symmetrical about Z-axis. Let the reference frame be the same with Fig. 5. In this case, five $r2 will also be symmetrical about Z-axis. One $r2 can be achieved by transforming another by a rotation around Z-axis. Based on the screw theory, assume O-YZ plane of reference frame parallel )1( 2r$ , and )1( 2r$ intersect with X-axis at [x, 0, 0], namely T)1( 2 ]00[ xmxnnmr \u2212=$ (29) Since any )( 2 i r$ can be achieved by transforming )1( 2r$ after a rotation around Z-axis, then )( 2 i r$ can be expressed as T)( 2 ][ xmxncxnsnmcms iiii i r \u03b1\u03b1\u03b1\u03b1 \u2212\u2212=$ (30) where \u03b1i is the angle between the first limb and ith limb; s\u03b1i, c\u03b1i denote sin(\u03b1i) and cos(\u03b1i), respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001555_iecon.2005.1569184-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001555_iecon.2005.1569184-Figure3-1.png", "caption": "Fig. 3. Wheel Model.", "texts": [ " Tassist = \u03b1 1 1 + \u03c4s Thuman (1) Here \u03b1 is power-assistance-ratio, Tassist is the assist force, Thuman is the input torque from the pushrim, and \u03c4 is the time constant. For better ride quality, the ascent should be fast and the descent slow. So in this work time constant \u03c4 is decided as the following (2), \u03c4 = { \u03c4fast = 0.08[s], d dtThuman > 0; \u03c4slow = 1.0[s], d dtThuman < 0. (2) 18330-7803-9252-3/05/$20.00 \u00a92005 IEEE B. System Model When a person is sitting, COM of the him and the wheelchair frame is supposed to be at the surface of his abdomen. And by neglecting two front wheels, the model of wheelchair can be represented as Fig. 3. Please note that although two front wheels are not shown in figure, operator cannot fall \u201dforward\u201d, and on horizontal plane there is \u03d50 \u2265 \u03d5. M : Mass of the driving wheel m : Mass of operator including wheelchair frame r : Radius of the driving wheel JM : Moment of Inertia of the driving wheel Jm : Moment of Inertia of operator and frame \u03b8 : Rotational angle of the driving wheels \u03d5 : Angle of the vertical to COM \u03d50 : Initial angle of the vertical to COM ax, ay : Accelerations of COM Ax, Ay : Accelerations of the wheel III" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001489_s00170-006-0516-4-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001489_s00170-006-0516-4-Figure5-1.png", "caption": "Fig. 5 The 3-RPRU simulation machine tool with a spiral curve tool path. a The pose as ey=60 cm, ex=75 cm. b The pose as ey=80 cm, ex=80 cm", "texts": [ "3 The 3-RPRU parallel simulation machine tool In light of the 3-RPRU parallel simulation mechanism, when z of m of the 3-RPRU simulation mechanism is replaced by a tool (such as a mill cutter or a drilling cutter) and its driving motor, a novel 3-RPRU parallel simulation machine tool is created, as shown in Figs. 5 and 6. How to use this 3-RPRU parallel simulation machine tool to machine S is a key problem to solve. When the tool path of a reciprocating straight-line is used to machine S, a simulation parallel machine tool is created, see Fig. 5, its creation processes are explained below. Step 1 Delete the driving dimension of ri, thus ri (i=1, 2, 3) of the 3-RPRU simulation mechanism are transformed into the three driven rods. Step 2 Connect the free end of z to S at point p1 in the 3D sketching environment, and give z a fixed dimension z=30 cm in length. Step 3 Constitute a tool driving line e, connect its two ends to S at point p1 and P0 at point p2 by adopting the linesurface coincident command. After that, set e coincident with z by using the coincident command" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002168_isie.2007.4374753-Figure12-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002168_isie.2007.4374753-Figure12-1.png", "caption": "Fig. 12. Current density distribution", "texts": [], "surrounding_texts": [ "Thermal experiments demonstrate that the motors present different characteristics with respect to conventional squirrel cage motor, since the current, that is the main cause of the increase of the temperature in the stator winds, practically does not change when changing the load. This makes these motors able to work in regimes with high slip without hardly changing their thermal conditions. This fact is almost impossible in squirrel cage motors. On the other hand the spiral sheet motors are more resistive than cage of squirrel motors, we can appreciate an increase of temperature with respect to motor of cage of squirrel when they are working in the same conditions, but it is in any case much smaller than the increase than undergoes other rotors like the solid rotor." ] }, { "image_filename": "designv11_61_0002821_j.ijsolstr.2008.05.021-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002821_j.ijsolstr.2008.05.021-Figure2-1.png", "caption": "Fig. 2. Circular beam contact model (a) and real conrod big end system (b).", "texts": [ " Such formula may be helpful in the early design stage by allowing fast parametric studies. To that end, a simple model of conrod is considered here. It consists of two elastic circular beams maintained by contact with friction and pre-stresses. The relative displacement of the beams on the contact interface is analysed when the internal beam, representing the bearing shell, is submitted to a cyclic rotating concentrated radial load and the external one, representing the conrod big end, is punctually clamped (Fig. 2a). As a first approach, this system represents a conrod big end assembly, loaded during an engine cycle (Fig. 2b). Recent works on the analysis of the relative slip of elastic solids maintained in contact with friction under cyclic loadings can be found in the literature. The analogy between relative slip mechanisms and classical long-term behaviours observed in . All rights reserved. toni). plasticity has been pointed out, plastic strains playing the role of slips, cf. (Koiter, 1960; Debordes, 1976; Maier, 2001; Nguyen, 2006; Bouby et al., 2006; Antoni, 2005; Antoni et al., 2007; Klarbring et al., 2007). The first part of the paper is devoted to a short presentation of the existing results on slip-shakedown theory", " In particular, the upper bounds derived from the slip-shakedown and limit analyses can be useful for a quick estimate of the critic al rotation load. In the last part of the paper, a step-by-step numerical calculation of an equivalent 2D-conrod model under the same cyclic load and Coulomb friction is performed by the finite element method in order to explore the validity of the assumption of small coupling and of the consistency of the beam modelling. Let us consider the modelling of the conrod system presented on Fig. 2a, within the framework of elastic curved beams under the assumption of small perturbations (Garrigues, 1999; Salen\u00e7on, 2002). The system is in quasi-static transformation under a cyclic radial force of amplitude F(b) located at angle b(s) at time s. The reference state is a self-equilibrated pre-stress state ensuring in particular a uniform normal force p0 > 0 on the contact interface Cc. The contact interface is assumed to be the whole interface when the pre-stress is strong enough. If the two different beam-centrelines with their respective radius are considered, the model will require an appropriated write-up of contact conditions. As a first approximation, a simpler modelling is adopted here. It is assumed that the mean lines of both beams are coincident with the contact interface. Let R be the corresponding radius (see Fig. 2a). Let h be the circumferential position and ( p(h,s),q(h,s)) be respectively the normal and tangential forces (per unit length) applied by internal beam (1) on external beam (2). The contact force~f 2!1 applied by solid (2) on solid (1) in the loaded state is ~f 2!1\u00f0h; s\u00de \u00bc p\u00f0h; s\u00de~e2 q\u00f0h; s\u00de~e1; 8h 8s \u00f01\u00de On the contact interface, a bilateral contact is considered. Additionally, a standard friction law is assumed: jqj < k) stick; jqj \u00bc k) possible slip \u00f02\u00de The tangential sliding force k is a constant or a given function for a Tresca friction and k(h,s) = lp(h,s) for a Coulomb friction with a friction coefficient l. Let (ui(h,s),vi(h,s)) be respectively the tangential and normal displacements of solid (i) (i = 1,2) at position h and time s with respect to the reference state, see Fig. 2a. If (Ni0,Mi0,Ti0) denote, respectively the normal load, bending moment and shear force of beam (i) (i = 1, 2) at the pre-stress state, the normal and generalized forces in the current state can be conveniently decomposed as p\u00f0h; s\u00de \u00bc p0 \u00fe Dp\u00f0h; s\u00de Ni\u00f0h; s\u00de \u00bc Ni0\u00f0h\u00de \u00fe DNi\u00f0h; s\u00de Mi\u00f0h; s\u00de \u00bc Mi0\u00f0h\u00de \u00fe DMi\u00f0h; s\u00de Ti\u00f0h; s\u00de \u00bc Ti0\u00f0h\u00de \u00fe DTi\u00f0h; s\u00de 8>><>>: \u00f03\u00de to obtain the governing equations of the quasi-static transformation of the system under a given cyclic history b(s) as \u00f0a\u00de Bilateral contact v1\u00f0h; s\u00de v2\u00f0h; s\u00de \u00bc 0;p0 \u00fe Dp\u00f0h; s\u00de > 0 \u00f0b\u00de Friction jq\u00f0h; s\u00dej < k\u00f0h; s\u00de ) _u1\u00f0h; s\u00de _u2\u00f0h; s\u00de \u00bc 0 jq\u00f0h; s\u00dej \u00bc k\u00f0h; s\u00de ) _u1\u00f0h; s\u00de _u2\u00f0h; s\u00de \u00bc kq\u00f0h;b\u00de; k P 0 k\u00f0h; s\u00de is a constant or a given function for a Tresca friction k\u00f0h; s\u00de \u00bc lp\u00f0h; s\u00de for a Coulomb friction: \u00f0c\u00de Constitutive equations \u00f0elastic beams\u00de o2vi oh2 \u00fe vi\u00f0h; s\u00de \u00bc Sf iDMi\u00f0h; s\u00de RStiDNi\u00f0h; s\u00de oui oh \u00bc vi\u00f0h; s\u00de \u00fe RStiDNi\u00f0h; s\u00de ( ; 8i \u00bc 1;2 \u00f0d\u00de Equilibrium equations Beam \u00f01\u00de 1 R oDN1 oh DT1\u00f0h;s\u00de R q\u00f0h; s\u00de \u00bc 0 1 R oDT1 oh \u00fe DN1\u00f0h;s\u00de R \u00fe Dp\u00f0h; s\u00de \u00bc 0 1 R oDM1 oh \u00fe DT1\u00f0h; s\u00de \u00bc 0 8><>: \u00f04\u00de Beam \u00f02\u00de 1 R oDN2 oh DT2\u00f0h;s\u00de R \u00fe q\u00f0h; s\u00de \u00bc 0 1 R oDT2 oh \u00fe DN2\u00f0h;s\u00de R Dp\u00f0h; s\u00de \u00bc 0 1 R oDM2 oh \u00fe DT2\u00f0h; s\u00de \u00bc 0 8><>: \u00f0e\u00de Continuity and boundary conditions u2\u00f0 p=2; s\u00de \u00bc 0 v2\u00f0 p=2; s\u00de \u00bc 0 dv2 dh \u00f0 p=2; s\u00de \u00bc 0 8><>: u2\u00f03p=2; s\u00de \u00bc 0 v2\u00f03p=2; s\u00de \u00bc 0 dv2 dh \u00f03p=2; s\u00de \u00bc 0 8><>: \u00bdDNi \u00f0h \u00bc b; s\u00de \u00bc 0 \u00bdDMi \u00f0h \u00bc b; s\u00de \u00bc 0 \u00bdDT1 \u00f0h \u00bc b; s\u00de \u00bc F\u00f0b\u00de 8><>: \u00bdui \u00f0h \u00bc b; s\u00de \u00bc 0 \u00bdvi \u00f0h \u00bc b; s\u00de \u00bc 0 dvi dh h i \u00f0h \u00bc b; s\u00de \u00bc 0 ; 8i \u00bc 1;2 8><>: where [f](x) = f(x+) f(x ) represents the jump in function f at point x, (Ei,Si,Izi) denote respectively the Young modulus, the cross-section area and the moment of inertia of beam (i) (i = 1, 2) and Sti \u00bc 1 EiSi ; Sf i \u00bc R2 EiIzi ; 8i \u00bc 1;2 \u00f05\u00de The determination of the solution of this evolving problem is complex since it requires a complete step-by-step calculation starting from any given initial state ut(h, 0) = u1(h, 0) u2 (h, 0) = ut0(h)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002697_s11249-009-9510-y-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002697_s11249-009-9510-y-Figure1-1.png", "caption": "Fig. 1 a The contact between the equivalent sphere with surface roughness and the semi-infinite half space. The equivalent sphere with two types of surface roughness is considered; b the sinusoidal surface roughness, and c Gaussian height distributed surface roughness", "texts": [ " For the analysis of surface with sinusoidal roughness, the effects of asperity amplitude and wavelength on the fatigue life are investigated. Through the analysis for spherical surface with a Gaussian height distributed roughness, the variation of the fatigue life and the crack initiation depth according to the center line average roughness Ra are revealed. 2 Analysis 2.1 Contact Model Physical contact model used in the present study is the contact between two spherical bodies. Generally, rolling contact of two spheres is considered as the contact between the equivalent sphere and the semi-infinite half space (Fig. 1a). In the present study, the equivalent sphere with two types of surface roughness is considered. One is sinusoidal surfaces roughness as shown in Fig. 1b. For the parametric study, various sinusoidal surfaces are generated with different asperity amplitudes and wavelengths. The other is a Gaussian height distributed surface roughness as shown in Fig. 1c. Gaussian surface with the expected standard deviation of surface heights and autocorrelation function are generated on the computer using a 2D digital filter technique [22]. The technique considers the heights of a rough engineering surface as a time series of random process, and the 3D random surfaces can be generated through a linear transformation system of a series of random input signals. For the simulation, various Gaussian rough surfaces with different value of Ra are generated numerically. Figure 1c shows 3D Gaussian surface generated with the center line average roughness Ra = 0.14 lm, are the correlation length b* = 31.5 lm, skewness Sk = 0 and kurtosis K = 3. The computer-generated surface contains 256 9 256 data points. By using these data points as nodal points there will be 255 9 255 small patches of equal size within the simulated area. During multiple-asperity contact, the length scale of surface roughness is a factor which contributes to the scale dependence of the contact performances such as the real contact area and the contact pressure" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002236_2007-01-3740-Figure15-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002236_2007-01-3740-Figure15-1.png", "caption": "Figure 15: TE-IVT Concept", "texts": [], "surrounding_texts": [ "Torotrak\u2019s full toroidal traction drive technology has achieved series production in the Outdoor Power Equipment market validating the traction drive technology, the components and \u2018two roller\u2019 Variator design. The Variator has also been proven to be scaleable and configurable for higher torque and power applications. A number of arrangements have therefore been designed to satisfy the requirements for low cost, efficient transmission solutions in the constrictive packaging space of transverse FWD vehicle applications." ] }, { "image_filename": "designv11_61_0002897_tmag.2007.916494-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002897_tmag.2007.916494-Figure1-1.png", "caption": "Fig. 1. Definition of angles in initial magnetization process.", "texts": [ " The energy equation in the magnetic substance can be expressed as follows: (1) In this equation, we need the curve to calculate the total energy . The curve includes the effect of the shape of materials. To overcome this inconvenience, we divide the integration term into the demagnetizing field term and the effective magnetic field term as follows: (2) where is the demagnetizing field and is the effective field [4]. To calculate the second term of the right-hand side of (2), the curve is required. This curve does not include the effect of the shape of materials. Fig. 1 shows the definition of angles of the magnetization vector and the magnetic field strength vector in the initial magnetization process. Accordingly, the demagnetizing field can be written as follows: (3) 0018-9464/$25.00 \u00a9 2008 IEEE where and are the demagnetizing factors: denotes the major axis and and denote the minor axes. The magnetizing energy can be expressed as follows: (4) where and are the magnetic susceptibility in the easy axis and the hard axes. Therefore the energy equation in the magnetic material can be written by (5) Accordingly, the conditions of and in each substance (finite element) can be calculated by the following simultaneous equations: (6) (7) B" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001874_ijcat.2007.015267-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001874_ijcat.2007.015267-Figure5-1.png", "caption": "Figure 5 Test result of fixture eccentric", "texts": [ "18 The following points can be noted: 1 The amount of eccentricity refers to the eccentric value between the centre of the claming fixture and the basic shaft of the machine tool. 2 Longitudinal eccentric refers to the eccentric in the flat, which is vertical to the upright column of the honing machine. Transverse eccentric is the reverse. We can conclude from the above that there is little influence on roundness, roundness and roughness. This result shows that, in a certain eccentric scope, the honing iron sleeve can obtain stable value in roundness, roundness and surface roughness. Using statistics chart law analysis, the result is shown as in Figure 5. After a week\u2019s inspection text in factory 856, the size precision of the workpiece is in the tolerance zone. The roundness and cylindricity is measured with pneumatic gauge. After a rough and a fine processing, the roughness reaches below Ra 0.1 \u00b5m. All of these meet the necessity. 1 The design of the prototype system is reasonable, the working frequency is 18.6 kHz and the oscillation amplitude reaches 10 \u00b5m. 2 Only one clamping when using micro power diamond oilstone strip to rough honing and micro honing a workpiece, gives a roughness of 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001573_acc.1995.531386-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001573_acc.1995.531386-Figure4-1.png", "caption": "Figure 4. Two Consecutive Bodies of the n-Body Robot", "texts": [ " The spatial velocity notation [5] is adopted in this paper to derive the generalized equations of motion of n-link robot manipulators. This notation is efficient in the derivation and computation of the equations of motion, because it provides a direct transition relation between the joint velocity and spatial velocity 161. 3.1 Recursive Kinematics in Spatial Vectors Let\u2019s consider a single-chain n -link robot manipulator, as shown in Figure 3. For simplicity, the base body (root) is assumed to be fixed at the origin of the global reference frame. Two consecutive bodies, say, body i and body i + 1, are shown in Figure 4. Both of the bodies have their own reference frames, denoted by XI, -y{ -z: and x\u2019 i + l - Y { + ~ - z;+~, respectively, located at their \u201cin-board\u2019\u2019 joints through which they are connected to their proximal bodies. Each joint is described by one or a set of generalized coordinates, depending on the degrees of freedom allowed on the joint. In this paper, for simplicity, a joint is assumed to be either revolute or transitional of one degree of freedom. In Figure 4, the generalized coordinate of the proximal joint of body i + 1, i.e., qi+l, is defined as the relative angle from the distal joint reference frame of body i , X\u201d - y\u201d - z \u201d , to the proximal joint reference frame of body i + 1 , x\u2019 - y\u2018 - z\u2018, which also is the body reference frame body i + 1. The generalized coordinate of translational joint is defined in similar way. Featherstone [5] defined a six-element column vector, called spatial velocity, for a body, say body i t as: where ri , ri, ai and are the Cartesian position, velocity, and angular velocity of body i , with respect to the global inertial reference frame" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003013_robot.2009.5152257-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003013_robot.2009.5152257-Figure1-1.png", "caption": "Fig. 1. Example of flexible circuit board", "texts": [ ". INTRODUCTION According to downsizing of various electronic devices such as notebook PCs, mobile phones, digital cameras, and so on, more film circuit boards or flexible circuit boards illustrated in Fig.1 are used instead of conventional hard circuit boards. It is difficult to assemble such flexible boards by a robot because they can be easily deformed during their manipulation process and they must be deformed appropriately in the final state. For example, the flexible circuit board shown in Fig.1-(a) must deform to the objective shape illustrated in Fig.1-(b) to install into the hinge part of a flip phone. Therefore, analysis and estimation of deformation of film/flexible circuit boards is required. In solid mechanics, the Kirchhoff theory for thin plates and the Reissner-Mindlin theory for thick plates have been used[1]. For very thin plates, the inextensional theory was proposed[2]. In this theory, it is assumed that the middle surface of a plate is inextensional, that is, the surface of the plate is developable. Displacement of plates can be calculated using FEM based on these theories", " As those parameters are defined continuously along the central axis, the maximum curvature direction can be determined arbitrarily. Consequently, our proposed model is suitable for representation of a deformed inextensible rectangular belt object. However, in that model, it is assumed that the maximum curvature direction at both ends is parallel to the central axis to simplify integration of potential energy. Such assumption restricts application of that model to any deformation of a belt object. Furthermore, film/flexible circuit boards have some curves, angles, and/or branches in general as shown in Fig.1. Therefore, in this paper, we extend the previous fishbone model and apply it to deformation of a belt object with angles. In this section, we explain the fishbone model to describe deformation of a rectangular belt object in 3D space and its modification. Let U and V be the length and the width of the object, respectively. Let u be the distance from one end of the object along the central axis in its longitudinal direction and let v be the distance from the central axis in a transverse direction of the object", " In this example, a belt object is wound helically around a cylinder. Then, at both ends, the rib line does not coincide with the transverse edge, that is, \u03b1(0) and \u03b1(U) are not equal to zero. This also implies that transverse curvatures \u03b4(0) and \u03b4(U) are not equal to zero. Our previous model proposed in [12] cannot be applied to such deformation. We can derive the deformed shape shown in Fig.6 by using the modified model in this paper. Some flexible circuit boards have angles like a polygonal line as shown in Fig.1 or branch into some parts. In this section, we apply the fishbone model to an angled and a branched belt object. Fig.7 shows a belt object with one angle. Let \u03bb be the angle between the spine line of the left part and that of the right part at point P(ub). To represent the angle of the object, Eulerian angles and transverse curvature of the left rectangular part ABGH, the central part BEFG, and the right rectangular part BCDE are defined separately, for example, \u03c9\u03b6(u) = \u03c9\u03b61(u) (0 \u2264 u \u2264 ua) , \u03c9\u03b62(u) (ua \u2264 u \u2264 uc) , \u03c9\u03b63(u) (uc \u2264 u \u2264 U) , (23) where a parameter with subscript 1 is for the left rectangular part, that with subscript 2 for the central part, and that with subscript 3 for the right rectangular part of the object, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002377_s12239-009-0084-3-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002377_s12239-009-0084-3-Figure1-1.png", "caption": "Figure 1. Space vector in the stationary abc, \u03b1\u03b2\u00e2, and the rotating xy coordinate frames.", "texts": [ " The torque equation will be further derived based on the phase quantities as: (4) Knowing that: (5) Then, the following equation will be obtained: (6) For the field oriented control, the electrical equations are projected from the three phase stationary frame to the two coordinates rotating frame. 2.2. Direct Current-Space-Vector Control for a BPM Machine The direct current-space-vector control presented in this paper utilizes the current vectors in a coordinate frame rotating synchronously (xy coordinate frame) as illustrated in Figure 1. The corresponding space-vector coordinates, qx(t) and qy(t), referring to the x- and y-axes, are obtained from the stationary coordinates by coordinate transformation as in (Chn and Soulard, 2003): (7) The xy current coordinates are then: (8) The electromagnetic torque of a BPM can be described by the interaction between the stator currents and the flux resulting from the stator currents (Krause, 1987): (9) where, , are the x- and y- axes stator fluxes. The stator fluxes are formulated as: , (10) where, Ls is the stator inductance and \u03c8x and \u03c8y are the xand y- axes flux linkages, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002001_12.709253-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002001_12.709253-Figure2-1.png", "caption": "Figure 2. Proxy-based haptic volume rendering: The proxy is constrained to move only perpendicular to the gradient vector", "texts": [ " The basic idea behind proxybased methods is that a virtual spring damper with damping constant k connects the haptic instrument to a virtual proxy. By the restriction of proxy movement forces are indirectly (by the spring) applied to the haptic instrument using Hooke\u2019s law f = \u2212k d (1) where d is the distance between proxy position xp and instrument position xt. While in haptic surface rendering planes defined by the surface normals restrict the proxy from moving through surfaces, in proxy based haptic volume rendering the proxy is constrained to move only perpendicular to the gradient vector at the proxy position \u2207V ( xp) (fig. 2). In detail, the force vector f from the spring simulation is split up in two orthogonal portions fN = f \u00b7 N\u0302 and fT = f \u00b7 T\u0302 where N\u0302 = \u2207V ( xp)\u2223\u2223\u2223 \u2207V ( xp) \u2223\u2223\u2223 and T\u0302 = d \u2212 N\u0302 ( d \u00b7 N\u0302 ) \u2223\u2223\u2223 d \u2212 N\u0302 ( d \u00b7 N\u0302 )\u2223\u2223\u2223 . (2) The new proxy position x \u2032 p is calculated in every time step (approx. 1000 times per sec.). The movement is split up into three steps that correspond to the simulation of penetrability (eq. 3), friction (eq. 4) and viscosity (eq. 5). x \u2032 p = { xp + N\u0302( d \u00b7 N\u0302 \u2212 TN/k) if TN < k( d \u00b7 N\u0302) xp otherwise (3) x \u2032 p = { xp + T\u0302 ( d \u00b7 T\u0302 \u2212 TT /k) if TT < k( d \u00b7 T\u0302 ) xp otherwise (4) x \u2032 p = { xp + R( XpO\u2212 xp) k| XpO\u2212 xp| if R/k < | XpO \u2212 xp| xpO otherwise (5) Proc" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000300_047174414x.ch15-Figure15.9-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000300_047174414x.ch15-Figure15.9-1.png", "caption": "Figure 15.9 Address-centric routing vs. data-centric routing [8]. (a) Address-centric routing; (b) data-centric routing.", "texts": [ " When data fusion is considered in conjunction with data gathering and dissemination, the conventional address-centric routing, which finds the shortest routes from sources to the sink, is no longer optimal. Instead, data-centric routing, which considers in-network aggregation along the routes from multiple sources to a sink, achieves better energy and bandwidth efficiency, especially when the number of sources is large, and/or when the sources are located closely to one another and far from the sink [8]. Figure 15.9 gives a simple illustration of datacentric routing versus address-centric routing. Source 1 chooses node A as the relaying node in address-centric routing, but node C as the relaying and data aggregation node in data-centric routing. As a result, a smaller number of packets are transmitted in data-centric routing. Existing research activities of data fusion can be categorized with respect to the following aspects: . Fusion Function. Data-fusion is generally applied for: (a) Basic Operations. The most basic operations for data fusion include: COUNT, MIN, MAX, SUM, and AVERAGE [26]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000569_msf.455-456.732-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000569_msf.455-456.732-Figure2-1.png", "caption": "Fig. 2. Experimental apparatus and springback.", "texts": [ " From the analyses of the yield surface shapes, it is possible to conclude that whilst the differences between the Hill48 and the D~L yield surfaces are generally small; the differences to the CB2001 yield surface are larger mainly in the neighbourhood of the equi-biaxial stress state. Table 1 shows the calculated anisotropy parameters. For comparison purposes, the parameters of the Hill48 yield criterion were also identified following the classical methodology (Hill 48 Ref.). The \u201cUnconstrained Cylindrical Bending\u201d benchmark was proposed at the Numisheet\u20192002 conference in order to investigate the springback phenomenon. Fig. 2a shows the main tools\u2019 dimensions. The total punch stroke is of 28.5 mm. Springback can be quantified by the difference between the bending angles after loading and unloading, as shown in Fig. 2b. Due to the geometrical symmetry, only one quarter of the total specimen was simulated. A FE mesh of 82\u00d7 30\u00d7 3 8-nodes hexahedron finite elements was used. To model the isotropic hardening, the Voce law without kinematic hardening was selected with the following parameters: 420.857 233.329 exp( 8.448 )pY \u03b5= \u2212 \u2212 . Simulations were carried out for specimens cut from three different angles (0, 45 and 90\u00ba). Table 2 summarizes the experimental [4] and numerical springback results. As one can see, the springback predicted with the CB2001 yield criterion is systematically closer to the experimental results than with the other yield criteria" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000117_135065003322620246-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000117_135065003322620246-Figure1-1.png", "caption": "Fig. 1 Examples of predicted journal orbits by various authors: (a) Ritchie [12] using an `improved\u2019 short bearing theory; (b) Fantino et al. [13] using a short bearing model and modelling the bearing elasticity using the FE method; (c) Pal et al. [23] using an FE oil \u00aelm model for a Ruston\u00b1Hornsby 6 VEB-X MkIII engine", "texts": [ " As a compromise, Ritchie [12] proposed an `optimized short bearing solution\u2019 that included a factor within the short bearing solution that modi\u00aeed the pressure distributions, particularly at high eccentricities, so that they were much closer to those calculated using a full solution. The results were compared with `exact\u2019 solutions based on the full Reynolds equation published by Campbell [7], and good agreement was shown. Ritchie\u2019s paper [12] was the \u00aerst to attempt to reconcile the advantages of an approximate solution with the accuracy of a full bearing solution. Examples of orbital trajectories published by various researchers are shown in Fig. 1, including results published by Ritchie (Fig. 1a). In 1983, Fantino et al. [13] devised a numerical `timemarching\u2019 solution based on a `short bearing\u2019 solution to Reynolds equation, but incorporated within the solution allowance for elastic deformations of the bearing owing to oil \u00aelm pressure forces. Theoretical results showed that the eccentricity ratio could be as high as 3.8 (i.e. journal displacement from the bearing centre by 3.8 times the nominal bearing clearance), thus implying substantial elastic deformation of the bearing. They also showed that the minimum oil \u00aelm thickness is reduced and that friction is slightly increased by bearing elasticity. An example of an orbital trajectory published by Fantino et al. [13] is shown in Fig. 1b. Also, in 1983 Martin [14] published a review of work undertaken on engine bearing design. He identi\u00aeed the areas in which most development work was taking place as being those concerned with oil \u00aelm history, inertia effects, bearing non-circularity, bearing \u00afexibility and lubricant supply port geometry. He commented that much of the work was being made easier with the more powerful computers that were then becoming more readily available. Moes et al. [15] continued the work initiated by Booker with his `mobility\u2019 method [2] by extending this to the analysis of diesel engine main bearings", " Results were published for connecting-rod bigend bearings for two engines, and results were compared with other data produced using the \u00aenite difference method [7, 14]. In the \u00aerst paper [21], results showed oil \u00aelm minimum thickness values that were considerably higher than those obtained by other researchers. Data were presented for the no-load condition (i.e. without allowance for gas pressure) and for a full-load condition, the respective maximum eccentricity ratios quoted being 0.55 and 0.57. The orbital trajectory of the journal was shown in reference [21] (see also Fig. 1c) relative to the engine axis, whereas most researchers show it relative to the connecting rod axis, and this makes it dif\u00aecult to compare the shape of the orbit with the results of other researchers. In the second of the two papers [22], the journal orbit shown more closely resembles those published by other researchers and has a maximum eccentricity ratio of about 0.9. Pal et al. [23] and Singh et al. [24] also describe results obtained using the \u00aenite element method. In reference [23] a journal locus using a Newtonian lubricant (see also Fig. 1c), similar to that published earlier [22], is shown to be similar in shape to one obtained when a non-Newtonian lubricant is used. However, minimum oil \u00aelm thickness is 17 per cent smaller when nonNewtonian lubricants are used. This is in marked contrast both to later theoretical results obtained by Roberts and Walker [25], showing that for a steadily loaded journal bearing the use of non-Newtonian lubricants does not adversely affect performance, and to experimental data collected by Williamson [26], showing that polymer-thickened oils increase bearing load capacity", " It is suggested, therefore, that further work be undertaken to tackle these issues. Many, if not all, of the most recent developments in bearing modelling capacity are incorporated within software utilized by leading bearing design organizations, and with modern powerful computers such organizations are able to run computer models of particular designs and loading conditions within a short time to meet the needs of their clients. Many such organizations also publish design guides (for example, those published by the Engineering Sciences Data Unit). Figure 1a is reprinted from Wear, Vol. 35, G. S. Ritchie, `The prediction of journal loci in dynamically loaded internal combustion engine bearings\u2019, pp. 291\u00b1297, 1975, with permission from Elsevier Science. Figure 1b is reprinted with permission from SAE 830307 # 1983 Society of Automotive Engineers. Figure 1c is reprinted from STLE Tribology Transactions, Vol. 31, 2, by Pal et al. `Analysis of big end bearing having non-Newtonian \u00afuids\u2019, pp. 296\u00b1302, 1987, with permission from STLE. Figure 2a is reprinted with permission from SAE 690114 # 1969 Society of Automotive Engineers. Figure 2c is reprinted from Tribology International, Vol. 17, Dede and Holmes, `On prediction and experimental assessment of engine bearing performance\u2019, pp. 251\u00b1258, 1984, with permission from Elsevier Science. Figure 2d is reprinted from STLE Tribology Transactions, Vol" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002862_isccsp.2008.4537214-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002862_isccsp.2008.4537214-Figure1-1.png", "caption": "Figure 1: Kinematics coordinate system assignments", "texts": [], "surrounding_texts": [ "The fuzzy controller which generates the virtual force toward the destination accepts the rotation angel ( ) and side distance ( s ) as its input parameters. The left or right side distance is calculated considering the direction of rotation of the AGV and the angel between the direction in which it is moving and direction toward the destination. Figures (2), (3) and (4) respectively depict membership functions for difference of AGV and destination, Side distance from obstacle and rotation angel. These figures show the membership functions for AGV moving toward the destination. Figure (5) shows the FIS used for guidance of the AGV toward the destination. Table.1 contains the fuzzy rules for guidance. TABLE1: ROTATION ANGEL MEMBERSHIP FUNCTIONS s Zero Small Medium Large Very Large Far Zero Low Mean High Very High Near Zero Zero Zero Zero Zero" ] }, { "image_filename": "designv11_61_0002222_978-0-387-46283-7-Figure4.28-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002222_978-0-387-46283-7-Figure4.28-1.png", "caption": "Fig. 4.28 Schematic diagrams indicated the experimental process for fabrication of TiO2 nanobowl arrays. (a) arrangement of polystyrene spheres, nominally 505 nm in diameter sapphire substrate, (b) growth of 20 nm TiO2 with 200 cycles, (c) ion beam milling to remove upper half of hemisphere, (d) dissolve polystyrene in toluene [79]", "texts": [ " Prior to growth, a 0.2 nm-thick film of alumina was deposited. The average deposition of tungsten was 0.72 nm per cycle, eventually forming a total thickness of 7 nm. A template method based upon nano- or microspheres can use the spheres as a mask to expose regions of the substrate to direct film growth [78]. Nanobowl films of TiO2, formed by ALD over a self-assembled monolayer of 505 nm-diameter polystyrene spheres, were removed in toluene following an ion beam etching of the spheres\u2019 top layer [79]. Figure 4.28 summarizes the fabrication process. High temperature processing was subsequently used to transform the amorphous material into polycrystalline titania. Size-tunable metal nanostructures were also defined utilizing these periodic arrays. Here the process is modified to include a PMMA sacrificial film [80], as depicted in Fig. 4.29. The junction between the PMMA film and the polystyrene sphere provides a good contact. After ALD coating, the nanospheres and PMMA were removed to produce a free-standing sheet of nanobowls, each with a small hole at the center" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002839_robio.2009.5420393-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002839_robio.2009.5420393-Figure6-1.png", "caption": "Fig. 6. Human index finger meshing", "texts": [ " Parameters used in simulation are shown in Table I. Young\u2019s modulus and Poisson\u2019s ratio have already been shown in the data sheet but the viscous parameter of this type of silicone is not identified. Therefore, we used 100 Pa\u00b7s as an approximate value for viscous modulus in this simulation. We utilized previous researches [4], [11] and MRI images (see Fig. 5) to determine parameters for human fingers, except for the viscous modulus shown in Table II and the shape of the mesh model of a human finger as shown in Fig. 6. This object consists of 91 nodes and 158 triangles. We used MSCvisualNastran4D, a CAE software program, to determine the mesh data of this object and imported these data into our simulation. The shape of this finger can be approximated by an ellipse, with major and minor diameters of 17.44 mm and 13.60 mm, respectively. As an approximate value, we set the depth of this object at 10 mm. We also considered the complex construction of a human finger, which is composed of epidermis, dermis, tissue under the skin, nail and bone, each of which has different parameters" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001733_cacsd-cca-isic.2006.4777101-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001733_cacsd-cca-isic.2006.4777101-Figure2-1.png", "caption": "Fig. 2. Satellite configuration.", "texts": [ " This enables the scheduled controller to have a small number of knots without conservatism. Furthermore, if we set G\u0393 \u2217 = G (a common matrix) in Theorem 1 and a solution to (8)\u2013(11) is obtained, the derived controller matrices Ac(\u03b8) and Cc(\u03b8) are simply piecewise-linear functions of \u03b8 since G\u0393 S becomes parameter-independent. In the next section, a design example with D\u0393 = {\u03b8, \u03b8}(N\u0393 = 0) and G\u0393 0 = G\u0393 1 is shown, which leads to a linearly interpolated gain scheduling controller. A. Satellite Model Figure 2 shows the large flexible satellite considered in this paper. The satellite has a typical configuration consisting of two solar array paddles and two large deployable antenna reflectors. The solar paddles are deployed in the pitch direction and rotate 360 degrees per day so that each wing constantly faces the Sun. The equation of motion for rigid-body rotation and the vibration equation of flexible components (solar paddles and antenna reflectors) are described as [15]: J(\u03b8)\u03a8\u0308 + \u2211 i Pi(\u03b8)\u03bc\u0308(i) = u, (12) PT i (\u03b8)\u03a8\u0308 + \u03bc\u0308(i) + \u03a92 i \u03bc (i) = 0, (i = n, s, a, b), (13) y = [ \u03a8 \u03a8\u0307 ] " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000285_cdc.2003.1271669-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000285_cdc.2003.1271669-Figure3-1.png", "caption": "Fig. 3. Devil stick model on the vertical plane", "texts": [ " In $U, we consider a model of the devil stick and make some assumptions, and derive a state equation of the devil stick. Control objective to rotate the center stick continuously is stated in $111, and based on it we try to control by output zeroing in SIV. The output function is derived from a good juggler\u2019s performance. And also we consider the angular velocity which realizes a stable rotary motion with analysis of zero dynamics in $IV. Finally, \u00a7V concludes this paper. 11. DEVIL STICK MODEL In this paper, we consider the devil stick model shown in Fig. 3 and use the parameters listed in Table I. In this model, the center stick moves on vertical plane and crosses the hand stick at right angles. (r , 0) and Cp represent a position of the center of gravity (COG) of the center stick in a polar coordinate and an attitude angle of the center stick respectively. F = [Ft, F,lT = [F,., F0IT is an applied force by the hand stick. A. Contact between the center stick and the hand stick We make the following assumptions about a contact between the center stick and the hand stick" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002189_3-211-38053-1_9-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002189_3-211-38053-1_9-Figure3-1.png", "caption": "FIGURE 3. Scheme and photo of the feeder (3 phases)", "texts": [ " ' i _ a 0 ^ E ~ x: 0 iy + - .D^ O CD 0) CO 0 CO II D ) CD \u2022D ^ O E ^ ? ^ 0 \u00a3 i: p ^ 0 CO 0 is CD CD r ^ CD) 0 _> y CD \"D CD 8 CO CD c 0 \u2022+= 0 CO 0 0 CD J O O) c CO CD CD CO 0 ^ Q. CD \" ^ ^ CO > Q^ Assembly of Microproducts: State of the Art and New Solutions 107 The contactless feeder is based on the property of materials to be attracted, by charge induction, towards regions with a higher electric field in the same way as a dielectric/conductive plate is attracted inside the two charged plates of a capacitor. The feeder (Figure 3) consists of a high voltage supply (1) that, by a PLC based switching system (5), supplies a series of V-shaped paral lel wires (the electrodes (4)) mounted over a conducting and vibrating platform (2). The vibration has been used to make the microparts free to move on the working surface following the existing potential fields (in this case electrostatic). Therefore, in order to minimize the energy of the whole system, the component is attracted and oriented under the charged electrode. The electrodes are charged one after the other according to a proper sequence that creates a step ping electrostatic field (that is a travelling capacitor) able to attract and transport the microparts when aligned", " Fassi Legend Electrostatic potential determined \" by the active electrodes Electrostatic field determined by the active electrodes Data Parallel straight elecrodes of infinite lenght Electrode distance p=2mm h=1.5mm Electric potential calculated on the plane z=0.1mm FIGURE 4. Evaluated electrostatic potential and field of the feeder Two CCD cameras have been used to monitor and analyse the part movement from two direc tions: the first one has its axis parallel to the x direction, the other to the y one (Figure 3). Each frame has been captured by a tailored software, based on a commercial vision system, has been the component recognised and its centre of gravity measured and plotted. This configuration allowed to observe the component behaviour on both x (feeding direction) and y (centring effect) corresponding to the variation of the process parameters. The typical results of feeding micro parts along a line are shown in Figure 5. The feeding of a steel microgear (mass=3.1mg, d= 1.20mm, 1=1.05 mm) is shown in FIGURE 5a, while the same dielectric sphere ((|)=0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000183_s0020-7462(02)00155-5-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000183_s0020-7462(02)00155-5-Figure2-1.png", "caption": "Fig. 2. Views of the :oppy disk.", "texts": [ " Once discretized, an ordinary diEerential equation integrator such as Gear\u2019s method can be applied [25\u201327,29,30]. What is e cacious and e cient about the equations as formulated is that viable solutions can be found without resorting to a large degree of freedom model. This bucks the current trend of \u201ckilling\u201d problems with degrees of freedom. In this section a simulation of a simpli2ed version of the above device is provided. Essen- tially the handle has been removed. The device now resembles a :oppy disk rolling on a plane. The situation is depicted in Fig. 2. The geometry of a 3:5 in computer :oppy disk was chosen. The hub of the disk was assumed loaded with a US quarter dollar, it just 2ts. The properties chosen are shown in Table 2. Consistent CGS units were utilized and thus the units are not reported. Those constants not reported in the table but needed in the simulation were generated via these parameters. The equations of motion were generated using the above reported procedure with the aid of computer symbolic manipulation tools [29]. Viscous damping was added to the disk 2eld equation and damping was added to the hub motion" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002412_icems.2009.5382984-Figure14-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002412_icems.2009.5382984-Figure14-1.png", "caption": "Fig. 14. Measuring points", "texts": [ " The effect of the measuring point location was verified by using the analysis result. We were able to confirm a huge acceleration by rotating the measuring point 45 degrees at 3040 Hz and a large amount of electromagnetic vibrations are caused by the resonance phenomenon. The result is shown Fig. 13. By using the single logarithmic graph, the acceleration at the resonant frequency seems much larger than others. Let us confirm the validity of the effect by changing the measuring point. The measuring point on deformation diagram at 3040Hz is shown in Fig. 14. The resonant frequencies are differed between the measurement and the analysis result because of the difference of the natural frequency. Accurately modeling the coil is more difficult than the other components, and determining the mechanical characteristics is also difficult because the coil is consist of the wires, the insulators, and so on. In addition, the coil has the air between wires. Therefore we investigate the influences of the modeling differences on the natural frequency. We compare the natural frequency of the stator consist of the stator core, the case, and the coil" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000839_cdc.2004.1428720-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000839_cdc.2004.1428720-Figure1-1.png", "caption": "Fig. 1. Coordinate frames", "texts": [ " The xe and ye axes are directed toward the North and the East, respectively, while the ze axis points downward. This frame is assumed to be inertial, because when studying the motion of marine vehicles the acceleration at a point on the earth due to the earth\u2019s rotation is generally considered negligible. The vessel\u2019s configuration in the EFF is \u03b7(t) [x(t), y(t), \u03c8(t)]T, t \u2265 0, (5) where x(t) \u2208 R and y(t) \u2208 R describe the distance traveled along the xe and ye directions respectively, and \u03c8(t) \u2208 R describes the rotation about the ze axis (see Fig. 1). The Body Fixed Frame (BFF) has its origin fixed at the vehicle\u2019s center of gravity (CG), the xb axis points forward, the yb axis starboard, and the zb axis downward (see Fig. 1). The vessel\u2019s velocity is defined in the BFF as \u03bd(t) [u(t), v(t), r(t)]T, t \u2265 0, (6) where u(t) \u2208 R and v(t) \u2208 R are the components of the absolute velocity in the xb and yb directions respectively, and r(t) \u2208 R describes the angular velocity about the zb axis. The vectors \u03b7(t) and \u03bd(t) are related by the kinematic equation [9], \u03b7\u0307(t) = J(\u03b7(t))\u03bd(t), t \u2265 0, (7) where J(\u03b7) \u23a1 \u23a3 cos \u03c8 \u2212 sin \u03c8 0 sin \u03c8 cos \u03c8 0 0 0 1 \u23a4 \u23a6 (8) is the rotation matrix from the BFF to the EFF. Using the form introduced in [9] and the previous nota- tion, the surface vessel\u2019s equation of motion is given by M\u03bd\u0307(t)+C(\u03bd(t))\u03bd(t) +D(\u03bd(t))\u03bd(t) + g(\u03b7(t)) = \u03c4(t), t \u2265 0, (9) where M is the mass matrix, C(\u03bd(t)) contains Coriolis, centripetal and added-mass terms, D(\u03bd(t)) is the damping matrix, g(\u03b7(t)) is the vector of gravitational forces and moments, and \u03c4(t) is the input vector" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002673_05698197008972283-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002673_05698197008972283-Figure5-1.png", "caption": "Fig. 5- Theoretical load distribu tion.", "texts": [ " a - Effect of lead de viati on . 0.6 ~ 5.;;; c: 1i B -0 ~ z:..;;; :ii 0.2 :5 ~ 0 0 2 4 6 Axial Po,ition. x1m (dimen,ionless) Tooth ~ - Fig. 7-Effect of meshing position. ing posinon, profile modification, facewidth and helix angle. The following values are used in all cases: Yrl m = 2.7, SclSo = 1.5, 11m = 3, Aim = 2. No -load separa tion is based on N = 50, epn = 20\u00b0 an d mi t = 200. The other data applicable to each figure is given in Table l. Figures 5 and 9 show comparisons with the \" thin slice\" theory (A = 0). Figure 5 shows that there are two points where the load distribution along a major line of contact, with typical profile modification, calculated by the new theory differs fun damentally from that predicted by the D ow nl oa de d by [ M ou nt R oy al U ni ve rs ity ] at 0 3: 12 0 1 M ay 2 01 3 Tooth Loading and Static Behavior of Helical Gears 73 o. 6...------.--,---.--,---,----,--.---, Fig. II-Effect of helix angle. miSSIOn error increases. In the case shown, th e load reaches the tips of th e teeth when wi t is approximately 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003828_acc.2012.6315448-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003828_acc.2012.6315448-Figure4-1.png", "caption": "Fig. 4. Sketch of the sets B0, \u039b, S0 and P\u039b", "texts": [ " Since dV/d\u03c4 \u2264 \u2212\u03bbl(\u03b7 + \u03f5\u03b7)\u03f5\u03b7 < 0 in S1\\B0 and S0 is a level set of V satisfying S0 \u2287 B0, the system (24) that starts from an initial point in S1 enters S0 at some finite time and remains in S0. Furthermore, it can be shown that the system (24) converges to a set smaller than S0. The set \u039b is defined as \u039b = {x | V \u2264 V0 and e(r)T r e(r) v \u2264 \u03c3a} , (42) where \u03c3a = \u22121 + V0/(\u03b1h\u0304 + 1) . (43) Proposition 4. Under the conditions e) to h), the system described by (24) that starts from a point in S1 reaches the set \u039b defined by (42) at some finite time and remains in \u039b as long as r \u2208 \u2126c. Proof: First, we will show that B0 \u2286 \u039b \u2286 S0 (Fig. 4). From the definitions of \u039b and S0, it is obvious that \u039b \u2286 S0. We consider the other half of the inequality, B0 \u2286 \u039b. Since 1\u2212 (e(r)T r e (r) v )2 \u2264 (\u03b7+\u03f5\u03b7)2\u2212\u2225de (r) v /d\u03c4\u22252 in B0, any e (r) v in B0 is included in the set defined by Bev = B0\u2229{x|de (r) v /d\u03c4 = 0}. Since, from (26) and (39), (\u03b1h\u0304+1)(1+e (r)T r e (r) v ) \u2264 V0 for \u2200x \u2208 Bev , the same inequality holds for \u2200x \u2208 B0. Then, noting that (\u03b1h\u0304+1)(1+\u03c3a) = V0 from (43), we obtain that e (r)T r e (r) v \u2264 \u03c3a for \u2200x \u2208 B0. Therefore, B0 \u2286 \u039b. Next, we will examine the behavior of x in S0, because x in S1 enters S0 and remains in S0 as shown above" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002953_ijsurfse.2009.026607-Figure11-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002953_ijsurfse.2009.026607-Figure11-1.png", "caption": "Figure 11 Construction of separate type laser rotary encoder (Encoder2): (a) grating disk and (b) detecting unit (see online version for colours)", "texts": [ " It may be, therefore, very difficult to eliminate the only influence of the eccentricity of the encoders after measurement. It must be eliminated before or during measurement. Only TE measured without the disturbance of the eccentricity of the encoders can be used in various evaluations. For solving the problem of the eccentricity mentioned earlier, the authors utilised the opposite reading. The 2nd adopted encoder (Encoder2) is available for the opposite reading with help of its special construction. A photograph of Encoder2 is shown in Figure 11. The grating disk and the sensor unit for signal detection of the encoder are separated with each other. The separation has the advantage that no unreasonable force is applied between the grating disk and the sensor unit. In addition, the disturbance by eccentricity on output pulses is quite less because laser beams are applied to two symmetric positions with respect to the centre of rotation of the grating disk. Output signal from each position has the same amount of influence of the eccentricity of the grating disk but the sign of the output signal is opposite because its location is symmetric" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003392_9780470876541.ch5-Figure5.22-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003392_9780470876541.ch5-Figure5.22-1.png", "caption": "Figure 5.22 Constant torque region.", "texts": [ "21, which is the cross section of the ellipse and the circle. And the torque is depicted as a reciprocal proportion curve in the current plane as shown in Fig. 5.21. 5.4.4.3 Constant Torque Region (ve vb) If d-axis current for the maximum torque, which is the current at the crossing point of the ellipse and the circle, is larger than the rated value of d-axis current of the induction machine, then the d-axis current reference should be set as the rated value to prevent the magnetic saturation of the induction machine. That is the case of point A in Fig. 5.22, where the torque, Te1 may be obtained by ieds1, but it is larger than i e ds rate and it would result in the severe saturation of the magnetic circuit of the machine. In this case, i*ds should be set as ieds rate as (5.86) and the operating point should be B in the figure generating torque, Te2. Also, the maximum available torque is decided only by the maximum q-axis current, which is given by the current constraint as (5.87 ), and the maximum torque is always the same as Te2 in this region. Hence, the region is called the constant torque region" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000796_s00170-003-1681-3-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000796_s00170-003-1681-3-Figure1-1.png", "caption": "Fig. 1 The spatial parallel mechanism made up of PTT", "texts": [ " Since most spatial parallel mechanisms have the same branches, the solving of the inverse screws can be abbreviated to solve only one branch. The whole process of this methodology, in reality, is very simple. In the following section we enumerate a representative mechanism to demonstrate the simplicity and effectiveness to deal with the singularity of spatial parallel mechanisms with this methodology. 3.1 The architecture of the spatial parallel mechanism made up of PTT A novel spatial parallel mechanism, shown as Fig. 1, is made up of three PTT (1 prismatic joint and 2 Hooke joints) branches. When the drives are given, the positions of the sliders are definite. To study the singularity of the spatial parallel mechanism, what we should do is only to search the imbalanced positions when the drives of the sliders are given. Because of the symmetry of the branches, we can simplify the process of analysis. For example, we select the branch P1B1 as the study object and create the local Cartesian coordinate shown as Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000165_s0531-5131(03)00300-5-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000165_s0531-5131(03)00300-5-Figure1-1.png", "caption": "Fig. 1. Initial alignment of patient and robot. The arrows show the variable degrees of freedom of the operating table and the robot.", "texts": [ " In order to cope with the problem of finding the correct robot position relative to the patient position, we developed a method which automatically locates the optimal robot position relative to the patient. Therefore, we established the following criteria for an optimal alignment: Feasibility of the robot path; No proximity to singularities; No collisions of robot and patient; No collisions of robot arm segments; Consideration of surgical access path. The algorithm considers the degrees of freedom of the mobile platform on which the robot is mounted (x-, y-translation and z-rotation) as well as the degrees of freedom of the operating table (z-translation and x-, y-rotation) (Fig. 1). The presented algorithm is an integral part of a viewer tool in which the surgeon can load and display the surface model of the patient skull, generated from CT-data, and the planned trajectory. Additionally, the robot is visualized and simulated on the basis of its Denavit\u2013Hartenberg model. Thus, the physician can watch the robot moving along the planned cut in virtual reality before the real treatment takes place. Furthermore, he can relocate the robot and the patient towards each other and assess different alignments" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000841_icsmc.2004.1401038-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000841_icsmc.2004.1401038-Figure2-1.png", "caption": "Figure 2. Schematic of the Tricept-like manipulator", "texts": [ " Then, the degrees of freedom of a mechanism is generally govemed by the following mobility equation: For the manipulator shown in Fig.1. n = 7 , j = 8 . Applying equation ( I ) to the manipulator produces: F = 6 ( 7 - 8 - 1 ) + 2 ~ 3 + 3 ~ 1 + 3 ~ 2 = 3 . Hence, the manipulator is a 3-DOF mechanism. Due to the arrangement of the links and joints, the leg 3 constraints rotation about the z axis, and the manipulator has IWO translational degrees and one rotational degree as Tricept robot. 3 Inverse kinematics As shown in Fig.2, the reference frame 0 - x y z is attached to the fixed base at point 0, located at the center of the fixed platform. And another reference frame a-+' is attached to the moving platform at point I?', the center of the moving platform. q, ,q2 , and q3 are considered the actuated joints. Assume that the initial location of the moving platform coincides with the fvted frame and that the final location is obtained by a rotation of a about the y-ais, followed by a second rotation of p about the x-a-xis" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001843_9780470185834.ch14-Figure14.2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001843_9780470185834.ch14-Figure14.2-1.png", "caption": "FIGURE 14.2 Examples of cell cell interaction models used in interface tissue engineering: (a) Coculture of fibroblasts and osteoblasts with temporary agarose gel divider; (b) triculture of fibroblasts and osteoblasts with chondrocytes encapsulated in 3D hydrogel; (c) microscale coculture of fibroblasts and osteoblasts using microfluidic patterning techniques. FB: fibroblasts, OB: osteoblasts, CH: chondrocytes.", "texts": [ " Interpretation of these coculture results, however, must take into consideration any dilution effect on cell response due to mixed culture as well as any metabolic differences between the various cell types. 14.3.1.2 Temporary Dividers Cell cell contact can also be controlled by establishing physical barriers that are used to organize cell-seeding patterns in coculture (Fig. 14.1). The barrier may be removed later to permit cell migration and controlled cell cell physical contact. This model has been used to examine interactions between fibroblasts and hepatocytes [26], as well as between osteoblast and fibroblast as shown in Fig. 14.2a [22]. It has the advantage of been able to excise greater control of the extent of heterotypic and homotypic interactions, while permitting both physical contact and soluble factor interactions. This system is, however, experimentally more challenging as a complete seal between the individual cell compartments is required. Moreover, cell response and soluble-factor transport in this model is a function of the physical and chemical properties of the divider material utilized. 14.3.2.1 Segregated Coculture To prevent cell cell contact, a segregated coculture system may be established by first forming individual cultures of each cell type, and later cocultivate them in the same environment", " The application of these models to determine the effects of cell cell contact and soluble factors for interface tissue engineering are reviewed below. To determine the role of fibroblast and osteoblast interactions in fibrocartilage formation, a novel coculture model permitting both cell contact and paracrine interaction was reported byWang et al. [22]. This model was designed to emulate the in vivo condition where the tendon is in direct contact with bone tissue during graft healing following ACL reconstruction. Briefly, as shown in Fig. 14.2a, osteoblasts and fibroblasts were first seeded on opposing sides of a tissue culture well. The cells were separated by a hydrogel divider preformed in the center of the well. Once the cells reached confluence on each side, the divider was removed, allowing the osteoblasts and fibroblasts to migrate and interact within the interface region. The two cell types communicated through paracrine and autocrine factors, as well as eventual physical contact in the interface region. It was found that osteoblast fibroblast interaction led to a decrease in cell proliferation, a reduction in osteoblast-mediated mineralization, accompanied by an increase in the mineralization potential of fibroblasts [23]", " It was found that while the chondrocytes continued to synthesize collagen type II, proteoglycan deposition was significantly lower in coculture. Alkaline phosphatase activity remained unchanged in the osteoblasts, while their mineralization potential was significantly reduced due to coculture. These results suggest that osteoblast fibroblast and osteoblast chondrocyte interactions are key modulators of cell phenotypes. Recently, Wang et al. reported on a triculture model of fibroblasts, chondrocytes, and osteoblasts; the three cell types dominant in their respective region of the interface, namely ligament, fibrocartilage, and bone (Fig. 14.2b) [44]. To ensure their phenotypic spherical morphology, chondrocytes were encapsulated within the hydrogel divider used in the previous osteoblast fibroblast coculture model (Fig. 14.2a) [22]. Once again, a reduction in the proliferation of both osteoblasts and fibroblasts due to heterotypic cell interactions was found, while the number of chondrocytes remained relatively constant over time in the hydrogel. Triculture led to reduced osteoblastmediated mineralization, accompanied by increased fibroblast mineralization. The chondrocytes continued to produce proteoglycans and the expression of both collage n types I and II were detected in the interfacial region. This model was subsequently used to compare the effects of fibroblast osteoblast interaction on the response of interface-relevant cells such as bone marrow stromal cells and ligament fibroblasts [45]", " The primary advantage of the microfluidic model resides in its ability to provide high resolution spatial control of fluid flow and factor concentration, as well as the potential for long term microscale coculture. For interface tissue engineering, a microscale coculture model is also physiologically more relevant, as the native human ligament-to-bone interface spans merely 200 300 mm [8]. Tsai et al. recently investigated the effects of osteoblast and fibroblast micropatterning on cellular function and organization [23]. This microscale coculture model (Fig. 14.2c) was fabricated using soft lithography and replica molding, and utilized microfluidics to exert spatial control and cell patterning. Not surprisingly, differences in cell growth and differentiation were detected at the microscale when compared to the macroscale model [22]. These observations emphasize the importance of considering interaction scale in coculture models. In addition to the interaction scale, the effects of 2D and 3D substrates on cellular communications during coculture must be considered" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000646_006-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000646_006-Figure5-1.png", "caption": "Figure 5. Free-body diagram for the lower pulley. The lower pulley has negligible mass. T1", "texts": [], "surrounding_texts": [ "Let us first list the quantities with which we shall calculate. We have the three masses\u2014-m2, m3 and m5. The tension in the lower cord is T1 and that in the upper cord is T2. The force of the ceiling bracket on the upper pulley is P . The magnitudes of the accelerations of the three masses in the inertial laboratory frame are +a2, \u2212a3 and \u2212a5 respectively. (We take upward accelerations as positive and downward accelerations as negative.) The upward acceleration of the lower pulley is +a5. We shall use the symbol g for the acceleration due to gravity. There is a second frame of reference that we must also consider. This is the accelerated, noninertial frame in which the lower pulley is at rest. An observer at rest in this frame would conclude that m2 accelerates upward with an acceleration a and that m3 accelerates downward with an equal and opposite acceleration \u2212a. The magnitudes of these accelerations must be equal in the pulley rest frame; otherwise the cord would snap. The acceleration +a2 of m2 in the inertial frame equals the vector sum of its acceleration in the pulley frame and the acceleration +a5 of the pulley in the inertial frame: +a2 = a + a5 (1) and similarly for +a3, \u2212a3 = \u2212a + a5. (2) Now consider the appropriate free-body diagrams for m2, m3, m5 and the lower pulley (figures 2\u2013 5). From these diagrams we can construct the net force acting on each body and equate it with the product of the mass with the acceleration. For m2 T1 \u2212 m2g = m2a2 = m2(a + a5). (3) Figure 2. Free-body diagram for m2. T1 is the tension in the cord. T1 m2g m2 For m3 T1 \u2212 m3g = m3(\u2212a3) = m3(\u2212a + a5). (4) For m5 T2 \u2212 m5g = m5(\u2212a5). (5) Even though the lower pulley is massless, it still obeys Newton\u2019s second law. Since its mass is zero, the net force acting on it is also zero. T2 \u2212 2T1 = 0. (6) This establishes that T2 = 2T1. (7) Equations (3), (4), (5) and (7) contain four unknowns: \u2212a, a5, T1 and T2. These are readily 290 P H Y S I C S E D U C A T I O N May 2004 solved and yield the results: a = (10/49)g = 2.00 m s\u22122 a5 = g/49 = 0.20 m s\u22122 T1 = [(120/49) kg]g = 24 N, T2 = 2T1 = 48 N. From equations (1) and (2) we find the accelerations a2 and a3 in the inertial laboratory frame as a2 = 2.20 m s\u22122 and a3 = \u22121.80 m s\u22122. At this point we have solved the problem as set. However, the problem has considerably more pedagogic value than just the above solution. The further topics we shall treat include inertial forces, d\u2019Alembert\u2019s principle, the breakdown of Newton\u2019s laws in non-inertial frames, centre of mass motion and the answer to the question of imbalance between equal masses. Inertial forces, d\u2019Alembert\u2019s principle and accelerated frames Rearranging equations (3) and (4) we obtain T1 \u2212 m2g \u2212 m2a5 = m2a (8) and T1 \u2212 m3g \u2212 m3a5 = m3(\u2212a). (9) Although these equations look as though they express Newton\u2019s second law, they do not. The left side of each contains two real forces and a \u2018fictitious\u2019 force. These two fictitious forces, \u2212m2a5 and \u2212m3a5, are essentially reversed effective forces that enable us to describe the motion of the 2 and 3 kg masses in the frame in which the pulley is at rest. We interpret these as inertial forces, forces that arise when we deal with accelerated frames. This artifice, first introduced by d\u2019Alembert in the eighteenth century, is the origin of the centrifugal force. In uniform circular motion we write Fnet = \u2212mv2/r. (10) If, following d\u2019Alembert, we write Fnet + mv2/r = 0 (11) we reduce the problem to one of equilibrium by introducing the fictitious inertial force, +mv2/r , known as the centrifugal force. To see the effect of inertial forces let us return to equations (3) and (4). These apply to figures 2 and 3. The net force is the vector sum of the tension and the weight in each case. It equals the mass times the acceleration in the laboratory frame. However, the acceleration in the laboratory frame may also be written as the vector sum of the acceleration in the non-inertial pulley frame (a for m2 and \u2212a for m3) plus the acceleration a5 of the pulley itself. This is how we arrive at equations (8) and (9). In words we can state equation (8) as follows. The net force on m2 (i.e. T1 \u2212 m2g) plus the inertial force (\u2212m2a5) equals the mass times May 2004 P H Y S I C S E D U C A T I O N 291 the acceleration in the non-inertial frame. It is only in the inertial frame of the laboratory that the net force equals the mass m2 times its acceleration in the laboratory frame a2 (a2 = a + a5). In other words, Newton\u2019s second law does not hold when applied to non-inertial frames. To describe motion in non-inertial frames we must introduce inertial forces. It is unfortunate that d\u2019Alembert\u2019s principle seems to have fallen into neglect in recent years. It is an extremely useful idea in classical mechanics that illuminates the meaning of inertial and noninertial frames. Though we shan\u2019t discuss it here, it is also important in the development of the concept of virtual work. For a fuller exposition of the principle see Goldstein [2] and Sommerfeld [3]." ] }, { "image_filename": "designv11_61_0002588_s10015-009-0707-9-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002588_s10015-009-0707-9-Figure4-1.png", "caption": "Fig. 4. The existence of two candidates for the solution", "texts": [ " 6, the component bi1 (or bi2) is represented as b P b Q b P b Qi i i i i i i i1 1 2 1 2 2 1 2= + = +( )or (7) where pi1,pi2, Qi1, and Qi2 are coeffi cient vectors. Thus, the unknown vector Bi = (bi1, bi2,0) can be derived to solve the quadric Eq. 4 with Eq. 7. Since the bottom end of link vector Bi can be derived, the link vector Li can also be derived by Eq. 2. 2.4 Condition of the solution In the previous section, we presented the derivation of the position of the bottom end of a link. This method provides two solutions for each link since it is a derivation from a quadratic equation. This situation is shown in Fig. 4. To avoid the three links interfering with each other, and from the structural constraint that the link cannot revolve toward the platform from the vertical line through the upper revolute joint, we choose the solution which is outside the vertical line. This section gives the numerical simulation results based on our analysis. For this simulation, the length of each link was set to 250 mm, and the three revolute joints were set at 30\u00b0 in the world reference frame at intervals of 120\u00b0 on the circumference of a circle with radius 31" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002336_1.2999933-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002336_1.2999933-Figure3-1.png", "caption": "FIGURE 3. Sketch of the \"elongated\" regime. Typical view of melt pool (scale: 1mm) and cross-section (scale: 0.5mm).", "texts": [ " More interestingly, when the ejected vapour plume collides with the melt pool and lifts it, one can observe the corresponding local heating of the liquid surface by this energetic heated vapour plume. So the vapour plumes not only transfers impulse momentum, but also energy probably due to its rather high temperature. c) Welding Speeds Between 9 to 11 m/min: \"Elongated Keyhole\" Regime. This regime is observed for welding speeds ranging from 9 to 11 m/min. It is characterized by a keyhole that is elongated, with its maximum length of about 2 mm, observed at 11 m/min (see Fig. 3). However its length presents also some fluctuations, but the resulting liquid oscillations are much less intense than previously, and the height of the induced swellings is much smaller. It is interesting to note that this elongated keyhole shows two characteristic zones that are heated: the first one corresponds of course to the inclined keyhole front wall, common to all regimes, and the second one is located at the rear end of this elongated keyhole, inside the melt pool. Moreover, vapour plume seems to be also emitted from the second heated spot and so directed frontward" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001951_20070822-3-za-2920.00059-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001951_20070822-3-za-2920.00059-Figure1-1.png", "caption": "Fig. 1. PVTOL aircraft.", "texts": [ " Using standard aeronautic conventions the equations of motion are given by y\u0308 = u1 sin \u03d5\u2212 \u03b5PVTOL u2 cos\u03d5 z\u0308 = \u2212u1 cos \u03d5\u2212 \u03b5PVTOL u2 sin\u03d5 + g \u03d5\u0308 = u2. (1) The aircraft state is given by the position (y, z) of the center of gravity, the roll angle \u03d5 and the respective velocities y\u0307, z\u0307 and \u03d5\u0307. The control inputs u1 and u2 are respectively the vertical thrust force and the rolling moment. The rolling moment u2 generates also a lateral force (the lift forces are not perpendicular to the wings) and \u03b5PVTOL is the coupling coefficient. Finally g is the gravity acceleration. In Figure 1 the PVTOL aircraft with the reference system and the inputs is shown. In (Hauser et al., 1992) the PVTOL was shown to be input-output linearizable when \u03b5PVTOL = 0, while in (P.Martin et al., 1996) it was shown that, when \u03b5PVTOL is not zero, suitable outputs (flat outputs) may be found, such that the system can be feedback linearized by means of dynamic extension. Using as flat outputs yf = y + \u03b5PVTOL sin(\u03d5) zf = z + \u03b5PVTOL cos(\u03d5) it can be shown, after some straightforward calculations, that for all \u03d5\u0307 and u1 such that u\u03031 6= 0, u\u03031 = u1 \u2212 \u03b5PVTOL \u03d5\u03072, the system is feedback linearizable and in fact equivalent to the two dimensional forth order integrator y (4) f = v1, z (4) f = v2 with suitable expressions for v1 and v2" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002542_00029890.2009.11920919-Figure14-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002542_00029890.2009.11920919-Figure14-1.png", "caption": "Figure 14. (a) Parallel zig lines and parallel zag lines through the hinges of a zigzag trammel. (b) A line through the rightmost rod RT intersects the y axis at A.", "texts": [ " Now let the free ends T and Z slide along the respective coordinate axes as indicated in Figures 13b, c, with all rods making equal angles with the x axis. We call this a zigzag trammel if the number of rods is greater than 1. In the examples shown, angle \u03b1 is allowed to increase from 0 to \u03c0/2, so the rightmost rod will stay in the first quadrant. For any fixed \u03b1 > 0, an ant starting at T and walking along the zigzag trammel toward Z decreases its x coordinate monotonically, but its y coordinate is piecewise monotonic, alternately increasing or decreasing as it crosses the hinges. Now refer to Figure 14. Any line parallel to the rightmost rod RT is called a zig line, while those parallel to the rod adjacent to RT are called zag lines. Because all rods make equal angles with the x axis, they generate a finite set of parallel zig lines and another finite set of parallel zag lines. Each hinge is the intersection of a zig line and a zag line, as indicated in Figure 14a. The configuration is like a movable latticework in which any two adjacent rods determine a parallelogram. In Figure 14b, A denotes the point at which a line through the rightmost rod RT intersects the y axis. By parallel projection suggested by Figure 14a, it follows that the length of AT is the sum of the lengths of the rods of the zigzag trammel. Consequently, we have the following property: Property 1. As a zigzag trammel ZT of length L slides along the axes, a line segment AT through the rightmost rod RT, from the y axis to T , is a standard trammel of length L sliding along the same axes, regardless of the number of rods and their relative sizes. Applications to folding doors. Zigzag trammels can be realized physically as folding doors. Examples are those treated in [5], where the left endpoint Z is kept fixed", " If the two rods joining alternate hinges of a zigzag trammel always have equal lengths, they form an isosceles triangle whose vertex is at the intermediate hinge. This configuration occurs when closet doors are divided into two or more hinged panels that fold or unfold as the door is opened or closed. Equation (12) can be used in particular to calculate the floor area swept by the rightmost panel of a door of n equal-sized panels by taking b = L/n and a = L(n \u2212 1)/n. We omit the details. A zigzag trammel has other interesting properties. For example, consider the rod RQ in Figure 14b adjacent to the rightmost rod RT. Let S denote the point where RQ or its extension to point Y on the y axis intersects the x axis. Then we have: Property 2. As a zigzag trammel ZT of length L slides along the axes, the point S remains on the x axis and SY has constant length L \u2212 2RT, so segment SY is a standard trammel sliding along the same axes. Proof. During the entire motion, triangle TRS is isosceles, with RS = RT, so S remains on the x axis. Also SY = YR \u2212 SR = (L \u2212 RT) \u2212 SR = L \u2212 2RT. The same analysis applies to each of the rods of a zigzag trammel. A line through the rod extended to meet the y axis produces a standard \u201cmini-trammel\u201d whose ends slide along the same axes. Consequently, each point on such a mini-trammel traces its own ellipse. It also has its own astroid as an envelope. One of them is shown by the dashed curve in Figure 14b. The next theorem tells how to determine the length of each such mini-trammel. Theorem 4. Consider a zigzag trammel of length L with rods of lengths z1, z2, . . . , zn, labeled from right to left. Then the segment between the x and y axes of the line through the i th rod is a standard trammel whose length Li is given as follows: for odd i (zig lines), L1 = L , L2k+1 = L \u2212 2(z2 + z4 + \u00b7 \u00b7 \u00b7 + z2k), k \u2265 1, and for even i (zag lines), L2k = L \u2212 2(z1 + z3 + \u00b7 \u00b7 \u00b7 + z2k\u22121), k \u2265 1. An inductive proof can be given, which is omitted" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003013_robot.2009.5152257-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003013_robot.2009.5152257-Figure3-1.png", "caption": "Fig. 3. Fishbone model", "texts": [ " In this paper, we assume that a belt object is inextensible. Then, the deformed shape of the object corresponds to a developable surface. It means that the object bends in direction d1 and it is not deformed in direction d2. Namely, a line the direction of which coincides with direction d2 is kept straight after deformation. In our studies, the central axis in a longitudinal direction of the object is referred to as the spine line and a line with zero curvature at a point on the object is referred to as a rib line as shown in Fig.3. We assume that bending and torsion of the spine line and direction of the rib line at each point specifies deformation of a belt object. This model is referred to as the fishbone model. Let \u03b1(u, 0) be rib angle, which is the angle between the spine line and direction d1 as shown in Fig.4-(a). Consequently, the shape of a straight belt object can be represented using five variables: \u03c9\u03be(u, 0), \u03c9\u03b7(u, 0), \u03c9\u03b6(u, 0), \u03b4(u, 0), and \u03b1(u, 0). Let us consider conditions five variables must satisfy so that the surface of a belt object is developable" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002748_ls.82-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002748_ls.82-Figure6-1.png", "caption": "Figure 6. Photoelastic observation apparatus.", "texts": [ " Moreover, tensile stress is induced along the interface of expanding bubble as a result of the viscoelastic solid transition. This enables the detection of the viscoelastic solid transition by photoelastic means in a dark fi eld of polarised light. The viscoelastic solid transition temperature of base oil P-N-0 in this method is TVE0 = \u221272\u00b0C. The photoelastic means was used to study the viscoelastic solid transition temperature under high pressure using DAC. A schematic diagram of the apparatus is shown in Figure 6. A drop of oil enclosed between a diamond anvil and oil in gasket contracts at high pressure. Above the glass transition, the solid-like behaviour of the lubricant can lead to non-hydrostatic stress in DAC.16 The difference in principal stress is induced in lubricant as a result of the viscoelastic solid transition. This enables the detection of the viscoelastic solid transition by photoelastic means in a dark fi eld of polarised light. Figure 7 shows the photoelastic pattern caused by increasing pressure in DAC", " The pressure\u2013temperature relation of viscoelastic solid transition temperature TVE was represented in the Yasutomi free volume model.17 T T A A pVE VE= + +( )0 1 21ln (2) The parameters obtained are A1 = 260.9\u00b0C and A2 = 0.549 GPa\u22121. The viscoelastic solid transition pressure from high-pressure density measurement at \u22127\u00b0C is 0.60 GPa, and the predicted value from equation (2) is 0.52 GPa. A result of an experiment of DAC and a result of an experiment of highpressure density measurement were corresponded. Again, the DAC experiment was considered without the analyser and polariser in the set-up (Figure 6) for P-A-10. A light was passed through the lubricant and gradually increased the pressure. Pictures Copyright \u00a9 2009 John Wiley & Sons, Ltd. Lubrication Science 2009; 21: 183\u2013192 DOI: 10.1002/ls were captured at different pressures, and the intensity of illuminance were measured by an illuminance meter. At the beginning of the experiment, light intensity was high and almost constant upon certain pressure, as shown in Figure 10. This phenomenon was introduced as the constant sol-state of the lubricant at the above condition" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002342_978-3-642-00196-3_53-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002342_978-3-642-00196-3_53-Figure4-1.png", "caption": "Fig. 4. FoV limits", "texts": [ " Handling Visibility Limits In general, there are two kinds of sensory limitations : limits on the maximum and minimum sensing range, range edges, and limits on the field of view, FoV edges. As these edges cannot be eliminated by the robot itself, the best strategy would be to delay the target\u2019s escape. A measure of this delay, the time taken by the target to escape through a gap (G), is called escape time (tesc). The robot then chooses actions that maximizes tesc, where we define tesc by, tesc = dist(T ,G) rel.vel(T,G) Field of View Limits: The FoV can be modeled as an annular sector with the robot at the center, with the visibility spanning from \u03b8min to \u03b8max, Figure 4. The only way to manipulate G is by rotating the visibility sensor towards the target. In case the visibility sensor has an additional degree of freedom over the robot, e.g. a pan mechanism, the angular velocity of the panning, \u03c9R is the action of the robot. On the other hand, if the sensor is attached rigidly to the robot base, the turning of the robot itself acts to rotate G. In that case we treat \u03c9R as the rotation of the robot. As we deal with the angular motion, it is natural to derive tesc using these rotational parameters" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002984_s00170-008-1573-7-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002984_s00170-008-1573-7-Figure1-1.png", "caption": "Fig. 1 Stewart platform", "texts": [ " These machine tools are not capable of delivering the required product quality and reduced manufacturing cost due to the factors such as poor rigidity, low pay load capacity, high inertia, and high vibration. Machine tools with closed kinematic chains and parallel actuators are being intensively researched as an alternate for open kinematic chain machine tools. These parallel kinematic machine tools have the advantages of higher stiffness, higher pay load capacity, and lower inertia to the manipulation problem [2]. Such parallel kinematic machine tools can meet the high demands on machining accuracy. Based on General Stewart platform [3], as shown in Fig. 1, many parallel machine tools are being built. These are generally known as hexapod machine tools, the schematic of which is shown in Fig. 2. In spite of the advantages possessed by these hexapod machine tools, they have a major drawback, i.e., the stiffness and workspace are highly dynamic and vary with the configuration of the machine tool structure. Because of this, locating the work piece optimally is very difficult. So, using a hexapod machine tool is not as simple as that of conventional machine tools [4]", " Further, the objective is also to plot 2D workspace slices, which would help the operator to locate the work piece optimally and conveniently. For constructing the hexapod machine tool, inverted Stewart platform is taken as the model. Spindle motor is fixed to the mobile platform. By varying the length of the legs, the position of cutting tool can be controlled. Since hexapod machine tool is an inverted Stewart platform, kinematics and stiffness of the hexapod machine tool will be as that of Stewart platform. The Stewart platform shown in Fig. 1 is a spatial 6-dof, 6SPS parallel manipulator. It consists of six identical extensible legs, connected to the mobile platform and fixed base by spherical joints at points Bi and Ai, where i=1,2, 3,\u20266. The legs are made up of two elements connected by a prismatic joint. Two Cartesian coordinate systems are attached to the mobile platform and fixed base. The position vector p of the centroid P and rotation matrix ARB of the mobile platform describe the transformation from the mobile platform to the fixed base", " Stiffness of the hexapod machine tool is governed by Jacobian, which is defined as the matrix that transforms the joint rates in the actuator space to the velocity state in the end-effector space. End effector velocity state expressed with respect to centroid P is given by, x \u00bc vp wB \u00f03\u00de where vp is the linear velocity of centroid P, and \u03c9B is the angular velocity of the mobile platform. Input joint rates are q \u00bc d 1 d 2 d 3 d 4 d 5 d 6 h iT ; \u00f04\u00de where d 1 ; d 2 . . . d 6 are sliding velocities of corresponding legs. The Jacobian matrix can be derived by formulating velocity loop-closure equation for each leg as shown in Fig. 1. p \u00bc ARBbi \u00bc ai \u00fe disi; \u00f06\u00de where bi is position vector of spherical joint Bi with respect to P, ai is position vector of spherical joint Ai with respect to the fixed base centre O, di is the length of the ith leg, and si is the unit vector pointing from Ai to Bi. Differentiating Eq. 6 with respect to time, vp \u00fe wB ARBbi \u00bc diwi si \u00fe d isi; \u00f07\u00de where \u03c9i is the angular velocity of ith leg with respect to the fixed base. To eliminate \u03c9i, Eq. 7 is dot-multiplied by si on both sides: sivp \u00fe ARBbi si wB \u00bc d i: \u00f08\u00de When assembled in the matrix form, Jx x \u00bc Jq q : \u00f09\u00de Where Jx \u00bc sT1 b1 s1\u00f0 \u00deT sT2 b2 s2\u00f0 \u00deT sT6 b6 s6\u00f0 \u00deT 2 66664 3 77775 \u00f010\u00de Jq \u00bc I 6 6 identify matrix\u00f0 \u00de: \u00f011\u00de Therefore, overall Jacobian matrix is as follows: J \u00bc J 1 q Jx \u00bc sT1 b1 s1\u00f0 \u00deT sT2 b2 s2\u00f0 \u00deT sT6 b6 s6\u00f0 \u00deT 2 66664 3 77775 : \u00f012\u00de The stiffness of the Stewart platform manipulator is given by, K \u00bc JTcJ ; \u00f013\u00de where c \u00bc diag k1; k2 " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002013_robot.2007.363148-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002013_robot.2007.363148-Figure1-1.png", "caption": "Fig. 1: (a) 3UPS spherical wrist; (b) 3UPU spherical wrist", "texts": [ " Unfortunately, to the author knowledge, the path-planning strategies available so far in the literature are mainly local, which means that if they fail to find a singularity-free path it is not always sure that a singularity free path does not exist at all. Obviously, the problem of singularity-free path-planning is strictly connected to the problem of identifying and characterizing the different disjoint regions into which the workspace is split by the singularity locus. These regions were named aspects and rigorously defined in [11]. The aim of this paper is to propose a global numerical method to avoid singularities of 3UPS and 3UPU parallel spherical wrists (Fig. 1). This method is based on Morse theory, and is able to identify the aspects of the workspace through the critical points of the determinant of the Jacobian matrix, which is a function of the orientations of the platform whose zero level-set defines the singularity locus. Once the different aspects are counted and identified, the proposed method is able to detect if any two orientations belong to the same aspect or not, and, in the first case, it is able to find a singularity-free path connecting them" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000841_icsmc.2004.1401038-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000841_icsmc.2004.1401038-Figure3-1.png", "caption": "Figure 3. The boundary produced by leg i", "texts": [ " q3 = W ) v v , =(q3 a B y Hence, w x ui can be written as : where : wxui = A i v , (10) Ai = (0 Reaio Rpaio) -sina - cosas inp C O S ~ C O S P -cosa s i n a s i n p - s inacosp o - s inacosp -s inadinp o - cosacosp - cosas inp 0 O I R =-= aR [ 0 aa -sin p and Ai is given by : Ai = ra x 1 1 B = [ s i n p 0 93 cos P -sinacospi -cosasinpsinp, -sinacospsinP, -cos a cos pi +sin a sin psin Pi -cos a cospsin Pi (11) 0 -sin p sin pi v = q3w3 + q3 (03 x w,) = Bv, The equation (7) can be written as : where : sin a cos p - g, sin a sin p cosacosp -9, s inacosp -q3cosas inp g3 cos a cos p To transform equation (1 1) into : Substituting the relationship from (12) into (IO), produces the following equation : Substituting (U), (13) into (X), (9), a equation is written as: qi = Jv (12) 1 ve = B- v w x u i = A,B% (13) where : J3 = w j The condition index K is defined for the manipulator as : oms = maxf i} omin = m i n k } Ai is the eigenvalue of J'J 5 Workspace analysis A schematic of the boundary produced by leg i is shown in Fig.3. Considering the constraint of the length of qi ( i = l , 2 ) , in the interval [qmin,gmar], each limb produces a minimum workspace boundary and a maximum workspace boundary. The manipulator workspace is found by intersecting the workspaces of the individual legs. In this paper, the workspace boundary is given by using the golden section search, as the steps by the following: Step 1: A reference frame 0 - r p z is used to instead of 0-XYZ . Search boundary points by finding z that can make g i solved by inverse kinematics equal to gmin and qmin with the fixed values of r and p " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000828_iros.2003.1248883-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000828_iros.2003.1248883-Figure3-1.png", "caption": "Fig. 3. Joint angles", "texts": [ " Each oscillator changes its speed depending on the touch sensor signal, and the effects reflected on the oscillator in each limb. As a result, the desired trajectory of each joint is adjusted so that global entrainment between dynamics of the robot and those of the environment is realized. In the following, the details of each controller are explained. I) Trajectory contmllec The trajectory controller calculates the desired trajectory of each joint depending on the phase given by the corresponding oscillator in the phase controller. Here, the trajectory of each joint is characterized by four parameters as shown in Fig. 3. For joints 3 , 4 and 5, which coincide with pitch axis, the desired trajectory is determined so that in the swing phase the foot trajectory draws a ellipse that has the radii, h in the vertical direction and /3 in the horizontal direction, respectively. For joints 2 and 4, which coincide with roll axis, the desired trajectory is determined so that the leg tilts from -W to W relative to the vertical axis. The amplitude of the oscillation, a, determines the desired trajectory of joint 1. The desired trajectories are summarized as following functions, e, = asin(@) (1) e, = Wsin(@) (2) ei = f i ($>h,P) ( i = 3,4,5) (3) e, = -wsin(@)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002843_acssc.2009.5470109-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002843_acssc.2009.5470109-Figure1-1.png", "caption": "Fig. 1. Exemplary comparison of conventional CP (left) and broadcast CP (right): with conventional CP, node #1 computes and transmits three different message pairs; with broadcast CP, node #1 computes and transmits only one message pair and nodes #2, #3, and #4 perform a postprocessing. (Nodes performing computations are encircled with thick solid lines.)", "texts": [ " (3b) This requires one broadcast transmission instead of |N (i)| transmissions from node i to all neighbor nodes j \u2208 N (i). We note that the current estimate in (2) can now be simply obtained as z\u0302 (n) i = K\u0303 (n) i /\u03bc\u0303 (n) i . To recover the original CP messages, all destination nodes need to perform the following post-processing steps: K (n) i\u2192j = K\u0303 (n) i \u2212 K (n\u22121) j\u2192i 1 + 1 \u03b2 ( K\u0303 (n) i \u2212 K (n\u22121) j\u2192i ) , (4a) \u03bc (n) i\u2192j = \u03bc\u0303 (n) i \u2212 \u03bc (n\u22121) j\u2192i K (n\u22121) j\u2192i K\u0303 (n) i \u2212 K (n\u22121) j\u2192i , (4b) A schematic illustration of the operational principle underlying conventional and broadcast CP is given in Fig. 1. We emphasize that broadcast CP yields exactly the same results as traditional CP and therefore all CP convergence studies apply to broadcast CP as well. However, if the messages are quantized or if there are transmission errors, the results do not coincide anymore. Due to lack of space, these issues are not addressed in this paper. It can be verified that conventional CP and our proposed broadcast CP scheme have the same overall computational complexity that scales linearly with the number of edges (links) in the network" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001740_icia.2005.1635112-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001740_icia.2005.1635112-Figure3-1.png", "caption": "Fig. 3 Moment of inertia about C.M.", "texts": [], "surrounding_texts": [ "When an object is rotated with angle \u03b8 by two robot fingers, the relation between the moment of inertia IG about C.M. and the moment NG around C.M. can be represented as IG\u03b8\u0308 = NG, (27) Here, let us denote the coordinates of fingertip 1 and fingertip 2 on a frame whose origin is set at C.M., by GR1, GR2, the distance from C.M to the Coulombfrictional resultant F c by Rcg, the distance from C.M to the viscosity-frictional resultant F \u03b3 by R\u03b3g. Thus, the moment around C.M. NG will be NG = F f1 \u00d7GR1 +F f2 \u00d7GR2 +F c \u00d7Rcg + \u03b3P\u0307 \u00d7R\u03b3g. (28) From (27) and (28), there is IG\u03b8\u0308 = F f1 \u00d7GR1 +F f2 \u00d7GR2 +F c \u00d7Rcg +\u03b3P\u0307 \u00d7R\u03b3g, (29) where, F f1, F f2 are obtained from fingertip force sensors, \u03b8\u0307, \u03b8\u0308 are from the variations of fingertip position, GR1, GR2 are from the obtained C.M. positions. The unknown parameters are IG, F c, Rcg and \u03b3R\u03b3g. By changing angular acceleration \u03b8\u0308, 2 equations based on (29) can be obtained as IG \u03b8\u0308(1) = F (1) f1 \u00d7 GR1 + F (1) f2 \u00d7 GR2 +F (1) c \u00d7 Rcg + P\u0307 (1) \u00d7 \u03b3R\u03b3g, (30) IG \u03b8\u0308(2) = F (2) f1 \u00d7 GR1 + F (2) f2 \u00d7 GR2 +F (2) c \u00d7 Rcg + P\u0307 (2) \u00d7 \u03b3R\u03b3g. (31) Since F (1) c and F (2) c are the same, we have: IG[\u03b8\u0308(1) \u2212 \u03b8\u0308(2)] = [F (1) f1 \u2212 F (2) f1 ]\u00d7 GR1 +[F (1) f2 \u2212 F (2) f2 ]\u00d7 GR2 + [P\u0307 (1) \u2212 P\u0307 (2) ]\u00d7 \u03b3R\u03b3g, (32) so that only IG, \u03b3R\u03b3g are unknown. For the 2 unknown parameters, we take 4 equations based on (29) respectively with different angular accelerations. From the differences between the equations, we have: \u23a1 \u23a3 [ \u03b8\u0308(1) \u2212 \u03b8\u0308(2) ] \u2212 [ [P\u0307 (1) \u2212 P\u0307 (2) ]\u00d7 ] [ \u03b8\u0308(3) \u2212 \u03b8\u0308(4) ] \u2212 [ [P\u0307 (3) \u2212 P\u0307 (4) ]\u00d7 ] \u23a4 \u23a6 [ IG \u03b3R\u03b3 ] = \u23a1 \u23a3 [ F (1) f1 \u2212 F (2) f1 ] \u00d7 GR1 + [ F (1) f2 \u2212 F (2) f2 ] \u00d7 GR2[ F (3) f1 \u2212 F (4) f1 ] \u00d7 GR1 + [ F (3) f2 \u2212 F (4) f2 ] \u00d7 GR2 \u23a4 \u23a6 . (33) Therefore, the moment of inertia around C.M. IG can be obtained by [ IG \u03b3R\u03b3 ] = \u23a1 \u23a3 [ \u03b8\u0308(1) \u2212 \u03b8\u0308(2) ] \u2212 [ [P\u0307 (1) \u2212 P\u0307 (2) ]\u00d7 ] [ \u03b8\u0308(3) \u2212 \u03b8\u0308(4) ] \u2212 [ [P\u0307 (3) \u2212 P\u0307 (4) ]\u00d7 ] \u23a4 \u23a6 \u22121 \u23a1 \u23a3 [ F (1) f1 \u2212 F (2) f1 ] \u00d7 GR1 + [ F (1) f2 \u2212 F (2) f2 ] \u00d7 GR2[ F (3) f1 \u2212 F (4) f1 ] \u00d7 GR1 + [ F (3) f2 \u2212 F (4) f2 ] \u00d7 GR2 \u23a4 \u23a6 . (34) By taking 2N (N \u2265 2) equations based on (29) with different angular accelerations, from the differences between the equations, IG, \u03b3R\u03b3 can be estimated with well accuracy as [ IG \u03b3R\u03b3 ] = D+ \u03b8P [DF1 \u00d7 GR1 + DF2 \u00d7 GR2], (35) D\u03b8P = [ [\u03b8\u0308(1) \u2212 \u03b8\u0308(2)] \u00b7 \u00b7 \u00b7 \u2212[[P\u0307 (1) \u2212 P\u0307 (2) ]\u00d7]T \u00b7 \u00b7 \u00b7 [\u03b8\u0308(2N\u22121) \u2212 \u03b8\u0308(2N)] \u2212[[P\u0307 (2N\u22121) \u2212 P\u0307 (2N) ]\u00d7]T ]T , (36) DF1 = [ [[F (1) f1 \u2212 F (2) f1 ]\u00d7]T \u00b7 \u00b7 \u00b7 [[F (2N\u22121) f1 \u2212 F (2N) f1 ]\u00d7]T ]T , (37) DF2 = [ [[F (1) f2 \u2212 F (2) f2 ]\u00d7]T \u00b7 \u00b7 \u00b7 [[F (2N\u22121) f2 \u2212 F (2N) f2 ]\u00d7]T ]T . (38) VI. Verification by Experiments We have done some experiments for verifying the validity of the proposed method. In the experiments, the object is a wooden box with sizes 200(mm)\u00d7150(mm) \u00d7100(mm) in which 1(kg) and 0.5(kg) weights have been put. Pushing operations are performed by a 6 joints manipulator (Mitsubishi, MELFA RV-1A, Position repeatability: +/-0.02(mm)), on the effector of which two fingers with 6-dimensional fingertip-force sensors (BL Autotech, NANO 5/4 Sensor, Resolution: 3.3(gf)) are equipped as shown by Fig.4. There are unknown frictions among the fingertips, the object and its environment. The environment for the pushing operation is a level acrylic plane, and the frictions on the contact plane between environment and object are uniformity. In the experiments, the robot arm will put its two fingertips on an side of object and push the object within 5(mm). If the two fingers have sensed contact forces respectively, a C.M. line will be between the two fingertips so that the object can be pushed straightly or with a certain rotation. Otherwise, the fingers will move 5(mm) right or left along the side of object and push again till a C.M. line between the fingers is found. Then the fingertips will still adjust thier positions to make the difference between two fingertip forces as small as possible. After deciding the fingertip\u2019s position on the object, the finger will slowly push the object slowly straightly or rotatively with different accelerations or angular accelerations, to obtain the information on the forces, positions, velocities and accelerations of fingertips. Over 8 different accelerations are taken out in each experiment. TABLE I" ] }, { "image_filename": "designv11_61_0000267_papcon.2002.1015139-Figure21-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000267_papcon.2002.1015139-Figure21-1.png", "caption": "Figure 21 : Placement of Thermocouple", "texts": [ " temperature under load using the following equation: Now there is enough information to calculate the That ~ AMBhot = Thot r ise Required data: Ambient cold and hot Resistance cold and hot Total temperature of winding when R, was measured, in \"C (hot). Resistance measured during test, in P. Reference value of resistance previously measured at known temperature tb in Q. K = 234.5 for copper If it is not practical to expose the motor connections to access the leads to apply the resistance bridge, an alternate approach would be to attach a thermocouple to the stator lamination at the center to the stator core as shown in Figure 21. For totally-enclosed motors, the lamination temperature at the center line (A, Figure 21) is usually 5 to 10 percent less than the average winding temperature. If the depth of the frame is known, a hole can be drilled down to the lamination (8. Figure 21). This method would be more than adequate for field measurements. Temperature of winding when reference value of resistance R, was measured, in \"C (cold). 234.5 for 100% International Annealed Copper Standard (IACS) conductivity copper. NOTE: For other winding materials, a suitable value of k (inferred temperature for zero resistance) must be used. - 127 The use of thermoaraDhv to obtain the lamination CONCLUSION temperature is another aiernatik In conclusion, a basic understanding of the impact of motor temperature on the stator, rotor and bearings can help achieve satisfactory motor life and pefiormance" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003705_s11249-013-0154-6-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003705_s11249-013-0154-6-Figure1-1.png", "caption": "Fig. 1 Configuration of a HRHB: a Rolling bearing supporting state at low speeds, b hydrodynamic bearing supporting state at high speeds", "texts": [ " When the load on the rolling bearing is heavier than the skidding critical load, the cage rotates at a theoretical speed of pure rolling, while it shows obvious skidding when the load on the rolling bearing is less than the critical load value [24\u201329]. Published data [22\u201329] show that a direct relationship exists between the load on the rolling bearing and the cage speed. Therefore, a method by which to identify operation modes for HRHBs based on monitoring the cage speed of rolling bearing was developed in this study. Experimental investigation was conducted to examine the variation in cage speed with the shaft speed, as well as the effects of external load and starting time on cage speed. As shown in Fig. 1, the HRHB consists of a hydrodynamic bearing and a rolling bearing with a clearance. The clearance of rolling bearing is smaller than that of the hydrodynamic bearing. There are two distinct bearing operation modes: (1) at low speeds, where because oil film pressure is not sufficiently high to raise the rotor, the shaft only rides on the rolling bearing without any rubbing with the hydrodynamic bearing (Fig. 1a); (2) at high speeds, where the hydrodynamic bearing raises the rotor so that it floats in the clearance of the rolling bearing, with the result that the rolling bearing does not interfere with the rotor (Fig. 1b). To ensure that the rolling bearing does not rotate with the rotor at high speeds, several magnets are set on the side surfaces of the outer ring to attract the rolling elements to the raceway of the outer ring; thus, the clearance of the rolling bearing can be kept between the raceway of the inner ring and the rolling elements. The load on the rolling bearing changes according to the operation modes [21]. In the rolling bearing supporting state, the load on the rolling bearing gradually decreases from full load to zero as the oil film force of the hydrodynamic bearing increases from zero to full load" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003159_2008-32-0061-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003159_2008-32-0061-Figure2-1.png", "caption": "Figure 2: Multibody model of MP3.", "texts": [ " In addition to these measured quantities, the inertial platform also outputs, as computed signals, the roll and the pitch angles. GPS: it gives the position of the installation point in a ground frame and evaluates the point velocity, the distance travelled and the trajectory. Steering sensors: these sensors give the steering angle and the steering torque applied by the driver. VEHICLE MODEL \u2013 A mutibody model of the vehicle was implemented in the MSC-Adams environment. The model consists of several rigid bodies (see Figure 2). The inertial properties for the main bodies were obtained on the basis of the 3D CAD drawings and technical data. The front four linkage architecture is modelled with two side arms and two horizontal arms connected by four revolute joints. The mechanism is also linked to the mainframe by other two revolute joints on the upper and lower arm. The two front suspensions can rotate around the relative side arm when a steering movement is imposed to the steering column. In the rear part of the vehicle model there is the rear suspension assembly, which includes the engine and the CVT transmission" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001418_1.2227009-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001418_1.2227009-Figure5-1.png", "caption": "FIG. 5. Impact test sample geometry, X=gap width, Y =weld depth.", "texts": [ " Laser conditions Laser RS 6000, Nd-YAG Laser power measured at the material surface : 2800 W Lens focal length 275 mm Focal point position On the material surface MIG conditions Equipment ESAB-ARISTO LUD 450 W/MEK 44C Pulse current 32 A Pulse frequency 130 Hz Pulse length 2 ms Background current 60 A Shield gas 65% He, 30% Ar, 5% CO2, 24 I /min Electrode wire OK 12,51, 1 mm diameter Material RAEX 420 MC plate , by Rautaruukki 10 mm thick Variables Welding speed 0.5\u20132.0 m/min Wire feed rate 4.5\u201320 m/min MIG voltage 35\u201350 V This article is copyrighted as indicated in the article. Reuse of AIP content is subject to the terms at: http://scitation.aip.org/termsconditions. Downloaded to IP: 130.209.6.50 On: Sat, 20 Dec 2014 00:10:57 The geometry of the laser and MIG arrangement is given in Fig. 4. B. Impact testing The welded samples were sectioned and machined to produce the type of samples shown in Fig. 5. The samples were then broken in an A.B. ALPHA Sundbyberg, Sweden Impact testing machine type 1H539 with the pendulum striking the top surface of the weld . Two sample groups were tested; one at room temperature and another after immersion in liquid nitrogen. Before and after impact testing the samples were examined by optical and scanning electron micrograph microscopy. Figure 6 shows macrographs of a typical weld cross section in its welded, machined, and post impact testing states. The results of the impact tests are given in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003361_978-90-481-9689-0_37-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003361_978-90-481-9689-0_37-Figure3-1.png", "caption": "Fig. 3 Analysis model of walking using crutches (double support phase).", "texts": [ " Sensors (n) such as accelerometers are attached to the user and a brake is attached to the spherical joint between the soleplate and the links for safety from external disturbances. 321 T. Iwaya et al. In order to clarify the motion which should be generated by WAMC, we measured swing-through crutch locomotion by a healthy person to ascend and descend steps. We modeled a human using crutches as a planar serial mechanism with six links and six revolute joints during the single support phase by the legs, and as a planar closed-loop mechanism with seven links and seven revolute joints during the double support phase (Figure 3a). In the figure, geometrical parameters of the step and positions of the foot and the crutch tips are defined. We put markers on the body of the subject and the crutches. Their positions were measured by a motion capture system. In order to investigate the planar motion in the sagittal plane, we transformed 3D measurement data to 2D data. A planar model with three moving links, three revolute joints and one prismatic joint shown in Figure 3b is used in dynamic simulation, in which the effect of user\u2019s body is included. In the figure, R corresponds to the center of the spherical joint (j) in Figure 2. We conducted experiments under conditions listed in Table 1. From the experimental data, we found that the change of velocity between a virtual point corresponding to R and a point S (shoulder), which correspond to the displacement and velocity of the linear actuator of WAMC, can be modeled by the parameterized 322 Development of a Walking Assist Machine Using Crutches curve shown in Figure 4", " As the result, we found that walkings for ascending and descending steps as well as on flat floor are made possible when appropriate values of parameters of this velocity profile are used. Though the parameters to be determined according to the walking conditions are Vex, Vin, \u03b8ex and Vout, we considered Vex, Vin and \u03b8ex as variables and Vin as a constant (Vin = 0.3 m/s) in the following. Dynamic simulations were carried out to determine values of parameters of the reference trajectory. Table 2 shows the values of the dynamic parameters of the model in Figure 3b used in simulation. First, we investigated the area of set of parameters, by which stable walking for ascending a step can be achieved. Figure 5a shows the minimum value of Vex vs. \u03b8ex for three cases of Vin for ascending a step of 323 T. Iwaya et al. hS = 80 mm. As shown in this figure, it is known that the minimum value of Vex is dependent on \u03b8ex, and almost independent of Vin. Since Vex affects the body of the user during the DS phase, smaller Vex is expected preferable to improve user\u2019s comfort" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002685_rnc.1378-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002685_rnc.1378-Figure1-1.png", "caption": "Figure 1. Inverted pendulum on a cart constrained on a curve.", "texts": [ " The following notations will be used: \u2200a,b\u2208R, a\u2227b=min{a, b}, a\u2228b=max{a, b} and [a,b]= {x \u2208R|a x b}, ]a,b[={x \u2208R|a0, there exists a positive-definite symmetric matrix X and Li =Ki X, which satisfy the following LMIs: Ai X+XAT i +B2Li +LT i BT 2 <( 2 )X (13) rX XAT i +LT i BT 2 AiX+B2Li rX <0 (14) I11i I21i I12i I22i <0 (15) Ai X+XAT i +B2Li +LT i B T 2 BT 1 C1X+D12Li B1 D11 XCT 1 +LT i DT 12 DT 11 <0 (16) where I11i =sin \u00b7 Ai X+X(sin \u00b7 Ai ) T +(sin \u00b7 B2)Li +Li T(sin \u00b7 B2) T I21i =cos \u00b7 Ai X cos \u00b7 XAi T +cos \u00b7 B2Li cos \u00b7 LT i B2 T I12i = IT 21i I22i = I11i i=1, . . . , N Suppose (X*, L*) is one feasible solution of the above LMIs. Then the matrix X* and the feedback gain K*i =L*i (X*) 1 are the positive definite matrix Xcl >0 and the vertices\u2019 controllers, respectively, which satisfy Theorem 2. For the pole-placement region S( , r, ) as shown in Figure 2, according to the relationship between MD(A, X) and fD(z), the following LMIs, which satisfy the pole-placement requirements, can be obtained from Definition 2 and Theorem 1: there exists XD >0 such that Acli XD +XD AT cli +2 XD <0 (17) rXD XD AT cli Acli XD rXD <0 (18) sin (Acli XD +XD AT cli ) cos (XD AT cli Acli XD) cos (Acli XD XD AT cli ) sin (Acli XD +XD AT cli ) <0 (19) Also, from Definition 3 and Theorem 1, the LMI, which guarantees Twzi (s) to possess a quadratic H performance Twzi < , is: there exists X >0 such that Acli X +X AT cli BT cli Ccli X Bcli I Dcli X CT cli DT cli I <0 (20) From (11), Acli =Ai +B2Ki, Bcli =B1, Ccli =C1 +D12Ki and Dcli =0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003411_s11771-012-0984-7-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003411_s11771-012-0984-7-Figure6-1.png", "caption": "Fig. 6 Photograph of spindle with vibration measurement equipment", "texts": [ "1\u2212200 mm/s, a vibration range of 0.05 mm/s, and an uncertainty of measurement of 0.03 mm/s with its equipment resolution. Also, the bandwidth of this device was 10 Hz\u221210 kHz and the measurement was performed by the unit of mm/s. In Korean Industrial Standards, the most reasonable method for evaluating the broadband vibration of a rotation machine is to consider the RMS (root mean square) value of its vibration speed [20]. This is due to the fact that the vibration speed is related to the vibration energy. Figure 6 shows the configuration of a vibrometer. As shown in Fig. 7, two vibration measurement points were determined to a vertical direction. Also, the spindle speed increased up to 1 000\u22125 000 r/min with an interval of 1 000 r/min. In the experiment environment, temperature and humidity were maintained as 21 \u00b0C and 45%, respectively. Figure 8 presents the values of the measurement at Point 1 and Point 2 according to spindle speeds. As the automatic variable preload device was not installed at a maximum speed of 5 000 r/min, vibrations were recorded by 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003947_icca.2011.6137946-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003947_icca.2011.6137946-Figure4-1.png", "caption": "Fig. 4", "texts": [ " (2) [Theorem 1] Each independent basic subsystem tend to their own consensus state, i.e there exists \u03b7s \u2208 Rm, lim t\u2192\u221e xi = \u03b7s ai \u2208 As , where As is an independent basic set of A on Ga j. (3) [Theorem 2] \u039b = {\u03b71, \u00b7 \u00b7 \u00b7 ,\u03b7q} is the set of the consensus states of all independent basic subsystems. vp(\u221e)(p= 1, ...,s) are vertexes of \u039e(\u221e), V (\u221e) = {vp(\u221e)|i = 1, ...,s}. V (\u221e)\u2282 \u039b (4) [Corollary 2] All states of the cooperative system tend to consensus if and only if there is only one independent basic subsystem in the system The left part of Figure 4 shows the edges which may appear in graph during the time interval [0, tr). But some of them will disappear over the time interval [tr,\u221e) for some tr 1. These edges may have no effect upon the collective behavior of system. The right part of Figure 4 only includes the edges which will persistently appear over the time interval [0,\u221e) although it may be intermittent. These edges are adopted by the adjoint graph Ga j of G(t). The collective behavior of the system is essentially determined by the topology of adjoint graph Ga j. The adjoint graph Ga j of system is given by Figure 5. Two rectangles represent two independent subsystems corresponding to the two independent subsets of Ga j. Each independent subsystem has only one independent basic subsystem" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003015_cae.20257-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003015_cae.20257-Figure2-1.png", "caption": "Figure 2 CAD model for assemble parts of automotive electric. [Color figure can be viewed in the online issue, which is available at wileyonlinelibrary.com.]", "texts": [ " Based on CyberCAD, VDS integrates the structure of a collaborative design system with Internet technologies by extending the client/server structure. For real-time co-design based on traditional commercial CAD software among many clients, a Java-coded software interface is developed to extend the single-location CAD software to a multi-location application through the Internet [25]. First of all, CAD models of each component and measuring tools (e.g., venire, CMM and so on) are built by CAD/CAM software in this study. Figure 2 illustrates the CAD models of the assemble parts with Figure 2a as the headlight, Figure 2b the distributor and Figure 2c the generator. Meanwhile, because the data (includes geometric features, dimensions, etc.) of component are key measuring or wiring elements, smooth information flow and efficient information management are essential. Combine the object-oriented approach with hierarchical structure of adaptive layer integration for realizing design data is proposed in this research. Object-orientation provides a systematic approach to integration in design data. A more detailed description of object-orientation of large information infrastructures has been provided in the forthcoming book on digital libraries (see Ref" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003117_ichr.2009.5379588-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003117_ichr.2009.5379588-Figure2-1.png", "caption": "Fig. 2. Model of the robot.", "texts": [ " \u2022 The robot should stably manipulate the door and pass through it without falling down. In this paper, we developed a trajectory planner that takes into account the above mentioned matters. The coordinate system in which the robot operates is set as in Fig.l, where the origin 0 = (xo,Yo) is the point in the middle of the walls. We denote with S = (xs,Ys) the starting point and with G = (xg,Yg) the position of the goal. To reduce the amount of computations, we model the robot as a parallelepiped whose base dimensions, ar, b- , comply with the physical sizes of the humanoid, see Fig.2. Its center R = (xr,Yr) coincides with the Center of Mass (CoM) of the robot and its orientation is expressed by er . M; and I; design the mass and the momentum of inertia, respectively, calculated with regard to the parallelepiped. To analyze the physical interaction with the door, we have to consider the space that the arms of the humanoid can span. The manipulable area is represented by a rectangle, that includes the above mentioned parallelepiped, and by two half circles, of radius T r , centered in the joints of the arms", "0) and S = (3.0, -2.0), to show how the robot interacts with the door, approaching it from different larger sampling time. For simplicity, we assume it is constant, and we set its value equal to O.8sec. Then, as references for position and posture of the robot at step i, we give the values at step i+ I to the controller. The walking motion at each step is defined by the pattern generator [7]. While opening the door, the arms are controlled to keep the door away from the manipulable area shown in Fig.2. Otherwise, the posture of the arms is kept constant. Because of the reaction force coming from the door, we have to use the stabilizer to keep the stability of the robot. The simulation result of the generated motion is in Fig.15. As shown in Fig.15(e), while the robot interacts with the door, the latter does not enter the manipulable area. The robot opens the door stably and passes through it without falling down. In Fig.14, the trajectory followed by HRP-2 is printed on the one generated by the path planner described in Section III" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000961_0065955x.1977.11982423-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000961_0065955x.1977.11982423-Figure2-1.png", "caption": "Fig 2. - Prismatic effect of spherical lens.", "texts": [ " The prismatic effect of a spherical lens is dependent on the strength of the lens and distance from the optic center of light passing through the lens. STRENGTH OF THE LENS.- The stronger the lens, the greater the prismatic effect. A 1-0 spherical lens decentered 1 cm will produce a deflection of a beam of rays parallel to its principal axis of 1 cm at 1 m. This is equal to the deflecting power of a 1-8 prism. A 2-0 spherical lens decentered 1 cm will cause 2 cm of deflection at 1 m or 28. Multiply the dioptric power of the lens by the decentration in centimeters, and the result is the pris matic effect. DISTANCE FROM THE OPTIC CENTER. Figure 2 shows that a spherical lens has a stronger prismatic effect as one approaches the periphery. At c and d the tangents will meet at 1, forming a prism with little deviating power. At e and f the tangents are more inclined towards one another and so form a prism with greater deviating power. However, one should realize that when prisms are divided between the two eyes, each lens has to be decen tered the same amount or the pris- PRISMS IN PARALYTIC SQUINT 55 matic effect also is divided in half. For example, a 10-D sphere decentered 2 cm gives a prismatic effect of 20a" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003788_1.3276412-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003788_1.3276412-Figure1-1.png", "caption": "Fig. 1. The setup of the magnetically controlled pendulum used to produce self-excited oscillations with a reed switch.", "texts": [ " Except for the data-acquisition system, the apparatus is simple and inexpensive. \u00a9 2010 American Association of Physics Teachers. DOI: 10.1119/1.3276412 I. INTRODUCTION Recently, a magnetically controlled pendulum was used in some experiments on nonlinear dynamics.1 The same pendulum can be used in many other experiments, some of which are described in this paper. The magnetically controlled pendulum consists of a thin 12 cm long aluminum rod with two permanent magnets attached to its ends, as shown in Fig. 1. The pivot of the rod, located 4 cm below the top of the rod, is attached to the shaft of a PASCO Rotary Motion Sensor CI-6538 . The centers of the upper and lower magnets are positioned 2.5 cm above and 7.5 cm below the pivot, respectively. The magnets are strong ceramic magnets 1.2 1.2 1.2 cm3 in size. The magnetic dipoles can be aligned parallel or perpendicular to the axis of the pendulum. An aluminum disk, 1.2 cm thick and 4 cm in diameter, attached to the other end of the shaft of the Rotary Motion Sensor can be subjected to the magnetic field of another permanent magnet", " Redistribution subject to AAPT lice The second method used to compensate for the energy losses uses a reed switch to provide suitable current pulses to the driving coil, as shown in Fig. 2 b . The switch consists of a pair of contacts on ferrous metal reeds in a hermetically sealed glass envelope. The state of the switch depends on the external magnetic field. We use a dry reed switch AssemTech GC 3817 that is open when the magnitude of the external magnetic field is small. When this method is used, the axis of the lower magnetic dipole is aligned vertically. The switch is located 2 cm away from the equilibrium position of the pendulum, as shown in Fig. 1, and is triggered by the lower magnet of the pendulum. The switch is connected in series with a dc power supply and the driving coil. Every time the pendulum reaches its lowest point, the switch is closed and an electric current of about 0.3 A passes through the driving coil. The switch breaks the current when the pendulum moves away from its lowest point. A 5 F capacitor, shunting the driving coil, suppresses emf pulses in the driving coil when the switch opens. To start the oscillations, the dc supply is switched on when the pendulum is in its lowest position" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000961_0065955x.1977.11982423-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000961_0065955x.1977.11982423-Figure7-1.png", "caption": "Fig 7 - Strength of horizontal prisms in 30 patients", "texts": [ " It is possible that during the exam ination the patient's deviation be comes more and more dissociated, especially with horizontal diplopia. In that case we have two results, a mini mal and a maximal separation. When the deviation is vertical, there will be hardly any difference and we prescribe the full correction. When the deviation is horizontal, we start with the minimal strength which corrects the existing deviation in the primary position. If the patient's complaints persist we may strengthen the prisms. When the deviation is variable, we also prescribe the weakest correction. Figure 7 shows the different strengths of horizontal prisms we prescribed for 30 patients. The peak lies around 5!J.. Figure 8 shows our use of vertical prisms for 50 patients. The peak here lies between 2!J. and 4!J.. The aim of the prisms is to eliminate any abnormal head posture and to create orthophoria in the more fre quently needed fields of gaze, that is in primary and downward gaze. In a vertical deviation, we should disregard diplopia in the upward gaze, that being of little importance. One must warn the patient that the ocular move ments under prism correction should beof small amplitude" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000443_s00170-005-0009-x-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000443_s00170-005-0009-x-Figure1-1.png", "caption": "Fig. 1 A spatial parallel mechanism with 4-PUU", "texts": [ " Moreover, Newton\u2212Raphson will offer a faster convergence speed than Newton-GMRES, which is described in theorem 1 and 2. Second, when the differential coefficient matrix G\u2032(X) is nearly singular at the nearby of the real solutions, the Newton\u2212Raphson method will be converged to the real solution at a very low speed, while Newton-GMRES algorithm is quite a good redeeming method to crack this problem because it can avoid calculating [11] the differential coefficient matrix. 3.1 The forward and inverse displacement of a kind of 4-PUU parallel manipulator A spatial parallel manipulator, shown in Fig. 1, is made up of 4-PUU (1 prismatic joint and 2 universal joints) kinematic chains. The absolute coordinate system oxyz are created as Fig. 1 shows, where z-axis is perpendicular to the guide plane P1P2P3P4, the origin is on the midline of the two guides, x is superposed with the midline of the two guides and y-axis is perpendicular to the two guides. According to the method presented above, the local coordinate system oc xc yc zc is shown in Fig. 2, where zc-axis is perpendicular to the plane of the manipulator M1M2M3M4, the origin is superposed with the geometric center of M1M2M3M4, xc and yc axes are parallel to the two orthogonal sides of the manipulator. Firstly, we will analyze the DoF of the manipulator as Fig. 1 shows. According to [9], we can find the dimension of the constraints spaces that all of the reciprocal screws, shown in Fig. 3, can be spanned is: d \u00bc dim span $rB1P1M1 $rB2P2M2 $rB3P3M3 $rB4P4M4 8>< >: 9>= >; \u00bc 2: (11) Therefore, F \u00bc 6 d \u00bc 6 2 \u00bc 4: (12) So, the manipulator shown in Fig. 1 has four DoFs, including three orthogonal translational movements and one rotational moment around zc-axis. Now we can select (xc yc zc) and the rotational angle around zc-axis, \u03b2, as the stance parameters. If we presume the length ofM1M2 to be 2a and the length ofM1M4 to be 2b, the local coordinates of the four vertexes of the manipulator can be obtained: rLM1 \u00bc b a 0\u00bd T rLM2 \u00bc b a 0\u00bd T rLM3 \u00bc b a 0\u00bd T rLM4 \u00bc b a 0\u00bd T : 8>>< >>: (13) The transform matrix from the local coordinate system to the absolute one is: A \u00bc cos sin 0 sin cos 0 0 0 1 2 4 3 5: (14) With Eq", " 5, there are: J4 4 \u00bc @G xc; yc; zc; \u00f0 \u00de @ xc; yc; zc; \u00f0 \u00de \u00bc 2 xc \u00fe b cos a sin x1 yc a cos b sin \u00fe c zc a yc \u00fe c\u00f0 \u00de \u00fe b x1 xc\u00f0 \u00de\u00bd sin \u00fe a x1 xc\u00f0 \u00de b yc \u00fe c\u00f0 \u00de\u00bd cos xc \u00fe b cos \u00fe a sin x2 yc \u00fe a cos b sin c zc a c yc\u00f0 \u00de \u00fe b x2 xc\u00f0 \u00de\u00bd sin \u00fe a xc x2\u00f0 \u00de b yc c\u00f0 \u00de\u00bd cos xc b cos \u00fe a sin x3 yc \u00fe a cos \u00fe b sin c zc a c yc\u00f0 \u00de \u00fe b xc x3\u00f0 \u00de\u00bd sin \u00fe a xc x3\u00f0 \u00de \u00fe b yc c\u00f0 \u00de\u00bd cos xc b cos a sin x4 yc a cos \u00fe b sin \u00fe c zc a yc \u00fe c\u00f0 \u00de \u00fe b xc x4\u00f0 \u00de\u00bd sin \u00fe a x4 xc\u00f0 \u00de \u00fe b yc \u00fe c\u00f0 \u00de\u00bd cos 2 6664 3 7775 (19) Let |J4\u00d74| = 0, we can obtain: xc \u00fe b cos a sin x1 yc a cos b sin \u00fe c zc a yc \u00fe c\u00f0 \u00de \u00fe b x1 xc\u00f0 \u00de\u00bd sin \u00fe a x1 xc\u00f0 \u00de b yc \u00fe c\u00f0 \u00de\u00bd cos xc \u00fe b cos \u00fe a sin x2 yc \u00fe a cos b sin c zc a c yc\u00f0 \u00de \u00fe b x2 xc\u00f0 \u00de\u00bd sin \u00fe a xc x2\u00f0 \u00de b yc c\u00f0 \u00de\u00bd cos xc b cos \u00fe a sin x3 yc \u00fe a cos \u00fe b sin c zc a c yc\u00f0 \u00de \u00fe b xc x3\u00f0 \u00de\u00bd sin \u00fe a xc x3\u00f0 \u00de \u00fe b yc c\u00f0 \u00de\u00bd cos xc b cos a sin x4 yc a cos \u00fe b sin \u00fe c zc a yc \u00fe c\u00f0 \u00de \u00fe b xc x4\u00f0 \u00de\u00bd sin \u00fe a x4 xc\u00f0 \u00de \u00fe b yc \u00fe c\u00f0 \u00de\u00bd cos \u00bc 0 As a result, the singularity criteria of the manipulator can be simplified as: zc x1 x2 \u00fe x3 x4 0 b x1 \u00fe x2 \u00fe x3 x4\u00f0 \u00de\u00bd sin \u00fe a x1 x2 \u00fe x3 \u00fe x4\u00f0 \u00de\u00bd cos 2a sin x3 \u00fe x4 2a cos 2c 2ayc \u00fe b x4 x3\u00f0 \u00de\u00bd sin \u00fe a 2xc x3 x4\u00f0 \u00de 2bc\u00bd cos 2b cos \u00fe x1 x4 2b sin b 2xc x1 x4\u00f0 \u00de\u00bd sin \u00fe a x4 x1\u00f0 \u00de \u00fe b 2yc \u00fe 2c\u00f0 \u00de\u00bd cos \u00bc 0 (20) Considering the real applications of the manipulator shown in Fig. 1, we can assume that zc\u22650. Therefore, if the manipulator is not working at the singular position and posture, Eq. 20 will not hold. So we can create an iterative function: Presume X \u00bc xc yc zc 2 664 3 775, the ith iterative value of X are denoted as Xi \u00bc xci yci zci i 2 664 3 775 ; we can get the Newton iterative function with Eq. 7: Xi\u00fe1 \u00bc Xi J 1 4 4G Xi\u00f0 \u00de: (21) In the following we will illustrate some numerical results to show the benefits and drawbacks of the Newton\u2212Raphson method and Newton-GMRES algorithm" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000224_tasc.2003.813067-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000224_tasc.2003.813067-Figure1-1.png", "caption": "Fig. 1. Transformer winding configuration.", "texts": [ " The transformer windings are made from multifilamentary, not twisted, BSCCO/Ag tape with an overall cross section of 4.1 mm 0.3 mm, A, supplied by American Superconductor. The primary winding is a helical coil with 96 turns in 4 layers, total tape length is 61.5 m, the secondary coil consists of 12 double-pancake coils, each containing 4 turns of two parallel tapes, the total tape length of the secondary coils is 73.65 m. Total tape length of the transformer windings is 135.1 m. The transformer configuration is shown in Fig. 1. The AC losses of the transformer at 50 Hz were measured using NORMA Power analyzer D 4000 at different temperatures obtained by pumping of N gas. The potential taps were attached to each winding including all pancake coils. For numerical calculation of loses in the transformer coils we used experimental data for a similar tape published in [4]. Using Fig. 5(1) and (2) from [4] we plotted transport current loss and magnetization losses at 26 A vs. the amplitude of the external magnetic field, (See Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002645_gt2009-60243-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002645_gt2009-60243-Figure7-1.png", "caption": "Figure 7: Vibration Instrument Locations, Units A and B", "texts": [ " 4 Copyright \u00a9 2009 by ASME rms of Use: http://www.asme.org/about-asme/terms-of-use Dow There are three close-clearance labyrinth seals in the engine. The shaft seal (see Figure 6, VIEW A) has labyrinth teeth on the shaft running against a gear/bearing bronze alloy. It separates the compressor inlet from the aft bearing compartment. The inner and outer compressor seals are located on the back face of the compressor, adjacent to the Hirth coupling. Both seals have teeth on the rotor that rub against abraidable coatings. As shown in Figure 7, vibration sensors are located at five planes along the engine assembly. Velocity pickups on the generator are shown in Figure 8 and are labeled 3HV and 3VV (aft end) and 4HV and 4VV (forward end). Figure 9 shows the turbine output shaft end displacement sensor (eddy current proximity probe) used to generate vibration plots. nloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 10/04/2017 Ter Details that bring clarity on the cause and effect of the shaft heating hypothesis can be found in Figure 10 through Figure 12" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002988_icsps.2009.53-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002988_icsps.2009.53-Figure8-1.png", "caption": "Figure 8. Fault in the inner race of a ball bearing with its inner race stationary.", "texts": [ "DIGITAL SIGNAL PROCESSING The FFT is a method widely used in the field of vibrations. It is an easy method to use and in many cases can give out the cause of the vibration. In the Fig. 7, frequencies of 22 Hz, 66 Hz and 110 Hz are excited. Frequency 22 Hz is the rotational one, 66 Hz is the frequency excited when there is a fault in the outer race and 110 Hz is the frequency excited when there is a fault in the inner race. The way the ball bearings were used in the experiment, with their inner race stationary, as shown in Fig. 8, faults in the inner race were more probable to exist, since the same area of the inner race was always in the load zone. Fig. 9 illustrates a ball bearing with faults in the inner race. The FFT and the envelope spectrum [8], illustrated, above and below respectively, in Fig. 10, are taken from the same signal, which was just minutes before the destruction of the bearing. Investigating only the FFT signal it\u2019s very difficult, if not impossible, to understand where is the fault. In the FFT we can see the rotational frequency, the outer race defect frequency and the inner race defect frequency, and therefore the estimation of the faulty is difficulty and risky" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000138_bc.2003.010-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000138_bc.2003.010-Figure7-1.png", "caption": "Fig. 7 Schematic Representation of Possible Electrostatic Interactions in BI-VI Deduced from the Present Study. Arrows without annotation indicate electrostatic interactions. Annotations: CS and pKa indicate the induced chemical shift changes and dissociation-induced influence toward the interacting partners, respectively.", "texts": [ " However, their primary application may not be in predicting pKa values but in aiding the interpretation of observed electrostatic effects. Furthermore, as described above, all the conformations in the ensemble did not satisfactorily match the pKexp values of some ionizable groups in the molecule. This suggests that each conformation in the ensemble should be analyzed individually for its validity in terms of the electrostatic properties. The putative electrostatic interactions in BI-VI are illustrated in Figure 7, which are in reasonable agreement both with the experimental and with theoretical experiments. Electrostatic Interactions between BI-VI and the At present, the 3D structure of bromelain has not been determined; however, the coordinates of the same cysteine proteinase papain are available from PDB. Papain is a close homolog of stem bromelain with approximately 40% amino acid sequence identity (Mitchel et al., 1970; Ritonja et al., 1989), and the catalytic mechanism in papain has been extensively studied (reviewed by Storer and M\u00e9nard, 1994)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003897_20120620-3-dk-2025.00176-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003897_20120620-3-dk-2025.00176-Figure1-1.png", "caption": "Fig. 1. Admissible region \u039b\u03b8 (Here, \u03b8i and \u03b8i respectively denote the maximum and minimum bounds of \u03b8i, and \u2206i denotes the maximum deviation of \u03b8i for one sampling step.)", "texts": [ " Similarly to the conventional problem setting for Linear Time-Invariant (LTI) observers [Zhou et al., 1996], the control input u is also supposed to be provided. Hereafter, the step index k is omitted if it is obvious. Similarly, \u03b8(k + 1) is denoted by \u03b8+ for simplicity of notation. The vector \u03b8 is supposed to lie in a convex set \u2126\u03b8 which is given a priori. Furthermore, the parameter deviation for one sampling step, i.e. \u03b8i(k + 1) \u2212 \u03b8i(k), is supposed to be bounded. Then, (\u03b8, \u03b8+) is therefore supposed to lie in a convex set and its admissible region is denoted by \u039b\u03b8. (See Fig. 1.) For LPV plant Gp(\u03b8), we consider a GS state observer using the available scheduling parameters. The scheduling parameters are supposed to be provided with some uncertainties. That is, the i-th scheduling parameter is provided as \u03b8i(k) + \u03b4i(k) with its uncertainty \u03b4i(k). The bounds of \u03b4i are supposed to be known a priori; that is, the following holds: \u03b4i \u2208 [ \u03b4i, \u03b4i ] , (2) where \u03b4i, \u03b4i are a priori given scalars. Then, the vector \u03b4 = [\u03b41 \u00b7 \u00b7 \u00b7 \u03b4k] T lies in a convex set \u2126\u03b4 which is known a priori", " The parameter dependency is summarized in Table 1, and the optimized \u03b3 is shown in Tables 2 and 3. As \u03b4 is allowed to move arbitrarily fast, the performance does not depend on \u03b6. Next, we design discrete-time GS state observers using Theorem 6 with parameter-dependent matrices set in Table 4. The LPV system is discretized using Euler approximation for x\u0307 with sampling period Ts = 0.01 [s]. As the bound of \u03b8\u0307 is set as |\u03b8\u0307| \u2264 \u03b6 and Ts is set as 0.01, the deviation of \u03b8 for one sampling, i.e. \u2206i in Fig. 1, is set as 0.01\u03b6. The results are shown in Tables 5 \u223c 7. The effectiveness of using PDLFs is confirmed from these tables as long as the bound of parameter deviation is not so large. To confirm the importance of considering disturbance effect, a numerical simulation using discretized state-space matrices with \u03be and \u03b6 being respectively set as 0.1 and 1 was conducted. For comparison, the following two observers were designed using Theorem 6: (i) robust GS state observer in which B1(\u03b8) = 0 is supposed, and (ii) robust GS state observer without any assumption for the state-space matrices" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003756_icelmach.2010.5607764-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003756_icelmach.2010.5607764-Figure4-1.png", "caption": "Fig. 4. Machine example 2102212 >\u2212\u2212\u2212< with two elementary machines (phases are highlighted in color)", "texts": [ " In common rotor concepts it is related to the number of magnets pmN or pole pairs p resp., as follows: .2/pmNpt ==\u03bd (16) This equation is not valid for fragmented magnet poles or in case of a consequent-pole-type rotor [6]. Here, pmt Np==\u03bd is essential. In order to achieve a compact analytical formalism of combining different winding systems among each other, a classification according to [10] is presented here. To understand the basic features of the system, a concrete motor is exemplarily chosen as shown in Fig. 4. The usual classification found in the literature refers to the number of magnets pmN against the number of slots sN only (e.g. 2/3, 4/3 or 8/9). However, for more complex pole / winding arrangements, this characterization is not sufficient anymore and the pair of numbers has to be replaced by a quadruple >\u2212\u2212\u2212< pmNrrcN 21 for a complete definition of one elementary machine, which is the smallest functional unit. Thus, the complete machine may comprise EN elementary machines. With respect to Fig. 4 these numbers are given to: cN : Number of coils, 1r : Number of coils per zone, 2r : Number of coil groups (each with 3 zones). The number of coils is 213 rrNc = . Furthermore the machines may show single- or double-layer windings \u2212 here mentioned as one or two coil-sides per slot, noted with subscript indices 1 and 2 (see Fig. 4). So the winding definition is fixed due to repetition factors 1r , 2r plus the type of winding (one or two coil-sides per slot). As will be shown later, also winding factors and ordinal numbers due to the winding scheme could be represented completely by this quadruple. Table I shows a selection of feasible elementary machines with winding factors greater than 0.8. Obviously, at least two machines could not be specified by the usual spm NN / -classification (e.g. 2102212 >\u2212\u2212\u2212< and 2101412 >\u2212\u2212\u2212< )" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001555_iecon.2005.1569184-Figure14-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001555_iecon.2005.1569184-Figure14-1.png", "caption": "Fig. 14. New Coordinate System of ax,ay .", "texts": [ "4 0.45 Times[s] Fig. 12. Xzmp. V. Lagrange\u2019s Equation of System on slope It is on the upslope, that the wheelchair is easy-tooverturn. So we will deduce the Lagrange\u2019s Equation of System on slope, and try to find a universal criteria, which can be used both on horizontal plane and slope. The total kinetic energy T is T = 1 2 Mx2 + 1 2 mv2 G + 1 2 JM \u03b8\u03072 + 1 2 Jm\u03d5\u03072. (17) Now it will be proven that no matter how the coordination system of COM is defined the result will not be affected. As shown in Fig. 14, here we define the coordination system of COM as vertical and horizontal, then the velocity of COM can be calculated as below. { vmx = d dt (x cos \u03b6 + l sin(\u03d5\u2212 \u03b6)) vmy = d dt (x sin \u03b6 + l cos(\u03d5\u2212 \u03b6)) (18) Or { vmx = x\u0307 cos \u03b6 + l cos(\u03d5\u2212 \u03b6)\u03d5\u0307 vmy = x\u0307 sin \u03b6 \u2212 l sin(\u03d5\u2212 \u03b6)\u03d5\u0307 (19) So velocity of COM vG is v2 G = v2 mx + v2 my = x\u03072 + l2\u03d5\u03072 + 2x\u0307l\u03d5\u0307(cos \u03b6 cos(\u03d5\u2212 \u03b6) \u2212 sin \u03b6 sin(\u03d5\u2212 \u03b6)) = x\u03072 + l2\u03d5\u03072 + 2x\u0307l\u03d5\u0307 cos\u03d5 (20) Consequently, no matter how to define the coordination system of COM, the total kinetic energy T does not change" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001017_papcon.2004.1338357-Figure9-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001017_papcon.2004.1338357-Figure9-1.png", "caption": "Fig. 9. Three-mass system with two spring connections", "texts": [ " If speed BW (Crossover frequency) is indiscriminately set to higher values, system response will become less stable, and sustained damaging resonance producing torques will likely occur. Responses shown in Figs. 7 and 8 (without overshoot) can be considered well-behaved. Time domain responses with overshoot, such as those shown in Figs. 5 and 6, are not appropriate in many instances, as will be shown in section IV. 111. EVALUATION USING A THREE-MASS MODEL The effect of the HS (High speed) shaft stiffness, and the reducer inertia can be explored using a 3-mass model (see Fig. 9). Now J1, J2, and J3 are the motor, reducer and roll inertias, and K1 and K2 are the torsional stiffnesses of the HS shaft and jackshaft, respectively. A. Anti-resonance and Resonance Frequencies For the 3-mass model two pairs of anti-resonance / resonance frequencies show up. Table V shows the corresponding values for the Size Press section. The low pair of anti-resonance and resonance frequencies is almost the same as found using the 2-mass model. The highfrequency anti-resonance /resonance pair that show up using the 3- mass model is related to the torsional oscillations between the motor inertia (51) and the reducer inertia (J2) due to the HS shaft elastic constant" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002175_robot.2007.363146-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002175_robot.2007.363146-Figure7-1.png", "caption": "Fig. 7. Assembly configuration to avoid the third case of actuation singularity", "texts": [ " Given )1( 2r$ , )2( 2r$ , )3( 2r$ , any )( 2 i r$ can be expressed as )( 23 )( 22 )( 21 )( 2 i r i r i r i r kkk $$$$ ++= (31) where k1=(s\u03b13c\u03b1i-s\u03b1ic\u03b13+c\u03b12s\u03b1i-c\u03b1is\u03b12-s\u03b13c\u03b12+c\u03b13s\u03b12)/t; k2= (s\u03b13-s\u03b13c\u03b1i-s\u03b1i+s\u03b1ic\u03b13)/ t; k3= (s\u03b1i-s\u03b12-c\u03b12s\u03b1i+c\u03b1is\u03b12)/ t; t= s\u03b13-s\u03b12-s\u03b13c\u03b12+c\u03b13s\u03b12. Thus, the max linear independent number of five $r2 is three. Hence, only four of six wrenches work efficiently. In other words, there are two uncontrollable DoF. Moreover, this type of actuation singularity can be avoided by arranging three limbs with clockwise and the other two with counter-clockwise at assembly configuration. As shown in Fig. 7, limbs 1, 3, 5 are assembled with one current and the other two with contrary current. Furthermore, for all existent 5-DoF 3R2T fully-symmetrical parallel manipulators, it is always a common constraint force is acted on the movable platform at general configurations. Singularity analysis for them is similar because of their similar constraint property. Hence, this study is helpful for singularity analysis of other 5-DoF 3R2T fully-symmetrical parallel manipulators. For example, actuation singularity of manipulator 5-PRR(RR)[3] is the same with 5-RRR(RR)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003175_6.2008-6318-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003175_6.2008-6318-Figure1-1.png", "caption": "FIG. 1. UAV Model in Horizontal Plane.", "texts": [ " In this work, the two-dimensional motion of UAVs in a horizontal plane is analyzed. It is assumed that all UAVs maintain the same altitude throughout the rendezvous process. In the horizontal plane, a UAV\u2019s motion can be completely defined by two variables: the heading angle \u03c6 and the airspeed V. It is assumed that each UAV has the capability to control each of these variables. While many papers assume constant velocity UAVs, there are papers, such as [10], [11], and [12], that assume UAVs have thrusting capabilities. Figure 1 gives a sketch of the horizontal plane UAV model with respect to an inertial Cartesian coordinate frame. The UAV equations of motion with respect to the x-y inertial frame shown in Fig. 1 are defined as sinx V \u03c6=\u027a (1) \u03c6cosVy =\u027a (2) \u03c6\u03c6\u03c6 cossin \u027a\u027a\u027a\u027a VVx += (3) \u03c6\u03c6\u03c6 sincos \u027a\u027a\u027a\u027a VVy \u2212= (4) The angular velocity \u03c6\u027a and velocity rate V\u027a in equations (3) and (4) are taken as the control inputs in this work. When studying scenarios involving multiple UAVs, it is typically required to compute the range and line-of-sight between multiple UAVs. The range and line-of-sight between the i-th and j-th UAVs are ( ) ( )22 ijijij yyxxr \u2212+\u2212= (5) \u2212 \u2212 = \u2212 ij ij ij xx yy 1 tan\u03bb (6) Further, by differentiating equations (5) and (6), one can find the range-rate and line-of-sight rate as ( )( ) ( )( )[ ]ijijijij ij ij yyyyxxxx r r \u027a\u027a\u027a\u027a\u027a \u2212\u2212+\u2212\u2212= 1 (7) ( )( ) ( )( )[ ]ijijijij ij ij xxyyyyxx r \u027a\u027a\u027a\u027a\u027a \u2212\u2212\u2212\u2212\u2212= 2 1 \u03bb (8) The goal of this work is to present a method where a set of Follower UAVs can rendezvous in a finite time in a desired formation about a Leader UAV" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003392_9780470876541.ch5-Figure5.2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003392_9780470876541.ch5-Figure5.2-1.png", "caption": "Figure 5.2 Four-pole surfacemounted permanent magnet synchronous machine (SMPMSM) and its rotor reference d\u2013q frame.", "texts": [ " Also, part of the stator current can be classified as the armature current with regard to the defined rotating flux linkage, which generates anMMFwhose position is perpendicular to the rotating flux linkage. Through the interaction of the rotating flux linkage and MMF by the equivalent armature current, torque can be generated as like a DC machine. At first, the principles of the instantaneous torque control of AC machine can be understood easily from SMPMSM as follows. The torque of an SMPMSM in Fig. 5.2 can be represented as Te \u00bc 3 2 P 2 lf irqs derived in Section 3.3.3.1. Because the pole number of SMPMSM in Fig. 5.2 is four, the d axis and q axis are apart by 45o spatially. After measuring the rotor position instantaneously, if the q-axis current in the measured rotor reference frame is controlled instantaneously as ir * qs \u00bc Te * 3 2 P 2 lf where Te * means the instantaneous torque reference, the torque of SMPMSM can be regulated instantaneously because the flux is kept constant by the permanent magnet. The current in the rotor reference d\u2013q frame, irds and i r qs, can be obtained through the calculation in (5" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002197_6.2007-2727-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002197_6.2007-2727-Figure4-1.png", "caption": "Figure 4. Proportional Navigation.", "texts": [ " Waypoint Tracking PN is a guidance law of an aerial vehicle heading towards a moving target.2,3 It dictates the aircraft to rotate at a rate proportional to the rotation rate of the line of sight, a line connecting the aerial vehicle and the target. One of the main characteristic of this scheme is that this system is simple, because only the rotation rate of the line of sight is needed for navigation. Therefore it is appropriate for small UAVs whose computer resources and measurement hardware are restricted. Fig. 4 shows the basic concept of this system. American Institute of Aeronautics and Astronautics 6 The guidance law of PN is represented as follows: v pn lk\u03c8 \u03b8= && (12) where v\u03c8& is the angle of velocity direction, pnk is a control gain called navigation constant, and l\u03b8 is the angle of the line of sight from a reference line. In this section, the guidance control law of an UAV in cruise phase is derived. Here, an UAV circling around the observation points with a constant altitude is considered. The imaginary targets are placed directly above the observation points, and PN is applied to track these points" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002182_ecctd.2007.4529762-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002182_ecctd.2007.4529762-Figure4-1.png", "caption": "Fig. 4. Simplified Ebers-Moll model of a npn transistor.", "texts": [ " It means finding all the element regions (the kn-values) involved in the occurrence of any operating point. 4) Knowing the specific element regions that must be considered in the DC analysis, the equilibrium system (5) is recast into an iterative linear system and it is solved by considering all these kn-value combinations. Consider the Schmitt-Trigger circuit shown in Fig.3. This circuit contains two transistors, each one of them is modeled by a controlled current source in series with a nonlinear element (a pn junction diode) as shown in Fig.4. The diode iD-vD characteristic is approximated by the PWL curve shown in Fig.5. In order to compute all the operating points in the SchmittTrigger circuit, it is necessary to replace every bipolar transistor by its equivalent circuit shown in Fig.4. Fig.6 shows the resulting PWL network after this device substitution. In this network there are two nonlinear elements and the variables labeled as vDn and iDn (for n = 1, 2) denote the voltages and currents of the PWL elements. After applying nodal analysis, the equilibrium equations (9) and (10) are obtained. vD1 = e2 \u2212 e1 = 3.33 \u2212 1000 (iD1 + iD2) (9) vD2 = e3 \u2212 e1 = 4 \u2212 1392iD1 \u2212 1096iD2 (10) where {e1, e2, e3} are nodal voltages. Notice that equations (9) and (10) can be rewritten into the form of equation (5) as follows: [ 1 0 0 1 ] [ vD1 vD2 ] + [ 1000 1000 1392 1096 ] [ iD1 iD2 ] \u2212 [ 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003994_2011-01-0982-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003994_2011-01-0982-Figure1-1.png", "caption": "Figure 1. Schematic of Vehicle Dynamics Model", "texts": [ " A 7-DOF model is used to represent the handling dynamics for the developed control system. The model includes the degrees of freedom of lateral and longitudinal motions, yaw rotation, and the rotations of the four tires. This model ignores suspension effects and therefore does not consider the pitch, heave, and roll of the vehicle body. Detailed derivations and discussions of the model are given in [7, 8, 12]. The longitudinal, lateral, and yaw equations of motion are: (1) (2) (3) The notations used in (1, 2, 3) are defined in Figure 1. The tire/wheel dynamics connect the wheel torque to the dynamics of the vehicle via the longitudinal tire forces. The tire/wheel dynamics are given by: (4) where i represents left front, right front, left rear, and right rear tires. Each wheel torque is determined from a vehicle speed controller (or driver model) and/or the vehicle stability controller described in the next section. Ideally, the wheel torques, Tw, will be achieved through driving and primarily regenerative braking, however, friction-based braking will be used as a supplement when regenerative braking alone cannot give the desired level of deceleration" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002588_s10015-009-0707-9-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002588_s10015-009-0707-9-Figure2-1.png", "caption": "Fig. 2. Simplifi ed diagram of the tripod parallel mechanism", "texts": [ " This transfor- Key words Tripod parallel mechanism \u00b7 Six degrees of freedom \u00b7 Inverse kinematics \u00b7 Geometrical constraint The Stewart-type parallel mechanism, which achieves 6- DOF motion by the coordinated movement of six actuators, has many advantages compared with the conventional serial link mechanism.1 These are: \u2013 a higher payload-to-weight ratio since the payload is carried by six cylinders in parallel; mation is known as the inverse kinematic problem, and in this case it is the calculation of the positions of the bottom ends of the three links from a given position and orientation of the movable platform. 2.1 Structure of the tripod parallel mechanism As a preparation for the motion analysis, we describe the structure of a tripod parallel mechanism. As shown in Fig. 2, the tripod parallel mechanism consists of a platform (endeffector) and three fi xed-length links, and these are connected by revolute joints. The bottom ends of the three links are given 2-DOF motion on the horizontal plane in any direction. In the case of our experimental setup, this motion can be actualized by a pair of linear actuators. The coordinated motion of the three bottom ends of the links is converted through a spatial mechanism with fi xedlength links into a 6-DOF motion of the platform" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000285_cdc.2003.1271669-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000285_cdc.2003.1271669-Figure7-1.png", "caption": "Fig. 7. Physical meaning of (Y", "texts": [ "03 [m] ....... I 8 I 0 1 2 3 4 5 6 1 8 0 0 .]I ...................................... -0.5 ............... ; . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . t P 6 I I I I 0 1 2 3 4 5 6 7 8 + 3 2 1 c 0 -1 I I I I I - 0 1 2 3 4 5 6 7 8 t B. Stabilizing the angular velocity In this subsection, we try to derive a condition about stable 4 by analyzing zero dynamics. From (18) (22) (27) and (28), zero dynamics becomes as Let a be fixed to simplify (29). Fig. 7 shows situations where output zeroing is achieved. When the center stick keeps rotating in these situations, we have to apply the force which balance with centrifugal force to keep distance T . If a = -7~12, we can apply the force toward rotation center, and this is suitable to realize our control objective. Under Q = -7r/2, we obtain simplified zero dynamics as To analyze zero dynamics, let Cp transform to 4' = 4 - 271.~7 (4jjump 5 4' < 4jump + 2 ~ ) . (31) Using the relations $ = qh, $' = 4, (30) (2) become as follows" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002544_s12206-008-0202-6-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002544_s12206-008-0202-6-Figure4-1.png", "caption": "Fig. 4. Typical cellular material with thick cell struts: (a) plan view; (b) geometry of unit cell.", "texts": [], "surrounding_texts": [ "The standard (implicit) version of the ABAQUS FE code was used to compute the maximum stress (von-Mises stress) on the strut surface and the plastic collapse stress for 2D cellular materials with thick cell struts. Plan views of 2D cellular materials analyzed in this study are shown in Figs. 4 [21] and 5. The FE analysis provides results that are in good agreement with the theoretical results. Table 3. The loads and boundary conditions (for the compliance matrix method) applied to the faces via the associated auxiliary nodes of the unit cell used to calculate the maximum stress on the strut and the plastic collapse stress. Loading direction and boundary conditions Stress factors Node1 Node2 Node3 Node4 Loading Mode 1/e , * 1( ) /pl ys YFIX(1) XCLOAD(2) FREE XFIX Direct (Uniaxial) 2/e , * 2( ) /pl ys YFIX FREE YCLOAD XFIX Direct (Uniaxial) 12/e , * 12( ) /pl ys XFIX(1) YCLOAD(2) XCLOAD YFIX Shear (Uniaxial) 1/e , 2/e 1 / ys , 2 / ys YFIX XCLOAD YCLOAD XFIX Direct (Biaxial) (1): XFIX and YFIX indicate fixity in the X1 and X2 directions. (2): XCLOAD and YCLOAD indicate the application of a concentrated point load equivalent to unit nominal stress on the unit cell in X1 and X2 directions, respectively." ] }, { "image_filename": "designv11_61_0003437_9781118516072.ch4-Figure4.32-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003437_9781118516072.ch4-Figure4.32-1.png", "caption": "Figure 4.32. Pitch angle controller model [8].", "texts": [ "37) If the wind speed increases too much, the electromagnetic torque is no longer sufficient to control the rotor speed from becoming too high and consequently, the generator can be damaged. Turbine blades can also be damaged because of high forces developed at high wind speeds. The solution is to reduce the power extracted from the wind. One way of limiting the forces acting on the turbine blades at wind speeds greater than the rated value is by changing the pitch angle b, thus reducing the performance coefficient Cp\u00f0l;b\u00de [17]. Rotation of blades around their longitudinal axis is performed by either hydraulic or electrical drives. Therefore, a pitch angle controller model (Figure 4.32) has to be integrated in the wind turbine system model [8]. As the wind speed cannot be measured precisely, the input to the controller can be the active power and the rotor speed. The lower part of the pitch controller shown in Figure 4.32 is the rotor speed regulator while the upper part is an aerodynamic power limiter. CONTROL OF WIND POWER EXTRACTION OF A DFIG. The rotor-side converter is responsible for controlling the mechanical torque as well as the stator terminal voltage or the power factor. The variable-speed operation of the DFIG is possible because the power converter decouples the mechanical rotor speed and the power system electrical frequency. The rotor speed control capability facilitates also the optimization of the power extracted from the wind [26,36]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003566_978-3-642-31401-8_14-Figure16-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003566_978-3-642-31401-8_14-Figure16-1.png", "caption": "Fig. 16. The particular of the lace seat in the heel area", "texts": [ " The choice of a semi-rigid rubber is dictated by the requirement that the net cannot collapse on itself when the device is not worn and to prevent tilt modifications due to the foot insertion in the device. This rubber has elastic properties similar to the material used for the upper of the shoes of the case study thus enhancing the feeling of wearing real shoes. Drive acting on a screw regulates the mechanism; by screwing or unscrewing, the length is modified. The last aspect to be described is the insertion of a lace in the heel area, which is shown Figure 16 by the red circle. The lace has the double function to improve the perception of wearing a real shoe and to maintain the FG tied to the foot during the walk. The choices made during the design phase of the FG have allowed to reach all the requirements described in the \u00a73. Tests have been performed both on the wearability and on the walkability of the device, obtaining the results that were expected. Through the compilation of a questionnaire, the testers have also validated that the FG provides the proper dimensional feeling about the shoe dimensions and that the FG is feasible to be used for finding the best fitting shoes" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000089_rspa.2002.1105-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000089_rspa.2002.1105-Figure2-1.png", "caption": "Figure 2. (a) Mode 1, axial slip. (b) Mode 2, rotation slip.", "texts": [ " If there is no slip, the structure is locked and the analysis follows that for other composite-type structures such as laminates and pultrusions. For rope deformation there is a slip between components, the analysis for which is detailed in Leech (2002) and summarized here. Two categories of friction exist, namely inter, where relative deformation occurs between two contiguous components, and intra, where relative deformation occurs within a component. The following slip modes are identified, and illustrated in figure 2. Mode 1. Slip between contiguous yarns and strands in the same layer due to rope stretch and rope twist (figure 2a). This acts axially along the components, but in opposite directions on opposite contact faces. On the component it will produce a shear or couple, whereas on the structure it will oppose the extensional motion. Mode 2. Slip in rotation of a strand/yarn in a rope/strand; the torsion developed within the strand is resisted by the friction torque at the end of the strand. This action opposes the unwinding of a twisted strand from its end. The degree of slip is length dependent since the friction (torque) developed is proportional to the strand length (figure 2b). Mode 3. Scissoring, where the relative angle between crossing strands changes due to rope stretch. This is most applicable in braided/plaited ropes, rope flexure and splices (figure 2c). Mode 4. Sawing due to the action of one yarn over another as they slide, due to rope stretch (figure 2c). This is not significant in geometry-preserving deformations, but since it results from flexure, and since geometry-preserving deformations are accompanied by flexure at the component level, it is present. Modes 5 and 6. Dilation and distortion, occurring as a result of change in strand cross-section as it is stretched in the helix, presses against contiguous strands and is squashed towards the final wedge geometry (figure 2d). Proc. R. Soc. Lond. A (2003) 1644 C. M. Leech These modes can be classified into inter modes (1\u20134), since these act between components, and intra modes (5, 6), since they act within a component. Mode 1 is most dominant for twisted, structured rope loading and for the estimation of hysteresis losses induced; mode 2 acts at a rope termination, break or join, or in the development of a splice. Mode 3 is probably the next most important but only for braided/plaited ropes and mode 4 is very important in rope flexure", " Measurement of friction can be achieved using \u2018yarn on yarn\u2019 testers (Flory et al. 1990) or their derivatives. There are little reported data (TTI 1992) on friction coefficients and what there are have been focused on modes 1 and 3. The friction force is given by the friction coefficient multiplied by the contact force, and this latter force (expressed in N m\u22121) will depend upon the direction of the Proc. R. Soc. Lond. A (2003) Modelling and analysis of splices used in synthetic ropes 1645 T \u03b8 N T scissoring sawing N Fs Fs (c) (d ) Figure 2. (Cont.) (c) Mode 3, scissoring, and mode 4, sawing. (d) Mode 5, dilation, and mode 6 distortion. contact action. For circumferential contact between components in the same layer, contact force = 4n\u03c0(p + t)2r (1 + \u03b5) \u221a 1 + (2\u03c0pr)2 \u00d7 component tension, where p is the pitch (turns m\u22121) of the component about the structure, r is the helix radius, t is the increased twist of the structure, \u03b5 is the strain of the structure, and n is the number of components in a layer. For contact between components in contiguous layers, the contact force is radial and there is no slip; it is given by contact force = 4\u03c0(p + t)2r (1 + \u03b5) \u221a 1 + ((2\u03c0(p + t)r)/(1 + \u03b5))2 \u00d7 component tension" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001494_s10809-005-0129-3-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001494_s10809-005-0129-3-Figure2-1.png", "caption": "Fig. 2. Stripping voltammograms of arsenic(III) and the supporting electrolyte (EDTA) in the presence of oxygen for different potential sweep modes: ( 1 ) potential step mode and ( 2 ) differential pulse mode (pulse amplitude 25 mV, pulse duration 50 ms). The conditions are the following: rotating gold-plated graphite electrode, E acc = \u20131.0 V, accumulation time 60 s, potential sweep rate 80 mV/s, and c As = 0.010 mg/L.", "texts": [ " A search for new conditions for obtaining an arsenic(III) signal in stripping voltammetry at a goldplated graphite electrode was performed. When a vibrating (or rotating) electrode and potential step mode (step height 8 mV, potential sweep rate 80 mV/s) is used, the oxygen reduction current overlaps with the arsenic(III) signal and distorts it. When the gold-plated graphite electrode is rotated at 2000 rpm in a differential pulse mode with 0.1\u20130.02 M EDTA as the supporting electrolyte, oxygen has no influence on the arsenic(III) signal in the concentration range of 0.005\u2013 0.5 mg/L (Fig. 2). The conditions for recording voltammograms are the following: rotating gold-plated graphite electrode, E acc = \u20131.0 V, potential sweep from \u20130.4 V to +0.6 V, differential pulse voltammetry (pulse amplitude 25 mV, pulse duration 50 ms, and potential sweep rate 80 mV/s), and E p = 0.05 V. Water-soluble food supplements were analyzed under these conditions without sample preparation (Table 3). The amount of arsenic(III) found coincides with the amount added within the error of the experiment; that is, the matrix has no effect" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003470_iccasm.2010.5619227-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003470_iccasm.2010.5619227-Figure3-1.png", "caption": "Fig. 3 calculate helix scan paths with the model layer", "texts": [ " In particular, its practical running time seems to grow only linearly. But this algorithm was used to only calculate Voronoi diagram of simply connected polygon, in practice, most of the model is connected region, thus it can use the extended wavefront-propagation algorithm by the author Xu Y P to calculate Voronoi diagram of multiply connected region slice. V5-686 Then Voronoi diagram of the model slice and the Recursive generation of Tool Path algorithm propose by the author in this paper are used to calculate helix scanning path of this slice (See Fig.3). And the Recursive algorithm is based on topological structural of all Voronoi edges and Voronoi polygons, and can accelerate the offset process. The algorithm is simple since it constructs topological relationship of edges and objects, and it is efficient since it takes only linear time and only limited storage for entire process to compute an offsetting in terms of the number of edges of the rings of Voronoi diagram. Finally Tool path is regarded as scanning strategy which is applied to metal powder melting technics, and processes the metal parts layer by layer" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002100_jmes_jour_1969_011_071_02-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002100_jmes_jour_1969_011_071_02-Figure1-1.png", "caption": "Fig. 1. Tip interference at fixed centre distance", "texts": [ " INTRODUCTION THIS paper deals with a type of tip interference not treated in the literature before. It will be proved that in an early stage of cutting, when the centre distance is still too low for involute generating, the cutter injures the internal teeth more severely than existing theories suggest. This phenomenon, called overcut, will be treated starting from the commonly used formulae. Dependent on kinetic conditions three cases are distinguished : (1) tip interference at fixed centre distance with rotating gears, Fig. 1 ; (2) overcut, that is tip interference at changing centre distance during the cutting process; (3) impossibility of radial assembly or disassembly, Fig. 2. A theory for the first case is well known (I)-@) and will be used here as a basis for further development. The second case was mentioned by Dietrich (4): . \u2018(translation) To check the possibility of either assembly of a pinion in an internal gear or manufacture of an internal gear with a certain cutter the calculation has to be made for the various actual centre distances" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001801_tmag.2006.879083-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001801_tmag.2006.879083-Figure1-1.png", "caption": "Fig. 1. Tubular linear actuator with Halbach array.", "texts": [ " Second, by analyzing this circuit using control parameters presented in [3], this paper predicts theoretical resonant frequency and investigates the variation of power factor, current, active and reactive power versus frequency. Finally, experimental results for dynamic characteristics such as current and stoke are presented for various values of frequency. In particular, the methods to predict an accurate resonant frequency are discussed fully in terms of moving mass and spring. Digital Object Identifier 10.1109/TMAG.2006.879083 As shown in Fig. 1, we chose the Halbach array as a mover and a slotless type as a stator of the tubular linear actuator for the reasons as follows. First, the fundamental field of the Halbach array is 1.4 times stronger than that of a conventional array, and thus the power efficiency of the actuator with Halbach array is doubled. Moreover, the magnetic field of the Halbach array is more purely sinusoidal than that of the conventional array, resulting in a simple control structure [4]. Second, the slotless stator eliminates the tooth ripple cogging effect, and thereby improves the dynamic performance and servo characteristics at the expense of a reduction in specific force capability [5]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002674_imece2009-11222-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002674_imece2009-11222-Figure5-1.png", "caption": "Figure 5: Illustration of the parameters to get \u2206n and \u2206s from centroid and face angle information.", "texts": [ " ( )2 x n nn nn ttT T T y t nt nt F F wl R R R wl F F wl \u03c3 \u03bb \u00b5 \u03b5 \u03bb\u03b5 \u03c3 \u00b5\u03b5 + + = = = (8) ( )2 2 N T N N M M n n t t t t n T T N N u u u u u u F wl n s s s s \u03bb \u03bb \u00b5 + \u2212 + \u2212 + \u2212 + \u2212 \u2212 \u2212 \u2212 = + + + \u2206 \u2206 + \u2206 \u2206 + \u2206 (9a) 1 2 N T N N M M t t n n n n t T T N N u u u u u u F wl n s s s s \u00b5 + \u2212 + \u2212 + \u2212 + \u2212 \u2212 \u2212 \u2212 = + + \u2206 \u2206 + \u2206 \u2206 + \u2206 (9b) During the setup the computer reads in a list of nodes and mapping of which nodes make up an element. From this information centroids, face lengths and angles, and neighbors are able to be determined. However since the vectors linking centroids is not always parallel to the line perpendicular to the face the offset must be taken into account as seen in Figure 5. Since the deformation of the cells is assumed negligible by the small strain approximation this calculation is only performed at setup when an adjusted r matrix is created and stored. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 5 Copyright \u00a9 2009 by ASME The vector connecting cell centers has a magnitude r and angle \u03c6 (x-axis datum).The n\u2206 becomes r*cos(\u03c6-\u03b8). However, for the Type II scenario, s + \u2206 and s \u2212 \u2206 combine to become r N *cos(\u03c6 N - \u03b8\u03c0 /2) or r T *cos(\u03c6 T -\u03b8\u03c0 /2)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002061_j.jsv.2006.11.032-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002061_j.jsv.2006.11.032-Figure1-1.png", "caption": "Fig. 1. Slicing the stator and rotor: (a) FE mesh of the stator and rotor and (b) the stator deformation is taken into account by averaging the stator inner core displacements.", "texts": [ " From this data, the current\u2013force frequency responses were calculated by applying Fourier analysis. The parametric model was then fitted to the frequency response data by utilizing least-squares algorithm. The 3D application of the EM force model presented by Eqs. (1) and (2) is accomplished by slicing the stator core and rotor shaft element models with each slice being perpendicular to the rotor shaft. The rotor shaft is modeled by beam elements and the stator bore is a cylinder around the rotor modeled by solid elements (see Fig. 1(a)). The cage current variables couple the slices manifesting implicit continuity conditions for cage currents. In the following the FE meshes of the stator core and rotor shaft are supposed to have a sliced structure consisting of N slices perpendicular to the rotor shaft. On each slice \u2018 \u00bc 1; . . . ;N there is a single rotor node ARTICLE IN PRESS A. Laiho et al. / Journal of Sound and Vibration 302 (2007) 683\u2013698686 with N\u2018 stator bore inner core nodes fn \u2018 sg, s \u00bc 1; . . . ;N\u2018 (see Fig. 1(b)). Let us align the z-axis along with the rotor shaft. In order to apply the 2D EM force model given by Eq. (1) we reduce the stator bore inner core node displacements to a single node by averaging procedure. Indeed, the averaged stator bore inner core transverse displacements in x- and y-directions for slice \u2018 are given by u\u0302\u2018x;s \u00bc 1 N\u2018 XN\u2018 s\u00bc1 ux\u00f0n \u2018 s\u00de, u\u0302\u2018y;s \u00bc 1 N\u2018 XN\u2018 s\u00bc1 uy\u00f0n \u2018 s\u00de, (3) where ux\u00f0n \u2018 s\u00de and uy\u00f0n \u2018 s\u00de are the x- and y-displacement degrees of freedom (dofs) of the node n\u2018s, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002241_978-1-84628-469-4_15-Figure15.3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002241_978-1-84628-469-4_15-Figure15.3-1.png", "caption": "Figure 15.3. Laterial-longitudinal coordinates", "texts": [ " We assume that the obstacles can be detected by a UAV obstacle sensor. The sensor range is shown in Figure 15.2. We are interested in maintaining close tracking of heading speed and heading angle of UAV, which are governed via (.) (.) (.) (.) v vv f g vc f g c (15.1) where and vc c are the commanded velocity and heading angle to the autopilots. and (.)fv (.)f are the nonlinear functions of the system, and(.)gv (.)g represent the control gains. With the lateral and longitudinal coordinates as shown Figure 15.3, we can establish cos sin x l y l , (15.2) where l , in which is the distance between the UAV and the departure point, and v l x and are the lateral and longitudinal positions of UAV with respect to the departure point. Clearly we have y cos ( sin ) sin ( cos ) x v l y v l , (15.3) By defining TX x y , , and T U vc c cos sin sin cos l J l , we get v X J , (15.4) v v X J J (15.5) where sin sin cos cos cos sin v l J v l . Combining Equations (15.1) and (15.5) leads to X F GU , (15.6) where (.) 0 0 (" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000241_ptg-48103-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000241_ptg-48103-Figure5-1.png", "caption": "Fig. 5 Illustration of the change of contact pressure p ijP )( of a point ij as a function or r.", "texts": [ " Relative sliding distance increment as gears rotate from position 1 mr to 2 mr is equal to the distance between pa and ga minus p mmpa s 1)( . If point pa enters the contact zone at position mr and remains within contact zone until position tr , the sliding distance that occurs when gears rotate from any position r to 1 r can be given in general terms as Rrtormr trms s r mq p qqpa p rpa g rgap rrpa 0,0 ,)()()( )( 111 1 XX (13) Noting that for the point pa represented by a discretized node ij, p rijP )( becomes non-zero for the first time at position mr and remains nonzero up to tr as illustrated schematically in Fig. 5. Here, the solid continuous line represents the actual contact pressure p ijP )( that the 5 Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?ur point experiences. The angular span needed for the contact zone to pass through the point ij is very small since the width of the contact ellipse is very narrow. It is also clear that p ijP )( varies significantly, starting from zero and reaching its maximum when the major axis of the contact (line of contact) is at point ij. Therefore, it is vital to have a large R (small p ) such that p rijP )( , which is a discretized snapshot of p ijP )( at position r in Fig. 5, is accurate enough to capture the pressure history of the point. In Fig. 5, R is such that there are 9 discrete positions available to approximate p ijP )( when point ij is in the contact zone. The sliding distance calculations must be carried out only for those nodes with nonzero gp rijP ,)( for at least two consecutive rotational positions. Sliding distance calculations for node ij of gear p are continued as r is increased until gp rijP ,)( becomes zero again. In calculating the sliding distance of point ga on gear g with respect to point pa on gear p, the same procedure is repeated by applying eq", " This iterative procedure is repeated until the maximum total wear depth on either of the two gears reaches a certain maximum allowable wear threshold value of tot . There are fixed surface grid nodes on the tooth that lie below the start of active profile and hence they never enter into the contact zone. For such points, the tooth sliding distance need not be calculated but assigned a value equal to zero. Given the procedure for discretizing the pressures over total sliding of any point on the tooth surface as illustrated in Fig. 5, eq. (2) can be written in the form: gp rij gp rij gp rrij gpgp rrij PPskh , 1 ,, 1 , 2 1, 1 )()()()( (14) Copyright \u00a9 2003 by ASME l=/data/conferences/idetc/cie2003/72062/ on 01/30/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use for a fixed surface grid node ij as gears rotate from position r to 1 r . Thus, the total wear depth reached at any point on the surface in c-th individual wear cycle is 1 0 , 1 , )()( R r gp rrij gp cij hh . (15) Eq. (14) and (15) are applied continuously C times until the maximum wear depth accumulated at any node of either one of the contacting surfaces after the -th pressure update equals " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002821_j.ijsolstr.2008.05.021-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002821_j.ijsolstr.2008.05.021-Figure5-1.png", "caption": "Fig. 5. Example of representation of self-tangential forces considered in the static approach.", "texts": [ " (Antoni and Nguyen, 2008): Z Cc Gq \u00f0h\u00deds \u00bc 0; 8G 8q 2 SF \u00f021\u00de the set SF of self-tangential forces is: q \u00f0h\u00de 2 SF() Z 3p=2 p=2 q \u00f0h\u00dedh \u00bc 0 \u00f022\u00de From (10) and (20), the static expression of the safety coefficient adapted to our problem is the following: ms \u00bc max q \u00f0h\u00de2SF m such that mjqel\u00f0h;b\u00de \u00fe q \u00f0h\u00dej 6 k\u00f0h;b\u00de 8 p=2 6 h 6 3p=2 8 p=2 6 b 6 3p=2 8>><>: \u00f023\u00de The solution of this convex optimization problem can be numerically obtained by standard methods. This resolution will not be considered in this paper since our principal objectives is to derive only some raw estimates of the critical load. Simple estimates could be obtained by a piecewise approximation of q*(h), see also Bj\u00f6rkman and Klarbring (1987). Let us consider here just two piecewise constant elements (see Fig. 5): q \u00f0h\u00de \u00bc q A if p=2 6 h < p=2 q B if p=2 < h 6 3p=2 \u00f024\u00de The condition q*(h) 2 SF is then equivalent to Z 3p=2 p=2 q \u00f0h\u00dedh \u00bc 0() q A \u00bc q B \u00f025\u00de Thus, for a Tresca friction such that k(h,b) = k, we have ems \u00bcmax q m such that k 6 m Q cos\u00f0b\u00de sin\u00f0h\u00de \u00fe b 3p 2 cos\u00f0h b\u00de \u00fe q 6 k 8 p=2 6 h 6 b 8 p=2 6 b 6 p=2 k 6 m Q cos\u00f0b\u00de sin\u00f0h\u00de \u00fe b\u00fe p 2 cos\u00f0h b\u00de \u00fe q 6 k 8b 6 h 6 3p=2 8 p=2 6 b 6 p=2 k 6 m Q cos\u00f0b\u00de sin\u00f0h\u00de \u00fe b 3p 2 cos\u00f0h b\u00de q 6 k 8p=2 6 h 6 b 8p=2 6 b 6 3p=2 k 6 m Q cos\u00f0b\u00de sin\u00f0h\u00de \u00fe b\u00fe p 2 cos\u00f0h b\u00de q 6 k 8b 6 h 6 3p=2 8p=2 6 b 6 3p=2 8>>>>>>><>>>>>>>: 8>>>>>>>>>><>>>>>>>>>: \u00f026\u00de where Q \u00bc F 2pR Sf1 Sf2 St1 St2 1\u00fe St1 St2 " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003411_s11771-012-0984-7-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003411_s11771-012-0984-7-Figure1-1.png", "caption": "Fig. 1 Detailed assembly of an automatic variable preload device [18\u221219]", "texts": [ " Based on the results of the analysis, an improved method that reduces such effects on the performance of the spindle is proposed. Foundation item: Project(2011-0027035) supported by the Basic Science Research Program through the National Research Foundation of Korea (NRF) funded by the Ministry of Education, Science and Technology, Korea Received date: 2011\u221205\u221224; Accepted date: 2011\u221210\u221210 Corresponding author: LEE Choon-Man, Professor, PhD; Tel: +82\u221255\u2212213\u22123622; E-mail: cmlee@changwon.ac.kr J. Cent. South Univ. (2012) 19: 150\u2212154 151 Figure 1 shows the detailed assembly of an automatic variable preload device using an eccentric mass. The device consists of a plate, bolts, and nuts. The principle of applying a preload is presented in Fig. 2. As shown in Fig. 2(a), the initial preload is applied using a constant pressure preload. The constant pressure preload is determined by the fastening force of a spring and lock nut. Figure 2(b) shows the change in the automatic variable preload device when a preload is applied. As the spindle begins to rotate, a centrifugal force is generated" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001952_s00604-007-0859-z-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001952_s00604-007-0859-z-Figure1-1.png", "caption": "Fig. 1. A schematic of the chemiluminescence flow system: (a) water; (b) luminol; (c) K3Fe(CN)6; (S) sample; (P1) peristaltic pump 1; (P2) peristaltic pump 2; (V) injection valve; (M) mixing tube; (F) flow cell; (D) detector; (W) waste; (PC) personal computer", "texts": [ " 3- Amimophthalic acid was purchased from Aldrich. Potassium ferricyanide (Xi\u2019an Chemical reagent Plant, China) stock solution (1.0 10 2 mol L 1) was prepared. A (5.0 10 3 mol L 1 stock solution of caffeine (Drug and Biological Products Examination Bureau of China) was prepared and stored in the refrigerator (4 C), and the working standard solutions were prepared daily from the stock solution by appropriate dilution with H2O immediately before use. Apparatus and procedure The flow system used in this work is shown in Fig. 1. There are two peristaltic pumps: one was used to deliver the flow stream of carrier stream at a flow rate of 3.5 mL min 1 (per tuber), and another was used to deliver the luminol stream and potassium ferricyanide stream at a flow rate of 3.5 mL min 1 (per tuber). PTFE tubing (0.8 mm i.d.) was used as connection material in the flow system. The flow cell is a flat spiral-coiled colorless glass tube (i.d., 1.0 mm; total diameter of the flow cell, 3 cm, without gaps between loops) and placed close to the window of the photomultiplier tuber (PMT)", " In order to apply the SCL emission for detection of reducing agents, the CL system was combined with flow-injection technique. The length of mixing tube of luminol and K3Fe(CN)6 and the concentrations of luminol, K3Fe(CN)6 and NaOH were all optimized for the sensitive and precise detection of reducing agents in the flow system. During each optimization, other conditions were remained constant. In the luminol-K3Fe(CN)6-reducing agents CL system, the PCL emission affected the SCL emission. In the flow system (Fig. 1), the effect of the length of mixing tube of luminol and K3Fe(CN)6 on the SCL emission was very important because reducing agent should be injected to the mixture at that time when the reaction of luminol and K3Fe(CN)6 has been completed sufficiently. So, the effect of the length of mixing tube investigated in the range 15\u2013120 cm at a fixed flow rate of 2.0 mL min 1. The result (Fig. 5) showed that the ratio of signal and blank (S=B) rapidly increased with increasing the mixing tube length in the range of 15\u201360 cm, probably because at too short mixing tube (<60 cm), the PCL emission did not finish and the baseline was higher; or because luminol and K3Fe(CN)6 did not react adequately and no enough intermediate diazaquinone produced", " The effect of K3Fe(CN)6 concentration was tested, and the result showed that when K3Fe(CN)6 concentration was 5.0 10 4 mol L 1, the SCL reaction had the maximum CL intensity. The effect of luminol concentration in the range of 5 10 7\u20132 10 5 mol L 1 was investigated. The result showed that the CL intensity increased with increasing luminol concentration, and above the concentration of 5 10 6 mol L 1, the CL intensity declined. So, the optimal concentration of luminol was 5 10 6 mol L 1. Under the optimum conditions given above and using the flow-injection system described in Fig. 1, the calibration graph of the SCL emission intensity (I, mV) versus caffeine concentration was linear in the range of 5 10 7\u20133 10 4 mol L 1, and the detection limit was 3 10 7 mol L 1. The calibration equations was I\u00bc 24.3\u00fe 72.8 C (C being the caffeine concentration, mg mL 1) with a correlation coefficient of 0.9986 (n\u00bc 8). The relative standard deviation (RSD) was 2.5% for 9 measurements of 5 10 6 mol L 1 caffeine. A complete analysis, including sampling and washing, could be performed in 1 min" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003411_s11771-012-0984-7-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003411_s11771-012-0984-7-Figure7-1.png", "caption": "Fig. 7 Vibration measurement points", "texts": [ "05 mm/s, and an uncertainty of measurement of 0.03 mm/s with its equipment resolution. Also, the bandwidth of this device was 10 Hz\u221210 kHz and the measurement was performed by the unit of mm/s. In Korean Industrial Standards, the most reasonable method for evaluating the broadband vibration of a rotation machine is to consider the RMS (root mean square) value of its vibration speed [20]. This is due to the fact that the vibration speed is related to the vibration energy. Figure 6 shows the configuration of a vibrometer. As shown in Fig. 7, two vibration measurement points were determined to a vertical direction. Also, the spindle speed increased up to 1 000\u22125 000 r/min with an interval of 1 000 r/min. In the experiment environment, temperature and humidity were maintained as 21 \u00b0C and 45%, respectively. Figure 8 presents the values of the measurement at Point 1 and Point 2 according to spindle speeds. As the automatic variable preload device was not installed at a maximum speed of 5 000 r/min, vibrations were recorded by 0.96 mm/s and 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002654_s0263574708004244-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002654_s0263574708004244-Figure1-1.png", "caption": "Fig. 1. Adopted AGV model.", "texts": [ " Dead-reckoning algorithm is very efficient using AGV internal instrumentation signals for autonomous navigation. Its systematical errors are most likely to induce positioning errors. In the work of Borenstein,3 two differential-drive mobile robots are physically connected through a compliantlinkage and using one linear and two rotary encoders provided signals adopted by an internal auto-correcting deadreckoning algorithm. An AGV control scheme for indoors positioning was presented in Azenha et al.13 In this paper, the AGV shown in Fig. 1 is chosen. The kinematic relationships for this type of AGV are (1), which state the dead-reckoning algorithm \u23a1 \u23a3x1odom x2odom \u03c6odom \u23a4 \u23a6 = \u23a1 \u23a2\u23a2\u23a3 x10 + R cos(\u03c60)( \u03b81 + \u03b82)/2 x20 + R sin(\u03c60)( \u03b81 + \u03b82)/2 \u03c60 + R 2b ( \u03b81 \u2212 \u03b82) \u23a4 \u23a5\u23a5\u23a6 (1) where (x1odom, x2odom, \u03c6odom) is the current AGV pose and (x10, x20, \u03c60) is the previous time-step AGV pose. Angles \u03b8i (i = 1, 2) are the wheel i angular position, parameter R is the drive wheel radius and b is as shown in Fig. 1. The AGV pose is defined as the position (x1, x2) and heading or orientation (\u03c6) in the Cartesian space. Figure 2 depicts the current heading (\u03c6) and the target heading (\u03c6t ) and (x1r , x2r ) is the reference AGV position in the Cartesian space. This paper deals with studying the contribution of dead-reckoning systematical induced errors to the final estimated localization error. It is shown that a good heading measurement accuracy contributes strongly to make triangulation more efficient for a differential-drive AGV model, keeping on quality of measurement. Once the order of magnitude for the localization error is estimated, then the need for performing the triangulation task12 is therefore evaluated and quantitatively studied. In the following text, the quantitative evaluation of dead-reckoning systematical errors for the AGV type of Fig. 1 is performed. The two AGV wheels have non-equal radius due to mechanical imperfections, mechanical tolerance, or wheels wearage so radius Ri is assumed to wheel i (i = 1, 2). This source of systematical errors is assumed to be the only one corresponding to dead-reckoning. Heading measurement errors, slips of wheels and uncertainty in other AGV kinematic parameters are not analysed in this paper. A study of dead-reckoning errors of robots caused by noise in the direct reading of wheels speeds can be found in Zhou and Chirikjian", " However, to extend it to the case of mismatched wheels radius, some computations were performed and the adopted AGV dynamic model is (8):[ \u03c41 \u03c42 ] = [ mc2 1b 2 + Ic2 1 + Iw1 mc1c2b 2 \u2212 Ic1c2 mc1c2b 2 \u2212 Ic1c2 mc2 2b 2 + Ic2 2 + Iw2 ] [ \u03b8\u03081 \u03b8\u03082 ] + [ Fv1 0 0 Fv2 ][ \u03b8\u03071 \u03b8\u03072 ] (8) where \u03c4i , Fvi , \u03b8i (i = 1, 2) are the drive dc motors torques, viscous friction coefficients and wheel i angular position, respectively, and: IC = mC 4b2 + l2 12 (9a) Iwi = mw R2 i 2 , i = 1, 2 (9b) Imi = mw ( 3R2 i + t2 w ) , i = 1, 2 (9c) I = IC + 2mwb2 + Im1 + Im2 (9d) m = mC + 2mw (9e) ci = Ri 2b , i = 1, 2 (9f) where l is the AGV length, tw is the transversal wheel length, mc is the AGV mass without wheels and drive motors and mw is the wheel with motor mass, Ri (i = 1, 2) are the drive wheels radius and b is as shown in Fig. 1. This model was derived according to the study presented by Yun and Yamamoto15 and, as it can be verified, significant parameters of the model are dependent on mismatched wheels radius. As presented in Section 2, dead-reckoning AGV navigation algorithm relies on odometry and on heading sensor instrumentation components. In the absence of measurement errors and in the absence of systematical and nonsystematical dead-reckoning errors, only the measurement of wheels angular displacements would be sufficient to accurately locate the AGV" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002481_kem.392-394.30-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002481_kem.392-394.30-Figure2-1.png", "caption": "Fig. 2 Schematic of turning center and sensors location", "texts": [ " ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) \u02c6 \u02c6 \u02c61 [ ] \u02c6 \u02c6 \u02c6\u02c6 \u02c61 [ ] \u02c6 \u02c6\u02c6 \u02c6 \u02c61 [ ] h h w w i i i h h h b b hi i hi h h h v v hi i hi h w k w k S c g a c w k b k b k S t g a t b k v k v k s t g a t v k \u03bc \u03bc \u03bc + = + \u2212\u23a7 \u23aa\u23aa + = + \u2212\u23a8 \u23aa + = + \u2212\u23aa\u23a9 (4) where , ,w b v\u03bc \u03bc \u03bc are the learning rate parameters and satisfies the inequality 1 2 1 0 m w i i c\u03bc \u2212 = \u239b \u239e< < \u239c \u239f \u239d \u23a0 \u2211 and w b v\u03bc \u03bc \u03bc\u2265 \u2265 . )(\u22c5g is the derivative of the nonlinear weighting function )(\u22c5f in (3). The PLS-NN method is applied on a CNC turning center for thermal error modeling. Four thermal-couple sensors are located on the different positions of the machine and those are the spindle t1, the coolant tank t2, the screw nut t3 and the work table t4. And two displacement sensors are set-up for the measurement of X and Z axis direction errors. Fig. 2 shows the schematic of turning center and sensors locations. In this paper the spindle radial(X axis) direction thermal error is modeled as the example for the application of this new modeling method. Thermal errors were measured with the interval of five minutes until the sensors\u2019 readings became changing slowly. The thermal errors are modelled with the proposed PLSNN method, and the fitting and measurment error curves are plotted in Fig. 3 with the residual curve. It can can been that the maximal fitting residual of this moel is about 1 \u03bcm, the average of residual in absolute value is 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002928_icinfa.2009.5205046-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002928_icinfa.2009.5205046-Figure6-1.png", "caption": "Fig. 6. General vision of the simulator system.", "texts": [ " \u2022 RF communication Block: It allowed the establishment of a bi-directional radio link for data communication. It operated in parallel with the commercial platform WIFI link. The objective of these communication links was to allow the use of remote control. The remote control has a high trajectory priority from other blocks, like supervisory control block, and can take the control of the mobile robot to execute, for example, emergency necessary movements or stop. To implement this block was used a low power UHF data transceiver module BiM-433-40. Figure 6 presents a general vision of the considered simulator system. The use of the system has beginning for the captation of main points for generation of the mobile robot trajectory. The idea is to use a system of photographic video camera that catches the image of the environment where the mobile robot navigate. This initial system must be capable to identify the obstacles of the environment and to generate a matrix with some strategical points that will serve of entrance for the system of trajectory generation" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003318_978-3-540-89933-4_2-Figure2.5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003318_978-3-540-89933-4_2-Figure2.5-1.png", "caption": "Fig. 2.5. Vector representation of PSO", "texts": [ " Vnew = w\u00d7Vcur + c1 \u00d7 rand1 \u00d7 (Pbest \u2212 Pcur)+ c2 \u00d7 rand2 \u00d7 (Lbest \u2212 Pcur) (2.2) Pnew = Pcur +Vnew (2.3) Where, Vnew - New velocity calculated for each robot Vcur - Velocity of the robot from the current iteration Pnew - New position calculated for each robot Pcur - Position of the robot from the current iteration w - Inertia weight c1 and c2 rand - Cognitive and social acceleration constants respectively - Generates a uniform random value in the range [0 1] A vector representation of the PSO algorithm in a 2-dimensional search space is given in Fig. 2.5. The robots are represented by stars and the target is represented by the labeled circle. The vectors Vpd and Vgd represent the effect of the \u2018Pbest\u2019 and \u2018Lbest\u2019 positions on the robots, respectively. To associate Eq. (2.2) with Fig. 2.5, the term Vpd in the figure represents \u2018c1 \u00d7rand \u00d7(Pbest \u2212Pcur)\u2019 and the term Vgd in the figure represents \u2018c2 \u00d7 rand \u00d7 (Lbest \u2212Pcur)\u2019. The robots use the values of \u2018Pbest\u2019 and \u2018Lbest\u2019 to find a new position closer to the target. The PSO algorithm can be summarized into the following steps: 1. Randomly initialize a swarm of robots in a defined n-dimensional search space with initial random velocities, Pbest , and Lbest values. 2. Evaluate each robot\u2019s current position using the fitness function (the Euclidean distance from the robot to the target(s) based on sensory readings)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000035_robot.1999.772527-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000035_robot.1999.772527-Figure4-1.png", "caption": "Figure 4: Determination of ME'S finger-tip trajectory.", "texts": [ "3 Finger-Tip Trajectory Prediction For an ME to work independently, it should also predict its finger-tip trajectories. This is also important for the reason that an ME can know only two variables, i.e. its own F, and Fy, while three variables are desired corresponding to the three object motions. Therefore, it is needed to restrict the motions of an ME to its own controller only, instead of the object motions. However, an ME needs to consider the object's trajectories while predicting its own. Consider again an ME in contact with the object as seen by the reference world frame, shown in Figure 4. As the motions of the object are considered in the world frame, hence the finger-tip trajectories can be found easily, knowing the position of the finger-tip with respect to the object. This is true when there is no slip a t the contact-points. Let \"pa = [ O x a '%IT be the position of the center of mass of the object and & be the rotation of the object frame E, with respect to the world frame, i.e. object's orientation. Then the finger-tip trajectory of i-th ME can be found easily as Opfz = \"pa + \"%(" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001408_j.mechatronics.2006.02.001-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001408_j.mechatronics.2006.02.001-Figure2-1.png", "caption": "Fig. 2. Model of golf swing robot.", "texts": [ " In the passive joint, there is no actuator but a mechanical stopper is used to realize cocking and uncocking of the golf club in the same way as the human wrist. A resolver is mounted on each joint to measure the rotation angle. The human skill of motion control, that is, two-step acceleration by dynamic coupling drive can be realized by the mechanism. First, the actuated joint is accelerated by the DD motor while the passive joint is cocked by a mechanical stopper in the initial acceleration period from top position to uncock position (Fig. 2). When the dynamic coupling torque exerted on the club is toward the direction of uncocking the club, the club is accelerated by the coupling torque and detaches from the stopper after the uncock position. For simplicity, the process after impact is considered as a scaled-inverse process of that up to the impact. The control of the robot is an under-actuated problem, that is, using one actuated joint to control motions of two joints. The control input and the motion trajectory must be generated from both kinematics and dynamics" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001142_cdc.2005.1582306-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001142_cdc.2005.1582306-Figure3-1.png", "caption": "Fig. 3. The adapted H-bridge setup scheme, top view.", "texts": [ " The controller (3) admits some freedom in the choice of the matrix K. This freedom can be used, for example, in tuning the controller to obtain desirable performance of the closed-loop system. Controller (3) is implemented in the experimental setup described in the next section. The experimental setup has been constructed by adapting an existing X-Y positioning system (the H-bridge setup) in the Dynamics and Control Technology Laboratory at Eindhoven University of Technology. The setup is shown in Fig. 2. The adapted H-bridge setup is schematically shown in Fig. 3. It consists of the following components. The two parallel axes Y1 and Y2 are equipped with Linear Magnetic Motor Systems LiMMS Y1 and LiMMS Y2 that can move along their axes. These two carriages support the X-axis. In all experiments that are performed on this setup, the Y1 and Y2 carriages are controlled to maintain a fixed position with a low-level PID controller. Therefore, in the sequel we assume that these two carriages stand still and that the X-axis is fixed. In the sequel we will refer to the X-LiMMS carriage moving along the X-axis as the cart" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001288_10402000600781382-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001288_10402000600781382-Figure3-1.png", "caption": "Fig. 3\u2014Geometry of a Hertz contact circle.", "texts": [ " Thus, a four-parameter model, which is in unidirectional shear, can be obtained as follows (Evans and Johnson (11)): \u03b3\u0307 = 1 Ge d\u03c4 dt + \u03c40 \u03b7eq sin h ( \u03c4 \u03c40 ) for \u03c4 < \u03c4L [1a] \u03c4 = \u03c4L for \u03c4 \u2265 \u03c4L [1b] with the four fluid parameters: the effective elastic shear modulus Ge of fluid plus ball and disc,the viscosity \u03b7eq, the reference Eyring stress \u03c40, and the limiting stress \u03c4L. The shear stress at an operating condition can be evaluated by determining these fluid parameters and solving Eq. [1]. A schematic of the contact geometry between the ball and disc in full EHL regime is shown in Fig. 3 together with the coordinate system. With no spin and lateral slide, many two-dimensional shear planes are considered for any ordinate y. The following assumptions are made: D ow nl oa de d by [ U ni ve rs ity o f U ta h] a t 0 3: 35 0 7 O ct ob er 2 01 4 1. Shear stress is considered only in the parallel region, and the flow caused by pressure gradient is ignored. 2. The base oil film thickness, which is assumed to be uniform in the Hertzian contact circle, is computed in this paper since the grease film thickness is close to its base oil film (Ge\u0301rard and Rene\u0301 (6)), and its value is the central film thickness according to Hamrock-Downson\u2019s equation (10): hiso R = 2", " The correction factor derived by Johnson and Cameron (13) can be conveniently adopted, and it is approximately 2.2, 2.1, 2 respectively for different rolling speeds of 20 m/s, 30 m/s, and 40 m/s. It should be pointed out, however, that the purpose of the mean effective equilibrium viscosity or the apparent viscosity above is to obtain the approximate values of the mean shear modulus. In the traction calculations the expression shown in Eq. [9] is used as the lubricant viscosity. The prediction values of traction force F can be estimated by integrating \u03c4\u0304 over the contact circle shown in Fig. 3. The corresponding values of the traction coefficient can be obtained by D ow nl oa de d by [ U ni ve rs ity o f U ta h] a t 0 3: 35 0 7 O ct ob er 2 01 4 Eq. [22]. F = \u222b a \u2212a (\u222b a+ \u221a a2\u2212y2 a\u2212 \u221a a2\u2212y2 \u03c40\u03c4\u0304dx ) dy [21] \u00b5 = F W [22] To illustrate the feasibility of the method for determining the rheological parameters of the grease and the reasonableness of the algorithm, six testing conditions were selected to carry out the comparison between model prediction and experimental observation. Figures 4 and 5 show the experimental traction curves obtained with the two lithium greases No" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001080_oceans.2005.1639997-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001080_oceans.2005.1639997-Figure1-1.png", "caption": "Fig. 1. One degree of freedom MMA.", "texts": [ " In conventional vehicles, the control moment is generated by control surfaces. Both the control moment and the lift generated by the body vary quadratically with speed so that, at some minimum speed, the body is unable to maintain depth. A moving mass actuator (MMA) can provide control mo- ments that are independent of vehicle speed. In this paper, we consider the problem of controlling longitudinal motion of a conventional streamlined AUV at low speed, where stern planes are prone to saturate. Consider the modular, one degree of freedom MMA shown in Fig. 1. When secured below a streamlined vehicle, such as the Virginia Tech Miniature AUV (VTMAUV) shown in Fig. 2, the actuator provides a constant gravitational pitch moment that depends only on the size and longitudinal location of the moving mass. A properly sized MMA can therefore extend the lower speed limit without greatly diminishing the vehicle\u2019s efficiency at its nominal speed. A two degree of freedom MMA can provide control in pitch and roll. For a neutrally buoyant vehicle, for example, a two degree of freedom MMA would enable tilt control in hover" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001139_iros.1992.587374-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001139_iros.1992.587374-Figure1-1.png", "caption": "Fig. 1: A Cartesian position of the TCP can be reached by 4 different kineniat,ic states.", "texts": [ "ce definition differ relating to the description of the robot\u2018s configura.t,ion. Often such a configura,tion is defined by the position of the joint,s of the robot. Thereby a unique specification of the robot\u2019s shape is possible, but. t>he resulting c-space ha.s no direct correspondence to the workspace. That, is the reason why we describe a configuration by the position of the tool-center-point. (TCP) [GI . Unfortunately t,he positlion of the TCP is not, related uniquely to a specific robot\u2019s shape - which is iniporta.nt, for t,he collision check - a.s shown in fig. 1. To overcome this problem an addit,ional coordinate G jkiiiemat,ic state) is introduced according t,o fig. 1. The symmetry of the robot\u2019s workspace wit,h respect t,o the first axis of t81ie robot, is t,a.ken int,o consideration by choosing cylindrical coordinates to describe t8he T C P (fig. 2). Furthermore a disc.ret,izat,ioii of work-space as well as c-spa.ce into \u2018elementary cells\u2019 is int,roduced. In the following an eleinenhry cell in work-spa.ce is denot,ed with e = ( T , \u2019p, z ) * , a cell in c-space with s = ( ,r , p, z , GIT. B. Classification of objects In order to minimize the computation t ime for obstacle ma", "ssificat 1) then which results in the manipulator work point being out of reach. a2 + d4 is related to the 8, singularity state of Cybotech W 1 5 [l]. Infact, desire to keep H in the interval 10, (aZ + d4)], for good maneuverability, we may choose H > a2 + d4 (17) Therefore we can find H 2 H = C = 2~ (a2 + d4) (18) 2 In this case the robot is able to stretch its arm back and forth and still avoid the singularity state of 83, the robot is then said to have good maneuverability. Notice also additional constraints exists on the solvability of 8,, these are: pX # 0 and pY # 0 (19) SO, equivalently, (p: +p$) > 61 > o (20) will seek to keep the manipulator out of the circle radius 6, centered about its base, hence avoid 01 singularities. If 8, is zero then joint rotations about axis four (first wrist roll) aligns with the rotations about axis six (the final wrist roll). At that time the rotation of joint four becomes colinear with the rotations of joint six. This is a singularity state, therefore we desire: 185 I > 65 (21) where the constant 65 is selected to produce desirable motion characteristics of the wrist, see [9]. IV. Singularity Avoidance And Control of the Redundant Joints We desire to maintain the robot close as possible to good maneuverability throughout the motion of the arm. This may be. achieved through appropriate coordination of the redundant joints. We will formulate this as a mathematical programming problem. The equality constraints is that: OR, O P W Ttrack B'T6 = [ 0 1 1 (22) 0 where Tt rack is the hoqogeneous transformation representing the track and BTet is the homogeneous transformation of the W 1 5 robot. The subscript w is used to denote the workposition, and the superscript 0 is used to denote the origin. This constraint can be subdivided further into: BR6 = O R w (23) and, ',P = (oRtrack)t (O_P - O _ P ) (24) T t rack consists of two transforma6 W track The transformation tions Z and &, i.e. (25) Ttrack = z& 0 ,e 3 4 where, Z = [\"d ';]and A,,= where ,e 3 = [0 0 lIt and Rz = rot(z,8,) , I T also, Ro = rot(x, - -) 2 and ,P = trans(x',aq) These transformation are defined for our example in Figure 6. O P - = P 5 +R,_e3dt (30) track z an& B'g = (RZRo)t(oPw - 'P, - RZ,e 3dt) (31) A mathematical programming model may now be constructed as follows: Let f(_x ) = Hz - C2 = (p: + p; + (p, - d,)' - C2)(32) Then, minimize f2(z ) with the equality constraints: 6 h ( _ x ) = [Px, Pyr PI] - (R,Ro)t [ o r -\"E _e 3 4 1 = 433) w z a and the inequality constraints: 6: - (P: + Pi ) g W = [ s5-p51 1. b] (34) where This is a standard problem of constrained minimization and a solution methodology exists [12). 5 = [Px, Py, Pz, dtlt 'Note here that , B'T6 = Rb BT6 Simplification of t h e Constrained Minimization for Singularity Avoidance a n d Redundancy Control On examination of the Kuhn-Tucker conditions we know that the inequality constraints need only be considered when they are active, This reduces to a practical strategy which is as follows: Find solution for fonc of the equality constraint, as noted in the above check if the inequality constraint is violated. If not discard the inequality constraint (iii) case(b) If inequality constraint is violated revize the solution to include inequality constraint. Intuitively, if we are far from 81 or 85 singularity than we may move the manipulator to maintain good maneuverability with respect to H. This is achieved by suitable motion of the track dt which satisfy the equality constraint. If the inequality constraints are violated, the track is moved to keep the manipulator at a safe boundary from the singularity conditions through considerations in case (b). We now find the analytic solution for both cases: Case(a): lem. Using Lagrange multipliers: Solution of t h e equality constraint prob- (a) f(x) = 0 or (b) ,e RO [pX, pY, pZ - d1I = 0 (35) 3 This results in two solution of dt, dt, and dtb: dtb =(E - )t R z ~ 3 -R0(3,3)dl (36) W z d,, = dlb i vd& 4- c2 - H2 - (a: - 2a,dx: + y i - d:b) These two solution of the track corresponds to two different regions in which they can be applied. There are two values of dt, in the region H = C, however when H < C there is only one solution dtb. Case (b): Solution with t h e Inequality Constraint Relating t o 8, (37) Now one more constraint added to the above: g(_. ) = 6:: - (P: + p:) s 0 (38) On consideration of, the feasible solution regions of 2 , h ( 2 ) = 0 and g ( 2 ) = 0, dt has only two possible solu- tions, where the constants are - CY = (CY, ,~y ,@z) = - (RZRdt(_F - ) - P = (P,,P,,PJ = P W , ~ ) , ~ o ( $ 2 ) , ~ J 3 , 3 ) ) Therefore the feasible region only contains two points and we need to choose a value of dt which makes the objective function the smallest f2(2 ) = (6; + (az + Pzdt, - d1)\u2019 - C2)\u2019 (40) Case (b): Inequality Cons t r a in t Related t o 05 expressed in terms of Cartesian parameters. (1) Execute equality constraints problem, disregard constraints related to g( ,x ) (2) Check inequality constraint, if it is not violated the solution produced by step (1) is acceptable (3) Otherwise solve a set of nonlinear equations h(_x ) = 0, g ( 2 ) = 0, to find feasible regions of dt. After substitution one nonlinear equation will remain. The solution to this may than be found by Brent\u2019s algorithm (see IMSL Library documentation). The constraint g ( 5 ) = 6, . 18, I S 0 can be The solution algorithm is as follows: VI. Conclusions In this paper we have presented an algorithm to coordinate a welding table and a seven degree of freedom manipulator. The motions of the table are constrained by the down hand welding. The motions of the redundant manipulator is selected from a Cartesian coordinate nonlinear optimization process to avoid robot singularities. This algorithm does not utilize generalized jacobian inverses like previously proposed schemes [4], [5], [7], [lo], 1111, [13], [14!. This way achieves the desired motion accuracies while only utilizing inverse kinematics. Our methodology is novel in that respect and it relies on our previous results on characterization of the manipulator singularities in terms of the Cartesian coordinates of the end effector. Simulation experiments have been carried out to verify this but have not been presented in this paper because of space limitations. Acknowledgements We would like to thank all our friends at Cybotech for many helpful discussion. Especial thanks go to Mike McEvoy, Paul Stahura, Rick Guptill, Larry Biehl and Jerry Mitchell. References Ahmad, S., Luo, S., \u201cAnalysis of Almost General Robots,\u201d TR-EE-87-19, Purdue University, West Lafayette, IN 47906. Chang, P.H., \u201cA closed-form solution for the control of manipulators with kinematic redundancy,\u201d 1986 IEEE Intl. Conf. on Robotics & Automation, p. 9. Freund, E., Hoyer, H., \u201cCollision avoidance in multi-robot systems,\u201d Robotics Research: 2nd International Symposium 1985, Ch. 3, pp. 135-146. Klein, C.A., Huang, C.H., \u201cReview of pseudo inverse control for use with kinematically redundant manipulators,\u201d IEEE V. SMC-13, p. 245. Liegeois, A., \u201cAutomatic supervisory control of the configuration and behavior of multibody mechanisms,\u201d IEEE V. SMC-7, p. 868. Tsai, L.W., Morgan, A.P., \u201cSolving the kinematics of the most general six- and five-degree-of-freedom manipulators by continuation method,\u201d J. of Mechanism, Transmissions, and Automation in Design, June 1985, Vol. 107, pp. 189-200. Yoshikawa, T., \u201cAnalysis and control of robot manipulators with redundancy,\u201d Robotics Research: International Symposium, 1984. [8] Zheng, Y.F., Luh, J.Y.S., \u201cControl of two coordinated robots in motion,\u201d Proc. 24th IEEE Conf. on Decision and Control, Florida. Stevenson, C., Paul, R.P., \u201cKinematics of Robot Wrists,\u201d Int. J. of Robotics Research, Vol. 2, No. 3, [IO] Whitney, D.E., \u201dResolved motion rate control of manipulators and human protheses,\u201d IEEE Transactions on Man-Machine Systems MMS-10, 1969, pp. 47-53. [11] Nakamura, Y., Hanufusa, H., \u201cInverse kinematic solutions with singularity robustness for robot manipulator control,\u201d ASME Journal of Dynamic Systems [9] 1983, pp. 31-38. Measurement and Control, September 1986, Vol. [12] Luenberger, D.G., \u201cLinear and Nonlinear Programming,\u201d Addison Wesley, Menlo Park, California, 1984. [13] Luh, J.Y.S., Gu, L.Y., \u201cIndustrial Robots with Seven Joints,\u201d Proceedings of the IEEE Conference on Robotics and Automation, St. Louis, MO, pp. 1010- 1015. [14] Fisher, W.D., \u201cThe Kinematic Control of Redundant Manipulators,\u201d Ph.D. Thesis, Purdue University, Aug. 1984. 198, pp. 163-171. 85" ] }, { "image_filename": "designv11_61_0002928_icinfa.2009.5205046-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002928_icinfa.2009.5205046-Figure3-1.png", "caption": "Fig. 3. Forward kinematics geometry.", "texts": [ " This paper presents the virtual environment implementation for project simulation and conception of supervision and control systems for mobile robots and focus on the study of the mobile robot platform, with differential driving wheels mounted on the same axis and a free castor front wheel, whose prototype used to validate the proposal system is depicted in Fig. 1 and Fig. 2 illustrate the elements of the platform. Suppose that the robot is at some position (x, y) and \u201cfacing\u201d along a line making an angle \u03b8 with the x axis (Fig. 3). Through manipulation of the control parameters ve and vd, the robot can be made to move to different poses. Determining the pose that is reachable given the control parameters is know as the forward kinematics problem for the robot. Because ve and vd and hence R and \u03c9 are functions of time, is straightforward to show (Fig. 3) that, if the robot has pose (x, y, \u03b8) at some time t, and if the left and right wheels have ground-contact velocities ve and vd 978-1-4244-3608-8/09/$25.00 \u00a9 2009 IEEE. 899 during the period t \u2192 t + \u03b4t, the ICC (Instantaneous Center of Curvature) is given by ICC = [x \u2212 R sin(\u03b8), y + R cos(\u03b8)]. (1) For better equation write, we can simplifying ICC = I , cos(\u03c9\u03b4t) = C and sin(\u03c9\u03b4t) = S, then, at time t \u2192 t + \u03b4t, the pose of the robot is given by \u23a1 \u23a3 x\u2032 y\u2032 \u03b8\u2032 \u23a4 \u23a6 = \u23a1 \u23a3 C \u2212S 0 S C 0 0 0 1 \u23a4 \u23a6 \u23a1 \u23a3 x \u2212 Ix y \u2212 Iy \u03b8 \u23a4 \u23a6 + \u23a1 \u23a3 Ix Iy \u03c9\u03b4t \u23a4 \u23a6 " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000985_ichr.2004.1442686-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000985_ichr.2004.1442686-Figure8-1.png", "caption": "Fig. 8. Reliability based on change in IR information taken from pyroelectric IR sensor", "texts": [ " The variables and the distribution are shown in Fig. 7. While the robot is moving, au is 0. Pyroelectn'c infrared Sensors The pyroelectric infra-red sensors are used to detect the moving objects by measuring changes in inka-red readings. The sensors' output is either l (detected) or 0 (not detected). We define these sensors\u2019 reliability distribution, fp(z), as 0.3 x (-30 5 xP 5 30), f p ( z ) { 0 (otherwise), (15) where ap is the weight value 0 whiIe the robot is moving and 1 while it is not moving. The distribution is shown in Fig. 8. After calculating all reliability distributions for all sensors, we combine the distributions into a reliability distribution of human existence around the robot by the following equation. 3.3. MCMCpmceas We predict the next positions of humans and limit the ranges of each sensor by utilizing particle filtering method [9,10] that is based on MCMC. A process of the prediction and the limitation is as follows: [l) Acquire a weighted sample-set {(sly\\, T!:\u2019,), n = 1,. . . , N ) a t time-step t - 1 from a previous iteration" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001246_iros.2006.281695-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001246_iros.2006.281695-Figure2-1.png", "caption": "Figure 2. Passive dynamic passenger model", "texts": [ " The effectiveness of the proposed method is confirmed through the experiments using WL- 16RII. II. IDENTIFICATION OF A PASSIVE DYNAMIC The passenger seat of WL-16RII is mounted on the waist through a force-torque sensor, and moves on a plane during walking motion over a flat and even surface. Therefore within the passenger's body, the lower limbs are assumed to be fixed on the robot, and the upper body is assumed to be a single particle with 2 DOF mounted on the seat, through 2 springs and dampers (Figure 2). The weight of this particle can be computed from the passenger's body weight, based on data stating that the human body's weight ratio of the upper body, including the head and upper limbs, is 65.7 % for a male and 63.9 % for a female [18] respectively. The equation of motion is as follows: mhrh (t) + C(t)(rh (t) - rw (t)) +K(t)(rh(t) -rw(t))= 0 (2) where mh is the weight of the upper body particle, rh (t) is the position of the same and rW(t) is the position of the robot's waist. K(t)=diag(K,(t),Ky(t),O) is the stiffness variable matrix and C(t)=diag(Cx(t),Cy(t),O) is the damping variable matrix respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002236_2007-01-3740-Figure16-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002236_2007-01-3740-Figure16-1.png", "caption": "Figure 16b: TE-IVT Concept Dimensions", "texts": [], "surrounding_texts": [ "Torotrak\u2019s full toroidal traction drive technology has achieved series production in the Outdoor Power Equipment market validating the traction drive technology, the components and \u2018two roller\u2019 Variator design. The Variator has also been proven to be scaleable and configurable for higher torque and power applications. A number of arrangements have therefore been designed to satisfy the requirements for low cost, efficient transmission solutions in the constrictive packaging space of transverse FWD vehicle applications." ] }, { "image_filename": "designv11_61_0000411_j.sysconle.2005.09.001-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000411_j.sysconle.2005.09.001-Figure4-1.png", "caption": "Fig. 4. Inverted pendulum in Example 1.", "texts": [ " In the other words, the recurrent neural network is, in fact, a closed-loop dynamical system without any inputs and with some initial conditions that are determined by the mentioned matrixes. Remark 5. From Eqs. (40) and (55), we see the speed of convergence rate in Lyapunov theorem is related to k and a larger k results a smaller convergence time. In this section, simulation results are given to demonstrate and verify the theoretical results discussed in previous sections. Example 1. In this example, we design an online pole placement to control a nonlinear system by instantaneous linearization of Eq. (27) (instantaneous pole placement). We have chosen an inverted pendulum (Fig. 4) with the nonlinear state as x\u03071 = x2, x\u03072 = \u2212g l sin(x1) \u2212 k m x2 + 1 ml2 T , (57) where y = x1 is output of system (angle in radians), u = T is the control input, x2 is the velocity (in rad/s), g =9.81, l =1.0, m = 1.0 and k = 0.5. Here there are two equilibrium points, (0, 0) and ( , 0) which are stable and unstable, respectively. As mentioned above we first transport unstable equilibrium point to the origin and then apply instantaneous pole placement, so we have x\u03071 = x2, x\u03072 = g l sin(x1) \u2212 k m x2 + 1 ml2 T ", " In this example, we show accuracy of this method for a third-order system with three inputs as A = \u23a1 \u23a3 \u22121 1 2 0 \u22123 \u22122 \u22120.1 \u22121 \u22122 \u23a4 \u23a6 , b = [1 0 0 0 \u22123 0 2 1 2 ] . First, we choose a stable closed-loop matrix with desired eigenvalues: A = \u23a1 \u23a2\u23a3 \u22122 0 0 0 \u22123 0 0 0 \u22125 \u23a4 \u23a5\u23a6 . Fig. 7 shows the output of recurrent neural network of Eq. (35) for this system. Example 4. In this example, we design an online pole placement with recurrent neural network to control a nonlinear system by instantaneous linearization. We have chosen the inverted pendulum of Example 1 (Fig. 4) with the same values for parameters and linearized system of Eq. (59). Fig. 8 shows the simulation result with several initial conditions without any controller. We see that the output of system converges to the stable point at ( , 0). Now we apply instantaneous pole placement as a controller to this system by using recurrent neural network. Because B is a time-invariant matrix, the results of Theorem 2 with recurrent neural network of Eq. (35) have been used for simulation. The following matrixes are chosen for Acl and B: Acl = [\u22122 0 0 \u22123 ] , B = [0 1 1 0 ] " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002253_978-3-642-04466-3_4-Figure16-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002253_978-3-642-04466-3_4-Figure16-1.png", "caption": "Fig. 16 Shear stress contours (Pa) on the rear of Balls 1 and 2, with separation points marked in black (indicating separation) and red (indicating separation at a seam)", "texts": [ " AS an example, Fig. 15 shows the oil flow visualisation (from four different views) around Ball 1 at 0\u0131 for Re D 1:0 106. This demonstrates the pathlines of particles released from the surface and shows clearly the separation position towards the rear of the ball. Each particle had its own \u201cparticle ID\u201d and the particles were coloured according to their ID. For each ball, images were then built up that indicated the position of separation at all points around the ball, and the influence of the seams were seen. Figure 16 shows shear stress contours (Pa) on the rear surface of Ball 1 and Ball 2, along with approximate separation points obtained from the oil flow visualisation, for x deg and x deg as examples. Black lines indicate approximate separation and grey lines indicate separation where the flow is particularly affected by the presence of a seam. The results suggest that seams that are perpendicular, or nearly perpendicular, to the flow had an effect on the position of separation, due to the sudden change in curvature of the surface", " In general, the seams of Ball 1 that were perpendicular (or nearly perpendicular) to the flow seem to have been more likely to hold back the flow and alter its position of separation than the seams of Ball 2. This was probably because they were larger and had more influence on the flow. This meant that the CS varied more often with angle for Ball 1. The seams of Ball 2 only rarely held back the flow and influenced separation, but was sometimes swayed heavily by the presence of long, perpendicular seams, e.g. at 40\u0131 and 50\u0131 (Fig. 16). The magnitude of CS was correspondingly lower for Ball 2. The significant \u201cstepped\u201d pattern in CS occurred when particular seams held back the flow in a certain position, and continued to do so as the ball was rotated. Eventually it reached a point where the seam suddenly lost its influence because it was too far away from the natural position of separation (i.e. where separation would have occurred without the presence of the seams), and the separation point then retreated to the previous perpendicular seam" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003678_tmag.2012.2201254-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003678_tmag.2012.2201254-Figure3-1.png", "caption": "Fig. 3. Simple example to detect the numerical error from mesh rotation.", "texts": [ " The merit of this method is that on the two edges of the period boundaries, the local coordinates of the nodes do not need to be the same. That means, when the mesh is generated, it is not necessary to consider the periodic boundary conditions. This simplifies the process of mesh generation with acceptable accuracy. A mesh with such antiperiodic boundary conditions, from the Team Workshops Problem No. 30 [11] and which will be used to verify the proposed method as described in Section B of the following paragraphs, is given in Fig. 2. A simple example as shown in Fig. 3 is used to study the numerical errors arising from rotor rotation. The stator and rotor are made of iron and their conductivity is set to zero. There is a 100-turn coil carrying a dc current of 100 A in the airgap. Firstly a coarse mesh with 903 triangles is used (Fig. 4(a)). In this study the rotor rotates at a speed of 360 degrees/s. The time step size is 0.001 s. For the first test, there are 90 master nodes and 90 slave nodes on the sliding interface. Interpolationmethod is used to deal with the matching boundary condition [7]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001530_icsmc.2004.1400740-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001530_icsmc.2004.1400740-Figure1-1.png", "caption": "Figure 1: An illustration of the 3DOF helicopter.", "texts": [ " The main par t of the controller is a real-time lookahead simulation of linearized plant dynamics with a feedback law. The paper also presents an evaluation of the remote laboratory facilities in a form of a comparison of the student experiences performing the experiments in a traditional laboratory setting and performing the same experiments using the remote laboratory. Keywords: Education; remote laboratory, system safety, Internet. 1 Introduction Since 1998, the MIT Department of Aeronautics and Astronautics has provided a laboratory facility with the 3 Degree-of-Freedom (3DOF) helicopters (see Figure 1) for the students in control classes. Three types of experiments are usually performed in the laboratory: a plant identification experiment, a lDOF and a 3DOF control experiments. The equipment has been used each year for a period of six weeks by a hundred students on average. Over the lifetime of the lab, two machines have had to be replaced because of the unintentional misuse by the students. The 3DOF helicopter is a complex system, which can be easily damaged if it hits the table support or stresses its own joints", "he students performing the traditional laboratory as the baseline for evaluating the feedback from the students performing the remote laboratory with the supervisory controller. 2 Laboratory Description As the first step in supervisory controller design, we develop a nonlinear dynamic model. Then, we present the design of a structured supervisory controller. Lastly, we describe the controller implementation as part of a remote laboratory. 2.1 System Model The 3DOF helicopter used in this research is shown in Figure 1. The helicopter is mounted on a table top, and its primary components are the main beam, the twin rotor assembly and the counterweight. The main beam is mounted on a bearing and slip-ring assembly which allows the rotor assembly to rotate in continuous circles without entangling the wires. This rotational motion is called travel ($). It occurs about the vertical axis which goes through the slipring and is perpendicular to the table. At the hearing and slip-ring assembly, there is a pivot point which allows the main beam to raise and lower" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002897_tmag.2007.916494-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002897_tmag.2007.916494-Figure4-1.png", "caption": "Fig. 4. Pulse magnetizer model.", "texts": [ " When the external magnetic field , which has the angle of from the easy axis, is applied to the single domain particle, the total energy can be written as follows: (13) where the first term of the right-hand side of (13) is crystal anisotropic energy, the second term is the shape anisotropic energy, and the third term is the magnetic energy. and are the demagnetizing factors: denotes the major axis and denotes the minor axis. is the anisotropy field. The angle of in each substance (finite element) can be calculated by the minimization condition of the total energy as follows: (14) To solve this equation, we have used the Newton method. The amount of remnant magnetization can be calculated by the following equation: (15) Fig. 4 shows the pulse magnetizer coil and a rare-earth permanent magnet. In the finite-element analysis, we assumed that the magnetization voltage was 1500 V, the capacitance was 3000 F, and the conductivity of the permanent magnet was S/m, respectively. The easy axis of the permanent magnet was assumed to be parallel to the -axis. The anisotropy field was A/m. Fig. 5 shows the initial magnetization curve in the easy and hard axis. In verification of this method, we compared the calculated results and the measured ones" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003735_2011-01-2652-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003735_2011-01-2652-Figure4-1.png", "caption": "Figure 4. AFADS Overall Layout", "texts": [ " A number of technology enhancements were incorporated, and the robot reach was extended by an additional 10 feet to give it the capability of depainting any of the fighter aircraft in the USAF fleet. AFADS is comprised of two identical robots (one positioned on either side of the aircraft), an aircraft positioning system, an operator control room, three levels of operator control, and an interface with the coating removal process. The media to be used in the process is MIL-P-85891A Type VII cornstarch-based media. The following sections provide a thorough description of the system. The AFADS system incorporates two identical robots as shown in Figure 4. Each robot carries an on-board media blast pot that is refilled automatically by a pneumatic bulk conveying system. The robot controller, manufactured by Kuka Robotics, is housed in a large electrical enclosure that also rides with the robot. Table 1 provides the general specification for each robot. SwRI determined that nine degrees of freedom are required for full accessibility of the robots around a typical fighter aircraft. The first two degrees of freedom (Axes 1 and 2) are used for gross positioning of the robot's vertical column; with the remaining seven axes providing coordinated motion of the end-effector around the aircraft" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002234_978-1-84628-469-4_16-Figure16.2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002234_978-1-84628-469-4_16-Figure16.2-1.png", "caption": "Figure 16.2. Formation flight in inertial frame", "texts": [ " In this chapter, two-module fuzzy logic control algorithms capable of maintaining desirable separation of wingman with the leading UAV are developed. The proposed method of trajectory tracking doesn\u2019t require the communication link, and compared with most existing methods, this formation control scheme is simpler in structure and easier to expand for implementation for other dynamic models. The formation geometry is determined by the relative position between the Leader and Wingman as shown in Figure 16.2. The formation control objective is to steer the Wingman (follower) to maintain a certain separation distance in longitudinal, lateral and vertical directions. By properly defining the body frame and inertia frame, the three equations given below describing dynamic behavior of wingman aircraft can be established [1]. The definitions and the subscripts of the variables used in this chapter are listed in Tables 16.1 and 16.2 respectively. It is assumed that each wingman aircraft could access the leader\u2019s complete flight condition at any time, and they are equipped with the same autopilot system which could execute the same command simultaneously: 1 1 1 (", "5), give the wingman\u2019s complete kinematical equations as follows: [ cos(w wc l l w y y )] wx V V (16.6) sin( )w l l w wc xy V x (16.7) where andwcV wc are the control laws of wingman\u2019s velocity tracking and heading angle tracking. w One of the important tasks in close formation is to separate the wingman away from the leading UAV with certain distance in lateral (x-axis), longitudinal (y-axis) and vertical (h-axis) directions to prevent possible collision. To address this issue, a relative frame as shown in Figure 16.2 is considered. The relative coordinates are defined by r lx x x , r ly y yw , r lh h hw . The formation control problem can be stated as follows. Design control algorithms to maintain the wingman\u2019s relative position ( , ,r r rx y h ) at the desired value with respect to the leader aircraft. Namely, the heading velocity control u , heading angle control u v , and altitude control u are to be designed so thath r l w dx x x x (16.8) r l wy y y yd (16.9) r l w dh h h h (16.10) where ,d dx y and are the desired formation distance in dh ,x y and coordinates" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000200_978-1-4757-3758-5_4-Figure4.7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000200_978-1-4757-3758-5_4-Figure4.7-1.png", "caption": "Figure 4.7. Schematic of sensor constraints for a group of three robots.", "texts": [ " Sensing Constraints. The sensor and communication constraints limit the observations that are possible. We must restrict the choice of control graphs to those that are compatible with the sensor constraints (e.g., a robot cannot follow a robot that it cannot see even if it is within communication range). We adopt a generic sensor model (omni-directional sensing [2] being a special case). The sensor4 has a limited range r s and a limited angular field of view parameterized by the angles a and f3 as shown in Figure 4.7. Notice that Rk detect Ri but not Rj. 4.2. Measures of Performance Measure of safety. In order to prevent collisions, we want to ensure that the separation Cij between robots Ri and Rj is above a threshold (see Figure 4.7). In addition, we will consider the rate of change of this separation and ensure that relative motion between the robots do not cause this separation to decrease below the threshold rapidly. Consider first the dynamics of the formation, where the group configuration is written as a! = [a! 1 a!2 ... a! N V with a!i as defined in Section 3, i E [1, N]. Then each robot Rj with control inputs U j has dynamics given by Egn. (7) which can be written as ;Vj = f(a!j)uj, (14) Suppose Rj has to maintain the separation constraint C ij = c( a" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001537_iicpe.2006.4685350-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001537_iicpe.2006.4685350-Figure8-1.png", "caption": "Fig. 8 shows the phase lag \u03bbsl of estimated by the LPF and the phase lag of \u03bbs estimated by the pure integrator. The phase lag", "texts": [ " There still remains the drift problem due to the very large time constant of the LPF. For the exact estimation of the stator flux, the phase lag and the gain of \u03bbsl in (12) have to be 90 and1 e\u03c9 , respectively. In addition, to solve the drift problem, the pole should be located far from the origin. In this algorithm, the decrement in the gain of the LPF is compensated by multiplying a gain compensator (G), in (15) and the phase lag is compensated by multiplying a phase compensator (P), in (16). The new integrator with the gain and the phase compensator can be given as (17). Figure.8 Vector Diagram of Programmable Low Pass Filter and Pure Integrator ( ) ( ) 2 2 ( ) (1 5 ) ( ) e x p (-j ) (1 6 )1 2 2 1 e x p (-j ) 1 s l ae G a in c o m p e n sa to r c e P h a se C o m p e n sa to r p ae v s ae e \u03c9 \u03c9 \u03c6 \u03c9 \u03c6 \u03c9 \u03bb + = = + = + \u239b \u239e \u239c \u239f \u239c \u239f \u239d \u23a0 (1 7 ) When the programmable LPF is transformed into the sampleddata model using the difference approximation, the sampleddata model has a modeling error, which, in turn, produces an error in the stator flux estimation. This error is more severe when the motor frequency is lower than the cut-off frequency of the LPF" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003061_vppc.2008.4677496-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003061_vppc.2008.4677496-Figure8-1.png", "caption": "Fig. 8 The meshed 3D model (without the stator windings) of the 165W, 4-pole BLDC motor", "texts": [ " The average input power increase and the computed core loss in the last period (80ms ~ 100ms) are compared with the measured data in Table I. The second application is for the no-load core loss computation of a 165W, 4-pole interior permanent magnet (IPM) brushless DC (BLDC) motor. At the ideal no-load operation, the stator currents are 0. In order to simplify the 3D model, the stator windings are removed from the stator core. According to the periodic condition, only one pole (90 degrees) is required for FEA. The meshed 3D model is shown in Fig. 8. At the no-load operation without considering the core loss effects, the electromagnetic torque of the machine has only the cogging torque component. The average value of the cogging torque is zero. When the core loss effects are taken into account, there exists an additional torque component due to the core loss. The waveform of this additional torque component can be derived by subtracting the torque waveform without the core loss effects from that with core loss effects, and is shown in Fig. 9, where negative value means that the torque component is in the opposite direction of the speed" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002490_s12010-007-8088-9-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002490_s12010-007-8088-9-Figure2-1.png", "caption": "Fig. 2 Surface response plot of encapsulated CRL hydrolysis of olive (a), canola (b), and soybean oil (b) showing the POH dependence on pH, temperature, and lipase loading", "texts": [ " For the olive and soybean oils, it was verified that with encapsulated lipase, the pH was the only variable statistically significant (Tables 5 and 6). The analysis of variance (Tables 4, 5, and 6) showed that the statistical significance of the responses for the percentage of hydrolysis is appropriate because a high determination coefficient R2=0.98229, 0.99926, and 0.98905 was obtained for olive, canola, and soybean oil, respectively. Statistical analyses showed significant effects for pH and demonstrate high statistical significance (p<0.05) at 95% confidence level. Tables 3, 4, 5, and 6 and Fig. 2 indicate a region of higher POH where the temperature is around 37 \u00b0C, and there is maximum loading of the entrapped lipase for the tested oils. These results were experimentally confirmed, and the POH was determined for different vegetable oils using an experimental design. Similar values have been found for fatty acids as previously obtained by our group in other studies [6, 17]. This study demonstrated that the statistical analysis is an efficient tool to unfold the influences of pH, temperature, and lipase loading on the POH", " The results of the experimental design for the encapsulated lipase were also analyzed by the RSM using the software STATISTICA\u00ae to find the region where a high POH for canola, soybean, and olive oils can be obtained as a function of pH, temperature, and lipase loading. A high percentage of hydrolysis is important for the biotechnological production processes of biodiesel and fatty acids. Analyzing the curvatures in Tables 4, 5, and 6, one concludes that the POH profile approximates to the optima region for the tested oils utilizing entrapped enzyme. RSM for entrapped lipase (Fig. 2) shows that the typical POH profiles is different from that of free lipase (Fig. 1), while the pH effect is very significant for the entrapped lipase (Fig. 2) with all tested oils; for the free enzyme, the POH is more affected by pH only for olive oil. Maximum hydrolysis was observed at lower pH for the entrapped lipase, whereas for the free enzyme, the maximum hydrolysis occurred at pH 7, for canola and soybean oils. POH was generally smaller for lower loadings of entrapped enzyme, as shown in Tables 3, 4, 5, and 6. This could be due to the limitation of substrate diffusion toward the biocatalyst surface and into the pores of the support because of its microporous structure" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000559_iros.2005.1545275-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000559_iros.2005.1545275-Figure4-1.png", "caption": "Fig. 4. The lower bound on u\u2217 2 is given by \u03b2, the projection of the ellipse onto the u3 axis.", "texts": [ " Also note that this property is local, and does not say anything about the possible evolution of the target position that may tend to make (\u03b8\u2212\u03c6) converge to \u00b1\u03c0 2 . As can be seen from the constraint given in (8), as the difference \u03b8 \u2212 \u03c6 approaches zero, the necessary value for u\u2217 1 approaches its minimum. As a consequence, we adopt an observer strategy that attempts to minimize | \u03b8\u2212\u03c6 |. This can be accomplished by setting u2 = u3. Using an analysis analogous to that used to derive \u03b1 as a lower bound for u\u2217 1, we derive \u03b2 as a lower bound on u\u2217 2. In particular, as shown in figure 4, we project the ellipse f(u1, u3) = 1 onto the u3 axis (since we have set u2 = u3), and after manipulations similar to those above we obtain \u03b2 = 1 l cos(\u03b8 \u2212 \u03c6) \u2264 u\u2217 2 (9) Proposition I If the constraints given by (8) and (9) hold at time t = 0, then in the absence of obstacles, the strategy u2 = u3 will guarantee that surveillance is maintained for all t. To maintain surveillance, it is necessary that the line segment connecting the observer and target (the rod) not intersect any obstacle in the environment (this would result in occlusion of the target)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003001_tasc.2008.922241-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003001_tasc.2008.922241-Figure2-1.png", "caption": "Fig. 2. The Cross-section of dipole Halbach magnet system with 64 sub-magnets.", "texts": [ " It means that the Halbach magnet system shows no dipole moment. In fact, because the distribution of the B field in the sub-magnets is varying with the changing of , the operating point of each of the sub-magnets is changed accordingly. The magnetization is circled by magnets. In the sub-magnets at the poles, operating points are lower; in the middle of a circle between two poles, the operating points are higher, which is shown in Fig. 1(a) and Fig. 1(b). 1051-8223/$25.00 \u00a9 2008 IEEE Taken dipole Halbach magnet system for example, Fig. 2 shows a 2D cross section of a dipole Halbach magnet system. According to the symmetry analysed above, the sum of the magnetic moment in Y direction is zero. Considering magnet 1\u201316, when , has a reverse direction to , , which makes magnet 1\u201316 have magnet moment in -X direction. Similarly magnet 17\u201332, 33\u201348, 49\u201364 have the same magnet moment in -X direction. So the dipole Halbach magnet system does have a dipole moment in -X direction. Equation (4) provides a method to measure the magnetic moment" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001040_0959651041568579-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001040_0959651041568579-Figure4-1.png", "caption": "Fig. 4 Schematic showing variables of the model applied to the drill bit with respect to the stapes footplate", "texts": [ " The dividing lines between these are: dxh dt = dx0 dt \u2212 d dtAFxk B (3) (a) the start of drilling; (b) the start of drill bit cutting across the full diameter Furthermore, the variables h1 and h2 give the angles of the drill bit; subtended from the axial centre of the burr tip on the (c) the start of equilibrium drilling; drill axis to the intersection of the orthogonal plane of (d) the start of breakthrough; the tissue interfaces on each side of the stapes footplate, (e) the completion of breakthrough. as shown in Fig. 4. An equivalent analysis for constant feed force ratherDepending on the compliance, the drilled material thickthan constant feed velocity is given byness and the cutting conditions, (b) and (d) may occur in the reverse order and (c) might not occur at all. dxh dt =F xA2vcpR B 1 2(h2\u2212h1)+sin 2h2\u2212sin 2h1 (4)The start of breakthrough and its identification are key to controlling the drilling process. The start of breakthrough determines the location of the far surface of the The drill torque equation (2) and the relationship given by equation (3) also apply" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002384_memsys.2009.4805354-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002384_memsys.2009.4805354-Figure3-1.png", "caption": "Figure 3: Chip assembly process. (a) Fixing the electrodelglass chip on the PCB by AB glue. (b) Wire-bonding and applying AB glue on the wires for protection. (c) Placing the PDMS lump on the working and reference electrodes, respectively. (d) Bonding the PDMS microchannel by oxygen plasma. (e) Connecting the tubes to the microchannel and then welding ICs on the PCB.", "texts": [ " The inlet/outlet section in the microchannel has a large circular area in which the interconnection can be easily implemented. To make a compact package, the interconnection passage in the PDMS is drilled in the form of an \"L\" shape. Biochip Integration The process for fabricating the implantable biosensor system should be carefully designed to resolve the incompatibility ofthe biomaterials and to achieve the system miniaturization. The electrode/glass chip is fixed on the printed circuit board (PCB) by AB glue (Figure 3(a)). The electrodes on the glass substrate are connected to the corresponding pads on the PCB by applying wire-bonding. The wires are then covered by the AB glue which serves as the protection layer in the following process steps (Figure 3(b)). The wire-bonding process should be carried out prior to bonding the PDMS microchannel on the glass substrate. Otherwise, the PDMS microchannel would hinder the view for aligning the electrodes with the bonding pads in the wire-bonding process. The AB glue must not cover the electrodes on the electrode/glass chip for the success of bonding the PDMS microchannel. Before bonding the PDMS microchannel, the Ag/AgCl reference electrode is chloridized. The enzyme polymerization on the Au working electrode should also be carried out prior to bonding the PDMS microchannel because it is difficult to proceed in the microchannel", " The polypyrrole (from Merck) and GOD mixture is electrochemically polymerized on the working electrode by controlling the voltage from OV to 1.2V in a cyclic voltammetry. After the GOD polymerization, the working electrode becomes dark brown. To bond the PDMS microchannel onto the glass substrate, the bonding surfaces should be modified by using oxygen plasma. It should be noted that the working electrode and the Ag/AgCl reference electrode should be covered for avoiding direct contact with the oxygen plasma (Figure 3(c)). Otherwise, the enzyme on the working electrode will be damaged, and the Ag/AgCl reference electrode will be oxidized. The microchannel which is bonded on the glass substrate is illustrated in Figure 3(d). Finally, the tubes are connected with PDMS microchannel. The interconnection is sealed by AB glue (Figure 3(e)). The polyethylene tube has 0.86mm inner diameter and 1.27mm outer diameter. In order to make the compact interconnection, the \"L\" shape passage of 1.25mm diameter is used. The diameter of the passage is slightly smaller than the outer diameter of the tube. The elastic PDMS passage can tightly clamp the tube and avert the AB glue from permeating into the microchannel by capillary force. After the microchannel and the electrode/glass substrate are assembled, the DEP control ICs (TPS2O010 from Texas Instruments and 74ACT04 from Fairchild Semiconductor) and electrochemistry control ICs (0PA4336 from Texas Instruments) are welded on the PCB (Figure 4)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002926_pime_proc_1970_185_113_02-Figure20-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002926_pime_proc_1970_185_113_02-Figure20-1.png", "caption": "Fig. 20. Analysis of power losses on the road", "texts": [ " These values are obtained at machine test equilibrium temperatures. However, on the road tyre temperatures are much lower due to windage, and using graphs of the form of Fig. 14, rolling resistance and effective rolling resistance values can be corrected to road temperatures. When converted to horsepower loss and plotted against speed the power losses of tyres can be seen in their true perspective Proc lnstn Mech Engrs 1970-71 Vol185 74/71 at UNIV OF VIRGINIA on June 5, 2016pme.sagepub.comDownloaded from 1012 J. A. TURLEY (Fig. 20). Here curve 1 shows the power required in the propeller shaft to maintain the vehicle at a constant speed. Curve 2 is the total loss due to four tyres (two rolling, two driven). The area between curves 2 and 3 is the extra loss due to the increase of tyre rolling resistance, with power transmission through the two driven tyres. The difference between the top two curves is a measure of the aerodynamic drag plus the axle differential power loss. \u2018Chunking\u2019 \u2018Chunking\u2019 is the overall term for the progression of tread damage caused by the combination of speed and torque on tyres over a period of time" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003411_s11771-012-0984-7-Figure9-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003411_s11771-012-0984-7-Figure9-1.png", "caption": "Fig. 9 Photograph of spindle with noise measurement equipment", "texts": [ " In addition, it is considered that if the density of the eccentric mass increases and the volume decreases, vibrations will be reduced. However, there are no problems in these vibration increases because the reference proposed at industrial fields is about 1.2 mm/s for about 5 000 r/min. Sound level meter (NL-31) by RION that presents a level range of 30\u2212120 dB and a frequency range of 20\u221212 500 Hz was used to measure the noise. The height of the measurement points was 1.2 m from the bottom, and the distance of the measurement points was 1 m from the front and side of the spindle. Figure 9 shows the sound level meter used in the J. Cent. South Univ. (2012) 19: 150\u2212154 153 measurement. The noise was measured according to the installation of the automatic variable preload device. Table 1 gives the noise for each rotation speed according to the specific measurement position. Figure 10 shows the comparison of the average value measured at the front and side of the spindle. As the automatic variable preload device was not installed at the maximum speed of 5 000 r/min, the noise levels at the front and side of the spindle were 73" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003789_ijtc2011-61146-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003789_ijtc2011-61146-Figure4-1.png", "caption": "Figure 4: Optical image of assembled foil bearing with Design II elastic foundation", "texts": [ "org/about-asme/terms-of-use 2 Copyright \u00a9 2010 by ASME 18m off-centered bearing sleeves were manufactured through a precision electro-discharge machining (EDM), and 50 m thick Inconel top foil was manufactured through coldforming and age-hardening. The three elastic foundations and top foils are assembled to the slots formed on the bearing sleeve resulting in one sub-foil bearing. Similar to Design I, two sub-foil bearings arranged as back-to-back configuration with 180 degree offset each other completes one foil bearing which generates self-generated preload of 18 m in vertical direction. Figure 4 shows optical microscope images of Design II elastic foundation made through X-ray lithography and electroplating. A test rig was designed with a brushless permanent magnet (PM) DC motor drive [7] developed by Power Electronic Systems Laboratory, ETH Zurich, Switzerland. The original PM rotor with two ball bearings had the first critical speed of 198,840 rpm according to [8] (including an axial turbine). To implement both radial foil bearings and thrust air bearings, rotor was redesigned (Figure 5(b)) with a thrust disc with 12 mm diameter with higher bending critical speed above 1,000,000rpm" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002988_icsps.2009.53-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002988_icsps.2009.53-Figure3-1.png", "caption": "Figure 3. (a) Basic frequencies of a rolling bearing. (b) Structure of a ball bearing and its features.", "texts": [ " However, this method, most times, cannot give any information about where the fault is, but only can answer if there is a fault Therefore, the FFT analysis is used in addition to the previous method, because can give scientists useful information of where the fault may be. There are five characteristic frequencies that can be excited when there is a fault in a ball bearing. The shaft rotational frequency (Fs), the fundamental cage frequency (Fc), the ball pass inner raceway frequency (FBPI), the ball pass outer raceway frequency (FBPO) and the ball rotational frequency (FB). These frequencies appear in Fig. 3a. [4] The theoretical values of ball bearings critical frequencies can be calculated from the equations (7) - (10). [3] The ball pass outer raceway frequency FBP\u039f = (NB / 2) *Fs*(1 \u2013 Db *cos\u03b8 / Dc) (7) The ball pass inner raceway frequency FBPI = (NB / 2)*Fs*(1 + Db *cos\u03b8 / Dc) (8) The fundamental cage frequency FC=1/2*{Fi*(1-Db*cos\u03b8/Dc)+Fo*(1+Db*cos\u03b8/Dc)} (9) The ball rotational frequency FB = (Dc / 2 * Db)* Fs * (1-Db 2 *cos2\u03b8/Dc 2) (10) Where, Db is the ball diameter, Dc is the bearing cage diameter, measured from one ball center to the opposite ball center and \u03b8 is the contact angle of the bearing, as illustrated in Fig.3b. Fi and Fo are the rotational frequencies of the inner and outer race respectively [5][6]. III.MEASURING SYSTEM In our measurement system there were used two accelerometers, one Data Acquisition (DAQ) card, one load cell and one press (to load the bearing using hydraulic pressure). The appropriate code developed using the Matlab software. The ball bearing 6200 2Z was used. The inner race of the bearing was held still, and the outer race was rotating. The measurement system is shown in Fig.4. A power supply needed to give the power to load cell" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002163_j.mechmachtheory.2007.03.005-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002163_j.mechmachtheory.2007.03.005-Figure2-1.png", "caption": "Fig. 2. The LLC of intersecting homologous lines described in [3].", "texts": [ " The vector connecting every pair of homologous points in space belongs to a helicoidal vector field of pitch \u00f0d=2\u00de= tan\u00f0/=2\u00de [4], and its corresponding LLC is usually referred to as the bisecting LLC. It has been shown that the bisecting LLC degenerates into the LLC of instantaneous kinematics; nevertheless, the LLC associated with homologous lines has been found to be different from the other two LLCs [3], and it does not take the form of Fig. 1. The LLC of homologous lines was introduced by Bottema and Roth [3] when investigating the two-position theory in kinematics. As shown in Fig. 2, given two-positions of a rigid body, one can determine the screw axis of the displacement and the corresponding rotation and translation parameters, / and d, respectively. For any point R in space, there exists a pair of homologous lines, L1 and L2, intersecting at R. The plane formed by L1 and L2 is denoted by a, whose normal at R is denoted by line N. For all the13 points in space, there exist13 such normal lines, which form a linear line complex [3]. We refer to it as the LLC of (intersecting) homologous lines" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000702_tro.2005.844677-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000702_tro.2005.844677-Figure1-1.png", "caption": "Fig. 1. Redundant set of triaxial accelerometersA attached to the rigid-body B located at points P , for i = 1; . . . ;m > 3.", "texts": [ " It appears, however, that the accelerometer orientation changes induced by the system flexural motion has been neglected in all of [11]\u2013[13]; this may result in erroneous estimation results. The use of accelerometers, either alone or along with strain gauges, to instrument structurally flexible manipulators has been reported extensively, e.g., in [14]\u2013[17]. However, in all these works the accelerometer signals are employed for acceleration feedback, not for estimating the flexural states. Now, let us assume that an accelerometer array\u2014i.e., a kinematically redundant set of more than three triaxial accelerometers\u2014as depicted in Fig. 1, is attached to the base of the micromanipulator of a macro\u2013micro structure, and that the array data are processed to obtain the twist\u2014velocity and angular velocity\u2014of the base. We report here on an algorithm which utilizes the accelerometer readouts and the twist data of the base to estimate the flexural states of the macromanipulator in an extended Kalman filter (EKF). It should be noted that the relations between the flexural states and the twist, which are the state\u2013output relations here, are nonlinear" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002978_ccdc.2009.5191607-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002978_ccdc.2009.5191607-Figure1-1.png", "caption": "Figure 1: The Single Inverted Pendulum", "texts": [ " Then we state some preliminaries of the energy well controller and SDRE controller in Section 3 and Section 4. The simulations and results are shown in Section 5. Conclusions and future scope of study are given in Section 6. 2 SYSTEM DYNAMICS The single inverted pendulum consists of one linked pendulum on a wheeled cart that can move linearly along a horizontal track and a force F to move the cart in order to balance the pendulum on the cart. The model of the single inverted pendulum is illustrated by Figure 1,where x refers to the displacement of the cart, \u03b8 is the angle of the pendulum from the vertical direction, m0 is the mass of the cart, m1 is the mass of the pendulum, l denotes the distance between the pivot and the center of mass of the pendulum and J is the moment of inertia with respect to the pivot point. The equations of motion can be obtained by the Euler-Lagrange formulation. The form of the EulerLagrange formulation is given as follows d dt ( \u2202L \u2202q\u0307i ) \u2212 \u2202L \u2202qi = Qi (i = 1, 2), (1) where L = T \u2212 U and T is the total kinetic energy of the system, U is the total potential energy of the system, q is the generalized coordinates of the system (q [x \u03b8]T ) and Qi is the generalized forced not taken into account in T , U " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000901_030932405x15936-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000901_030932405x15936-Figure1-1.png", "caption": "Fig. 1 Overlapping triangular distribution of n equal elements of pressure; each element is of halfwidth b [1]", "texts": [ " The elastic displacements of corresponding points on the two surfaces then satisfy the relationship uz1 uz2 \u00fe h\u00f0x\u00de \u00bc \u00bc 0 within the contact > 0 outside the contact \u00f01a\u00de \u00f01b\u00de ( where uz1 is the displacement for the rough half-plane, uz2 is the function defining the displacements of the smooth half-plane and is the approach of distant reference points in the two bodies. In the method used, it is first necessary to find the compliance matrix of the system. The nominal contact area is divided into n equal elements and, in this window, a continuous piecewise linear distribution of pressure is chosen, which produces surface displacements everywhere smooth and continuous. Such a distribution of pressure can be built up by the superposition of overlapping triangular traction elements, as shown in Fig. 1 [1]. The matrix of influence coefficients Cij is required; they express the displacement at a general point i due to a unit pressure element centred at point j. The total displacement uzi, at position i, is then expressed by uzi \u00bc \u00f01 2\u00deb E X j Cij pj \u00f02\u00de where b is the width of the single element, pj is the pressure on the jth node, and E and are Young\u2019s modulus and Poisson\u2019s ratio respectively for the material. Difficulties arise in line contact in evaluating the coefficients Cij because the displacements are defined except for an arbitrary constant (the rigid body displacement at infinity is unbounded because of the logarithmic nature of the Green\u2019s function)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002490_s12010-007-8088-9-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002490_s12010-007-8088-9-Figure1-1.png", "caption": "Fig. 1 Surface response plot of free CRL hydrolysis of olive (a), canola (b), and soybean oil (c) showing the POH dependence on pH, temperature, and substrate concentration", "texts": [ " For free lipase, at 70% oil\u2013water ratio (not shown at Table 2), all oil suspensions have shown a percentage of hydrolysis (POH%) lower than that observed at 50% oil\u2013water ratio, suggesting substrate mass transfer limitations at the highest oil concentrations. Analysis of variance (Tables 4, 5, and 6) shows that the statistical significance for the responses of the percentage of hydrolysis is appropriate because a high determination coefficient (R2) of 0.98542, 0.90664, and 0.95227 was obtained for olive, canola, and soybean, respectively. In this part of the study, RSM was used as an approach for determining the region where the percentage of hydrolysis is maximized for the oils tested (Fig. 1), confirming that in the ranges studied, temperature (30\u201344 \u00b0C) is at its optimum and pH (4 to 10) is optimum in the lower range of the 4 to 7 interval. According to Fu et al. [16], a complete hydrolysis can be achieved by either lengthening the reaction time at low free enzyme concentration or increasing the enzyme concentration for a shorter reaction time, to achieve 90\u201398% hydrolysis with coconut oil and other oils, and lipase from Aspergillus sp. The former is preferable for industrial production when using an expensive enzyme", " The results of the experimental design for the encapsulated lipase were also analyzed by the RSM using the software STATISTICA\u00ae to find the region where a high POH for canola, soybean, and olive oils can be obtained as a function of pH, temperature, and lipase loading. A high percentage of hydrolysis is important for the biotechnological production processes of biodiesel and fatty acids. Analyzing the curvatures in Tables 4, 5, and 6, one concludes that the POH profile approximates to the optima region for the tested oils utilizing entrapped enzyme. RSM for entrapped lipase (Fig. 2) shows that the typical POH profiles is different from that of free lipase (Fig. 1), while the pH effect is very significant for the entrapped lipase (Fig. 2) with all tested oils; for the free enzyme, the POH is more affected by pH only for olive oil. Maximum hydrolysis was observed at lower pH for the entrapped lipase, whereas for the free enzyme, the maximum hydrolysis occurred at pH 7, for canola and soybean oils. POH was generally smaller for lower loadings of entrapped enzyme, as shown in Tables 3, 4, 5, and 6. This could be due to the limitation of substrate diffusion toward the biocatalyst surface and into the pores of the support because of its microporous structure" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002993_vppc.2009.5289840-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002993_vppc.2009.5289840-Figure1-1.png", "caption": "Figure 1. Vector diagram of IPMSM in traction and braking stages", "texts": [ " When the electric drive system changes from traction stage to regenerative braking stage, the output torque order of the speed regulator is a negative value braking torque, correspondingly negative the input order of the current regulator which generates the relevant control voltage vectors, while the back-EMF still remains positive, thus the motor is in the state of regenerative braking and the capacity and battery of the DC link are charged by the feedback electric energy, which is illustrated in Fig. 1. There are three usual regenerative braking modes in the electric drive system: maximum feedback power braking, maximum feedback efficiency braking and constant torque braking [4]. The braking current of the first mode is heavy, the braking distance is short, but the maximum efficiency of the feedback energy is less than 50%; the feedback efficiency of the second mode is highest, and the braking current is low yet the distance is long; the performance of constant torque braking mode falls in between" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003065_tec.2008.2008881-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003065_tec.2008.2008881-Figure6-1.png", "caption": "Fig. 6. Installation of the detector. (1) Detector. (2) Slot wedge. (3) Batten or ripple spring. (4) Top conductor bar. (5) Bottom conductor bar.", "texts": [ " The tapping pole would vibrate vertically depending on the wave type, frequency, and amplitude of the input signal given by the control system. Since the vibration of the stator slot wedge is usually featured with very small amplitude, short-time action, and high frequency, the microphone and the displacement sensor should be qualified with not only high sensitivity and resolution but also fast response. In addition, the two sensors should not be installed on the surface of the tested slot wedge but be mobile with the detector. Fig. 6 shows how the detector is installed in HV generators for work. With the two sides of magnets on the iron, the detector is placed over the slot tightly, while the tapping pole and two sensors are just located on the slot wedge and prepared to test. The control system mainly consists of a signal conditioning board, a DSP motherboard, an LCD display screen, and a touch kit (see Fig. 7). The signal conditioning board has three important missions. First, it offers working power supply for the two sensors through two sets of different circuits" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000226_eurbot.1999.827617-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000226_eurbot.1999.827617-Figure1-1.png", "caption": "Figure 1: Stewart Platform configuration", "texts": [], "surrounding_texts": [ "This paper presents a solution for both compensation of accelerations and precise and powerful manipulation. In order t o achieve a compensation a Stewart Platform is mounted on t o p of a mobile robot.\nUsually a Stewart Platform is used t o generate accelerations. In this application the acceleration vector of the robot is inverted and sent to the filter, which performs the platform motion depending on the robot\u2019s movement. This filter handles all s ix degrees of freedom, so every movement can be taken into account, if it is determined b y acceleration or inclination sensors.\nFurthermore this combination is also a very precise docking system, since the Stewart Platform can easily correct the uncertainty in the robot\u2019s position.\n1 Introduction Mobile service-robots are mainly used for flexible transportation tasks. Their sophisticated navigation systems let them move in complex environments on trajectories which are different on every run. Therefore it is not possible to predict the movement and the accelerations which affect the payload of the robot.\nThere are passive systems to balance the payload, but they are all based on a flexible suspension and a low center-of-gravity. These systems can not be adapted to the characteristics of the payload.\nTherefore an active system would be necessary, that could do more than only compensate accelerations, e.g. perform precise docking.\nStewart Platforms are mair,,j used for fligA or driving simulators and become more and more important for machine tools as a more flexible instrument mounting point. The advantages of Stewart Platforms are their excellent relationship between weight and payload capabilities and their precise positioning. Their small workspace is not a disadvantage in this application.\nIn simulators Stewart Platforms are used to transfer accelerations to the person on the platform. Thus a Stewart Platform can also generate an antiacceleration, if the base below the platform is moving.\nThe solution presented in this paper deals with the kinematics and the dynamics of the combination of Stewart Platform and mobile robot. In addition to this the filter for motion generation is discussed.\n2 State of the art Mobile service-robots have to perform several tasks, mainly transporting objects in a more flexible way than driverless transport systems do. If the robot moves between people in narrow passages, its trajectory will not be a straight line, but a very complex curve with lots of accelerations and decelerations.\nDepending on the payload these accelerations should be small or even zero. In a medical application, where the payload is an injured person, a very smooth but also fast movement of the robot is aspired.\n2.1 Compensation of accelerations The first attempt to compensate the accelerations of a moving robot was done by Terashima et. al. and is called sloshing control [TS94]. They measured the fillings of an open tank with sensors to determine the movement of the surface of the liquid inside.\nWith this information and a model of the liquid they interrupted the acceleration of the vehicle to decrease\n0-7803-5672-1/99/$10 0 1999 IEEE", "the movement of the surface. If the movement turns to the opposite direction, the vehicle is accelerated again, until the maximum speed is reached. In addition to the behaviour of the liquid different shapes of the tank were investigated like rectangulars [THS94] and cylinders [THS95].\nObviously this is a really inharmonic movement and not practical for the transport of sensible objects or even persons. Furthermore, this method does not compensate for any accelerations, but only reduces their consequences.\nThe first attempt of an active compensation was done by Fukuda et. al. [FSSSO]. But they also did not compensate for the accelerations but for the sloshing of the surface.\n2.2 Stewart Platforms The Stewart Platform is a six degree of freedom parallel manipulator [Ste65]. Its advantage is to handle payloads up to its own weight and to position them very precise. In contrast to a sequential manipulator, Stewart Platforms have no singularities in their workspace.\nThe inverse kinematics of a Stewart Platform can be calculated easily. Let (AI , . . . , As) be the lower mount points of the joints and (B1, . . . , B6) the upper ones in their respective coordinate system. A is the coordinate system of the lower platform and B that of the upper one.\nThe servos' lengths SI, . . . , sg result from the amount of the distances sk = I A k B k 1. If r describes the position and the 3x3 Matrix R describes the orientation between B and A, the lengths of the servos can be calculated by 3\nsk=Ir+Rbk-akI w i t h H = l , . . . , 6 .\n2.3 Combination of both systems Until now no combination of Stewart Platform and mobile robot exists. The idea was proposed the first time in [GD97]. There are several combinations of a mobile robot and a manipulator in order to increase the manipulator's workspace or to improve its configuration.\nThe possible applications of this combination are not only the compensation of accelerations; but also the positioning and manipulation of the payload.\n3 Combination of Stewart Platform and a mobile robot\nThis section describes the separate systems and afterwards their combination and coordination.\n3.1 PRIAMOS The mobile robot PRIAMOS' is an experimental platform for testing sensors and navigation strategies [DKW92]. Its holonomous driving concept with Mecanum wheels is a suitable base, since all three degrees of freedom (z, y, a) are completely independent. Its movability encloses that of every other diving concept.\nThe mecanum drive is based on four wheels with free rotating barrels mounted at the angle of 45 degrees along each wheel's periphery. Figure 2 shows the placement of the wheels.\nIf the wheels on one side rotate a t the same speed in opposite direction, the robot moves sidewards, as if a worm gear was installed between wheels and ground.\nPRI permutation of IPR, Autonomous Mobile System", "If the diagonally opposite wheels rotate in opposite direction, the robot turns on the spot. According to the four wheel speeds V I , . . . ,214 (dimension 5) the velocity can be described by\n21, = Ct * T w h e e l * ( V I + ~2 + 03 + v4)/4 (1) \"y = ct * r w h e e l * (-211 + U 2 + 213 - V4)/4 ( 2 ) 0, = Cy * T w h e e l * (VI - + 03 - ~ 4 ) / 4 (3)\nwhere Twhee l is the radius of the mecanum wheels and Ct,,. are constant values for the translational and rotatoric movement.\nThe derivatives of equation 1 to 3 deliver the accelerations of all three degrees of freedom\nThese values describe the behaviour of the vehicle and can be used as input for the filter (see section 4). The maximum accelerations of PRIAMOS are amazt,y = 0.5 5 and a m a x , = 1.2 9.\n3.2 SPIKE The Stewart Platform SPIKE2 is a full 6 DOF motion platform (see figure 3) . It is driven by 6 motors and spindle drives. The workspace of the platform is: z, y, z E [-100mm.. .100mrn] and cy,P,y E ( - 2 5 O . . . 2 5 O ] .\nThe gradient of the spindle is s = 2mm. The number of revolutions of the motors is U = 3000 6 = 50 3 . Therefore the velocity of the servos is Vma, = 100 y. This velocity is reached in At = 0.5s, so the maximum\n2Stewart-Platform of the IPR KarlsruhE\nacceleration of the servos is\ni 7 . s At - - a m a x - m mm = 200- = 0.2-\nS2 S 2\n3.3 Combination The combination of PRIAMOS and SPIKE is a highly dynamic system, since both parts have no restrictions in their kinematics. Figure 4 shows a possible application carrying a glass filled with liquid. The coordinate systems will be described in section 4.\nThe mobile robot generates only accelerations in 2,- y, and &-direction. The Stewart Platform can generate accelerations in all six directions. If sensors are used instead of odometric calculations, the motion of the vehicle can also be described by six accelerations.\nThe parameters of the robot position are z, y and cy, the parameters of the platform position are z, y, z , j , p and P with 3 corresponding to cy.\nIf inclination sensors are installed on the vehicle, its movement consists of six parameter also. In order to take the whole movement of the vehicle into ac-" ] }, { "image_filename": "designv11_61_0000510_icems.2005.202511-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000510_icems.2005.202511-Figure1-1.png", "caption": "Fig. 1. Structure and flux linkage of an SPMSM.", "texts": [ " In this model, the saliencies of the motor are reflected by the variation of the stator winding inductances with respect to the rotor positions. In order to obtain the relationship between the inductance and rotor position, an inductance pattern is established based on the experimental measurement of a surface mounted PMSM (SPMSM) and Fourier series. Dynamic perfornance simulation using the derived model is carried out and the results are compared with those based on the model without considering the saliencies. II. NoN-LINEAR MODEL OF A PMSM Fig. 1 illustrates schematically the structure and flux linkage of a SPMSM. When the motor saliencies are taken into account, the conventional model of PMSM is inaccurate since the inductances of the motor are no longer constants and will be the functions of the rotor position and stator current due to the saliencies. Therefore, a new PMSM model is required. A. Flux Linkage ofPMSM In a PMSM, the rotor field is the dominant field which deternines the operating point. Although the characteristic of the magnetic core is non-linear, the magnetic circuit can be considered as piecewisely linearized around the operating point P at a given rotor position, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002574_1.3238251-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002574_1.3238251-Figure1-1.png", "caption": "FIG. 1. a Schematics of the motor; b the simulation box.", "texts": [ "7 However, as the acoustic waves are propagating in the air, the air is compressed and rarefied periodically, which leads to that the temperature at any point changes periodically. Since the dynamic viscosity of the air depends on temperature, it is expected that the dynamic viscosity varies both temporally and spatially. In order to investigate the effect of such changing viscosity on the performance of the motor, the viscosity is taken with a temperature-dependent form, which is described by the Sutherland formula. Furthermore, it is also interesting to check how the heat conduction and the size of the motor can change the performance of the motor. Figure 1 a shows the schematic structure of the acoustic streaming-driven motor, where the outer annulus is the stator with an ultrasonic wave propagating on the inner surface, and the inner annulus is the rotor with the radius of R. Air is filled in the gap between the stator and the rotor, with the thickness of h. As used in Luchini and Charru,6 such an axisymmetric system is simplified to two parallel plates, as depicted in Fig. 1 b , since h is much smaller than the perimeter of the stator. The comparison of the pure shear motions in the Cartesian and polar coordinates indicates that such a simplification introduces a relative error on the order of h / 8R , i.e., 2 10\u22124, in the present scenario.7 Thus, the problem becomes that an ultrasonic transverse wave is propagating to the right on the inner surface of the upper a Electronic mail: chenglp@nju.edu.cn. 0021-8979/2009/106 7 /074506/4/$25.00 \u00a9 2009 American Institute of Physics106, 074506-1 [This article is copyrighted as indicated in the article" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002499_elan.200804525-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002499_elan.200804525-Figure1-1.png", "caption": "Fig. 1. Schematic diagram of the FIA set-up for the determination of copper(II) ions: A) FIA-lab 2500 instrument; B) peristaltic pump; C) flow-injection valve; D) sample loop; E) carrier stream; F) flow-through cell; G) thin-film microsensor; H) Ag/AgCl reference electrode; I) Keithley multimeter/data acquisition instrument; J) computer; K) waste.", "texts": [ " FIA is the most widely used analytical technique under nonequilibrium conditions, in which a definite sample volume is injected to a carrier stream to be detected using a flowthrough cell. The measurements based on FIA reveal better precision, higher analysis throughout and a smaller sample volume, in comparison with batch analysis. The measurement set-up consists of a commercial flow-injection analyzer FIA-lab 2500 in combination with a Keithley multimeter/ data acquisition system connected to a personal computer to monitor the potentiometric response of the thin-film organic/inorganic sensor hybrid (see Fig. 1). The sensor chip is tightly sealed inside a miniaturized home-made flow through cell with a size of about 2 cm 1 cm 0.5 cm and a standard Ag/AgCl reference electrode has been placed into the waste container. Under optimized conditions (pump flow rate 2 0.2 mL/ min and carrier stream of 10 1 mol L 1 KNO3), the organic/ inorganic sensor hybrid has been characterized under dynamic mode of operation by injection of a series of standard copper(II) ion working solutions (10 6 \u2013 10 1 mol L 1). Electroanalysis 2009, 21, No", " Moreover, these values are much smaller than 1.0 indicating that the hybrid microsensor array exhibits excellent selectivity towards copper(II) ions over all the interfering tested cations. Thus, the merit offered by the realization of the new type of hybrid microsensor array is the high selectivity of the organic membrane-based microelectrodes in combination with the high stability of the inorganic chalcogenide glass-based microelectrodes. The latter has been investigated and published by some of the authors in [21]. Figure 1 sketched the schematic diagram of the FIA set-up with the home-made flow-through microcell. Under optimum conditions, the copper ions are detected downstream with the proposed thin-film organic/inorganic sensor hybrid as copper(II) ion FIA detector. The optimized FIA parameters of the home-made set-up and flow-through cell are Table 2. Selectivity coefficients of the Cu2\u00fe-sensitive organic/ inorganic sensor hybrid towards typical interfering ions. Interferent cations Selectivity coefficient Inorganic-type thin-film microsensor Organic-type thin-film microsensor Cu2\u00fe 1 1 Co2\u00fe 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001268_1.2125887-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001268_1.2125887-Figure3-1.png", "caption": "Fig. 3 Photograph of ball-on-flat rig", "texts": [ " Trapping a film of lubricant between a dropping steel ball and a transparent sapphire flat can produce a thin layer of lubricant under high-pressure conditions. Using a sapphire flat as one of the media allows higher loads than glass to be applied to the contact, but still maintains a transparent interface so that optical interferometry can be used to measure the film thickness of the trapped lubricant layer. Figure 2 shows a schematic of the test apparatus employed and a photograph of the actual rig is shown in Fig. 3. It can be seen to consist of a base plate and column, a lever-arm and ball, and a threaded lever-arm holder. The sapphire plate diameter 20 mm, Fig. 2 Schematic diagram rom: http://tribology.asmedigitalcollection.asme.org/ on 08/30/2017 Terms thickness 3 mm is mounted within a recessed hole in the middle of the base plate. The column was fixed to the base plate and, via a fulcrum pin, supported the lever-arm which held the 50 mm diameter steel ball at 21 mm from the pin. A threaded rod acted as a lever-arm holder, which was used to fix and tighten between the lever and the base plate see Fig. 3 . This helped prevent the lubricant escaping from the contact area, and allowed the rig to be easily moved around so that the ultrasonic water-bath test could be performed. The drop height of the ball was controlled by the initial angle of the lever, and weights could be added to the end of the beam to vary the impact force and to help prevent the escape of lubricant from the contact area. Various conditions of lever angles and weights were investigated in relation to the films and entrapped volume formed 25 " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000980_mhs.2004.1421274-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000980_mhs.2004.1421274-Figure8-1.png", "caption": "Figure 8. Learning motion. Figure 9. HOAP1: Humanoid for Open Architecture Platform.", "texts": [ " In successful experiments, the allocation of basis functions with AE-GSBFN differs little from A-GSBFN. Statistics however indicated that AE-GSBFN achieved G 18 times for 20 repetitions, but A-GSBFN achieved only 9 times. It can be seen that AE-GSBFN performs better than A-GSBFN, and we consider that AE-GSBFN avoided a fall into local minima through the above experiments. 4.2 Controlling Humanoid Robot In this section, as learning of continuous high-dimensional state spaces, AE-GSBFN is applied to a humanoid robot learning to stand up from a chair (Figure 8). The learning was simulated using the virtual body of the humanoid robot HOAP1 made by Fujitsu Automation Ltd. Figure 9 shows HOAP1. The robot is 48 centimeters tall, weighs 6 kilograms, has 20 DOFs, and has 4 pressure sensors each on the soles of its feet. Both of sensors of angular rate and acceleration are mounted in its breast. To simulate learning, we used the Open Dynamics Engine [9]. The robot can observe the following vector s(t) as its own state: s(t) = (\u03b8W , \u03b8\u0307W , \u03b8K , \u03b8\u0307K , \u03b8A, \u03b8\u0307A, \u03b8P , \u03b8\u0307P ), (15) where \u03b8W , \u03b8K and \u03b8A are waist, knee, and ankle angles respectively, and \u03b8P is the pitch of its body (see Figure 8). Action uj(t) of the robot is determined as follows: uj(t) = (\u03b8\u0307W , \u03b8\u0307K , \u03b8\u0307A), (16) One trial terminates when the robot fell down or time passed over ttotal = 10 [s]. Rewards r(t) are determined by height y [cm] of the robot\u2019s breast: r(t) = 20\u00d7 | lstand \u2212 y lstand \u2212 ldown | (during trial) 20\u00d7 |ttotal \u2212 t| (on failure) , (17) where lstand = 35 [cm] is the position of the robot\u2019s breast in an upright posture, ldown = 20 [cm] is its center in a falling-down posture. We used umax j = 1 36 \u03c0 [rad], \u03b3 = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000923_12.547996-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000923_12.547996-Figure1-1.png", "caption": "Figure 1. Model of Jeffcott rotor with crack and constant driving torque (a), crack section in inertial and rotating coordinates (b).", "texts": [ " In addition, the behaviors of a stalled rotor\u2019s angular velocity as well as the rotor\u2019s maximum amplitude of unbalance are presented as functions of the crack characteristics. The consequences of torsionally induced slant cracks are not considered in this study. 2. CRACK MODELS AND STIFFNESS MATRIX The theoretical model, called the Jeffcott rotor, employs a flexible rotor composed of a centrally located, unbalanced disk attached to a massless elastic shaft. In turn, the shaft is mounted symmetrically on rigid bearings (see Fig. 1(a)). For this study, the shaft has a transverse crack running across its cross-section located near the disk. The stiffness of the uncracked rotor system is symmetric (isotropic) and the damping due to the air resistance effect is assumed to be viscous. The rotor is driven by a constant external drive torque. The angle between the crack centerline and the line connecting the bearings and shaft center (Fig. 1(b)), ( )arctan y z\u03c8 = \u03a6 \u2212 , is used to determine the closing and opening of the crack. At any instant of time, the \u03be-axis remains perpendicular to the face of the crack. This allows the body-fixed rotating coordinate frame (\u03b6,\u03b7,\u03be) to spin with the same velocity as the rotor. The weight dominance, assumed in almost all previous analyses of horizontal cracked rotors, is not required. Furthermore, the influence of the whirl speed on the closing and opening of the crack is included. This nonlinear system with time-varying stiffness coefficients is studied numerically, with particular focus on the effect of different crack depths on the rotor stalling", " of SPIE Vol. 5393 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 07/17/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx ( ) ( ) ( ) ( ) 2 2 2 cos sin 0sin cos sin cos I p a M Mg M J M M z y T \u03b5 \u03b8 \u03b8 \u03b5 \u03b8 \u03b8 \u03b5 \u03b8 \u03b5 \u03b8 \u03b8 \u03a6 + \u03a6 + + = + \u03a6 \u2212 \u03a6 + + \u2212 + = Mq Cq K q (7.1) The matrices, 0 0 M M = M and 0 0 C C = C , are the rotor mass and damping matrices, respectively; ( ) Tz y=q is the vector of the disk\u2019s displacements; and \u03a6 is the rotor spin angle (see Fig. 1(b)). Using Eq. (7.1) one can obtain the following equation relating torque and the angular displacement: ( ) ( ) ( )( ) ( ) ( )2 1 2 1 2 ( )sin cos sin cos sin 2 sin 2 sin cos cos 2 sin cos 2 cos p a f KJ T Mg K y z C y z k y z k k z k k y \u03b5 \u03c8\u03b5 \u03b8 \u03b5 \u03b8 \u03b8 \u03b5 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03a6 = + + \u2212 + \u2212 + \u00d7 \u2206 \u03a6 \u2212 + \u2206 + \u2206 \u03a6 \u2212 \u2206 \u2212 \u2206 \u03a6 (7.2) where \u03b8 \u03b2= \u03a6 \u2212 , arctan y z \u03c8 = \u03a6 \u2212 , , \u03b8 \u03b8= \u03a6 = \u03a6 , and 2 pJ Mr= . Normalizing displacements with respect to unbalance eccentricity, and introducing nondimensional time, damping, and torque, the equations (7" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003958_09205071.2013.753662-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003958_09205071.2013.753662-Figure1-1.png", "caption": "Figure 1. Bearing elements.", "texts": [ " In spite of its implementation which is relatively expensive and needs a high piezoelectric sensors range, the vibration monitoring technique is the most precise and does not requires an additional artificial intelligence technique to detect the fault. For example, the current monitoring technique which is the simplest one but in some cases especially for small defects does not see the defect signature. In this paper, the vibration diagnosis technique for bearing faults is implemented and validated experimentally. The majority of electric machines use rolling element bearings. Each bearing consists of two races called inner and outer races (show Figure 1) [10]. A set of balls or rolling elements placed in raceways rotate inside these races. Bearing failures are responsible for the highest incidence of recorder electric motors failures as shown in Figure 2 [11,12]. Bearing failure leads to the vibrations of the rotor; these vibrations are transmitted to the stator via the outer race of bearing and are added to the vibrations of different source in the machine [13,14]. In this work and for theoretical calculation of vibration, two types of imperfections, or defects, are considered in the geometry of the bearing components" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002250_978-3-540-92841-6_559-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002250_978-3-540-92841-6_559-Figure2-1.png", "caption": "Fig. 2. Generation Method of Human Model Using Laser Range Finders", "texts": [ " Based on the information of the user, the robot systems could support the walking appropriately. In the following parts of this paper, we introduce the human linkage models for controlling the RT Walker and the Wearable Walking Helper, and experimental results with both systems illustrate the validity of them. II. USER MODEL FOR WALKER-TYPE ASSIST SYSTEM In this section, we explain the generation method of the human model that has seven rigid body links in sagittal plane. In this research, we set up two Laser range finders (LRFs) on the walker as shown in Fig.2(a). One of them is attached to the same height with hip joint of the user with standing state and measures the relative distances between the human surface and the walker along the vertical direction. This laser range finder measures the relative distances from 50[deg] to 30[deg] every 10[deg] under the assumption that the height of the hip joint of the user with _________________________________________ IFMBE Proceedings Vol. 23 ___________________________________________ standing state is 0[deg]", " For generating the human model, firstly, the data of LRFs are collected. Three points on the surface of the lower body and three points on the surface of the upper body are obtained from the upper part LRF. The equations of the straight lines that show the upper-body and the lower half of the body are derived to detect the position of the human hip joint during the motion of the user using walker. These equations are derived based on the least square method. The intersection of these straight lines is calculated as shown in Fig.2(b), and the intersection point is regarded as the position of the hip joint of the human model. The position of the shoulder is calculated by using the detected hip joint position, known length of the upper body link and the inclination of the straight line. From the shoulder position and the hand position, the angles of the arm joints are calculated by solving the inverse kinematic equations as shown in Fig.2(c). The angles of the legs joints are obtained by solving the inverse kinematic equations based on the position of the hip joint and the position of the feet obtained by LRF attached to the lower part of the walker as shown Fig.2(d). If we can estimate the user states during the usage of the walker, we can change the function of the walker appropriately based on the user states. In this paper, we consider the estimation method of the user states based on the human model. For the simplicity of the discussion, in this paper, we consider thirteen states of the user during the usage of the walker as shown in Fig.3. Main states of the user during the usage of the walker are Sitting state, Standing state and Walking state. The user sometimes faces the dangerous situations including the falling accident, which are estimated by the system as Emergency states" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001726_10946978_19-Figure19.3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001726_10946978_19-Figure19.3-1.png", "caption": "Fig. 19.3. Structure of force sensor", "texts": [ " This type of sensor has to be installed under the front end of the U-shaped supporting arm so that all the forces and torques act on the force-sensing beam. This configuration is unfavorable because the vertical load produces a large torque about the horizontal axis on the force sensor in addition to the load itself. This additional torque lowers the sensing precision of the forward force and the turning torque. It is possible to find a high-grade product that avoids these problems, but its cost is high. To solve this dilemma, the authors have developed a new force-sensing system. Figure 19.3 illustrates the structure of the developed force sensor. It is embedded in the U-shaped supporting arm. It comprises upper and lower members connected with four elastic joints at their corners and four gap sensors that detect vertical and horizontal (forward/backward) displacements between the members. The elastic joints are designed so that their vertical stiffness is larger than the horizontal one. The upper member is covered with a soft pad, and the lower member is attached to the walker body" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000278_1-84628-179-2_7-Figure7.6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000278_1-84628-179-2_7-Figure7.6-1.png", "caption": "Fig. 7.6. T-Wing demonstrator vehicle.", "texts": [ "5 kg). Although the petrol vehicle is still very much a concept demonstration platform, this increased take-off weight should allow an endurance of up to several hours carrying a 5 lb payload. The vehicle is built primarily of carbon-fibre and glass-fibre composite materials with local panel stiffness provided by the use of Nomex honeycomb core material. The airframe has been statically tested to a normal load factor in excess of 8 G\u2019s [73]. A picture of the completed T-Wing vehicle is shown in Figure 7.6. During hover, the vehicle is controlled in pitch and roll via elevon control surfaces on the wing which are submerged in the prop-wash of the propellers. Yaw control of the vehicle is effected via fins and rudders attached to the nacelles and which are also submerged in the propeller slipstream. Additionally the tips of the fins provide the attachment point for the landing gear and hence determine the footprint of the vehicle on the ground. From the beginning of the T-Wing concept in mid-1995 it was proposed that the vehicle be allowed to transition from vertical to horizontal flight via a stall-tumble manoeuvre" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000746_12.619143-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000746_12.619143-Figure2-1.png", "caption": "Fig. 2 Schematic of a cemented doublet lens mounted as in Fig. 1.", "texts": [ " For example, consider the design shown schematically in Fig. 1 . Here, a biconvex lens is clamped axially with some nominal PA in a cell between a shoulder and a threaded retainer. The glass-to-metal interfaces appear in the figure to be sharp corners, but tangential or toroidal interfaces would be more appropriate in an actual design. Key mechanical changes that can occur in this design and that contribute to the magnitude of its K3 factor are as indicated in Table 1. Designs, such as those for cemented doublet lenses (see Fig. 2) or lenses separated by spacers (see Fig. 3), would add components that further increase complexity and provide additional elastic variables. Prior discussions of K3 considered only the first two of these contributing factors. 6 Section 2, we summarize the prior theory that defmes a first approximation for K3 resulting from bulk compression of the lens and bulk elongation of the cell wall. Neglecting factors from Table 1 other than the two bulk effects is believed to make the value for K3 for any lens assembly greater than actually would occur in real life" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002607_978-0-387-77747-4_6-Figure6.9-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002607_978-0-387-77747-4_6-Figure6.9-1.png", "caption": "Fig. 6.9 Simple model of an ultra-short hydrostatic gas journal bearing [31]", "texts": [ " Scaling laws for the stiffness and damping coefficients and the natural frequency of hydrostatic micro-gas journal bearings are established to guide the design and conception of engineering solutions critical for stable high-speed bearing operation. The derivation of the scaling laws is based on the above analysis and assumptions made for ultra-short bearing geometries. A simplified Lomakin bearing model is discussed first to shed more light on the underlying mechanisms for the hydrostatic stiffness of ultra-short gas journal bearings. Assuming that the rotor is stationary and that the center of the rotor is displaced from the center of the journal by a small offset e as shown in Fig. 6.9 on the left, the bearing can be approximated as two half-seals acting in parallel between the same bearing pressure drop p. With a nominal bearing clearance C, the upper half of the seal experiences on average a reduced clearance C \u2013 e whereas the lower half of the seal experiences on average an enlarged clearance C + e. Considering the bearing axial cross-section shown in Fig. 6.9 on the right and an inlet stagnation pressure loss coefficient \u03c9inlet = (pt1\u2013 pt2)/(1/2 \u03c1 v2 2), the externally imposed total pressure drop p = pt1 \u2013 pt2 + pt2 \u2013 pt3 can be written as p = 1/2 \u03c1 v2 2 \u00b7\u03c9inlet + p2, where p3is assumed zero and v2 = v3from continuity. The imposed total pressure drop p is the same for both the upper and the lower half of the bearing. However, the flow rate through the upper half of the bearing will be lower compared to the lower half due to the increased resistance (reduced clearance). As a consequence, the stagnation pressure loss at the inlet will be reduced and an increased axial pressure gradient will prevail in the upper half of the bearing as shown in Fig. 6.9 on the right. The pressure gradient in the bearing gap sets up a Poiseuille type flow with a radially averaged axial velocity of v(x) = \u2212 C2 12\u03bc \u2202p \u2202x (6.11) From continuity and the assumption of purely axial flow, the local flow rate remains constant along the gap. Thus, the axial pressure gradient is constant as well and the pressure distribution is linear in the axial direction with \u2202p/\u2202x = \u2013 p2/L. The pressure balance across the bearing between stations 1 and 3 can then be written as p = p2 + 1 2 \u03c1\u03c9inlet C4 144\u03bc2 p2 2 L2 (6.12) The above equation can be used for both the lower and the upper half-bearings by substituting C \u00b1 e for C. Assuming that there is no reverse flow, the inlet pressures p2 of the two half-bearings become pU 2 = A0 (C \u2212 e)4 \u00b7 \u23a1 \u23a3 \u221a 2 p(C \u2212 e)4 A0 + 1 \u2212 1 \u23a4 \u23a6 (6.13) pL 2 = A0 (C + e)4 \u00b7 \u23a1 \u23a3 \u221a 2 p(C + e)4 A0 + 1 \u2212 1 \u23a4 \u23a6 (6.14) where A0 = 144\u03bc2L2/\u03c1 \u03c9inlet. Assuming linear pressure profiles in the axial direction as sketched in Fig. 6.9 and using p3= 0, the direct-coupled hydrostatic restoring force becomes Fhs= DL/2 \u00b7 (p2 U \u2013 p2 L). The hydrostatic stiffness Khs=\u2202Fhs/\u2202e then yields for small shaft offsets e/C << 1 Khs \u2248 4DLA0 C5 \u00b7 f ( 2 pC4 A0 ) (6.15) where f (z) = z/2 + 1 \u2212 \u221a z + 1\u221a z + 1 (6.16) For the range of z of interest f(z) is approximately linear. Rearranging terms the hydrostatic stiffness is found to scale with the bearing parameters according to Khs poC \u221e ( L D )( C R )\u22122 ( p po ) (6.17) where the stiffness is non-dimensionalized by the ambient pressure p0 and the bearing clearance C" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002163_j.mechmachtheory.2007.03.005-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002163_j.mechmachtheory.2007.03.005-Figure4-1.png", "caption": "Fig. 4. The helicoidal vector field associated with homologous lines.", "texts": [ " The slope of vector nb, considering the z-axis as the rising direction, is c \u00bc dWffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0X 2 \u00fe Y 2\u00de q sin / \u00bc dffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0X 2 \u00fe Y 2\u00de=W 2 q sin / \u00bc d r sin / \u00f013\u00de where r denotes the distance from the screw axis to the point. Note that the slope is inversely proportional to r. All the vectors nb, corresponding to all points in space, constitute a helicoidal vector field, which consists of the tangents of equal-pitched coaxial helices, as shown in Fig. 4. The pitch of the helicoidal vector field is p \u00bc d sin / \u00f014\u00de The above finding indicates the following theorem: For any finite displacement screw, there exists a helicoidal vector field associated with homologous lines, and the pitch of the helicoidal vector field is the translation divided by the sine of the rotation. The helicoidal vector field associated with homologous lines is similar to that in instantaneous kinematics, except for the difference in pitch. Furthermore, when the rotation and translation parameters of the finite displacement screw become infinitesimal, the pitch degenerates into that of an instantaneous screw" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002984_s00170-008-1573-7-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002984_s00170-008-1573-7-Figure2-1.png", "caption": "Fig. 2 Schematic of hexapod machine tool", "texts": [ " Machine tools with closed kinematic chains and parallel actuators are being intensively researched as an alternate for open kinematic chain machine tools. These parallel kinematic machine tools have the advantages of higher stiffness, higher pay load capacity, and lower inertia to the manipulation problem [2]. Such parallel kinematic machine tools can meet the high demands on machining accuracy. Based on General Stewart platform [3], as shown in Fig. 1, many parallel machine tools are being built. These are generally known as hexapod machine tools, the schematic of which is shown in Fig. 2. In spite of the advantages possessed by these hexapod machine tools, they have a major drawback, i.e., the stiffness and workspace are highly dynamic and vary with the configuration of the machine tool structure. Because of this, locating the work piece optimally is very difficult. So, using a hexapod machine tool is not as simple as that of conventional machine tools [4]. Several researchers have studied the features of the hexapod machine tools. Conti et al. [5] presented a method to evaluate the variations in the workspace of an octahedral hexapod machine tool during machining and demonstrated the shift in size and location of the workspace as the S" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003565_978-3-642-13769-3_39-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003565_978-3-642-13769-3_39-Figure3-1.png", "caption": "Fig. 3. Second model. The two IRs, and the press machine with the pair of dies.", "texts": [], "surrounding_texts": [ "5.2 Second Model" ] }, { "image_filename": "designv11_61_0002089_1464419jmbd49-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002089_1464419jmbd49-Figure1-1.png", "caption": "Fig. 1 Degrees of freedom in the motorcycle model", "texts": [ " The front suspension is of the telescopic type, whereas the rear suspension is a swing arm with a monoshock spring and damper combination. A four-bar mechanism produces a non-linear spring characteristic. The dampers are linear and viscous. Furthermore, the chain force in the transmission is incorporated to capture its influence on the rear suspension travel. Only conditions for nominally running straight ahead on a level road at a constant forward speed and for negotiating a curve of constant radius are considered here. The motorcycle model is shown schematically in Fig. 1. It consists of six rigid bodies: (a) the rear frame including the rider, (b) the steering head with handlebar and the suspended part of the front fork, (c) the unsuspended part of the front fork, (d) the rear swing arm, and (e, f) the two wheels. It is assumed that the rider is rigidly attached to the rear frame. For nominally running straight ahead, the position of the rear frame with respect to a road-fixed coordinate system is described by the three displacements, X, Y, and Z, of a point rigidly connected to the rear frame, directly below its centre of gravity at the road surface in the plane of the road when the motorcycle is in its upright nominal configuration", " A motorcycle model without suspension system can be obtained by constraining the coordinates xsw and sf to be equal to zero. A stationary motion is characterized by linearly varying values for X, xr, Proc. IMechE Vol. 221 Part K: J. Multi-body Dynamics JMBD49 # IMechE 2007 at The University of Iowa Libraries on March 16, 2015pik.sagepub.comDownloaded from and xf and constant values for all other state variables. For cornering, the displacements X and Y are replaced by polar coordinates C and R, as shown on the right-hand side of Fig. 1. In this case, the origin of the global coordinate system is at the centre of the curve, with the directions of the axes aligned with those of the rear frame in the reference configuration. Instead of the large radius R, the difference from the nominal radius of the curve R0, y \u00bc R02 R, is taken as a generalized coordinate. The yaw angle is similarly split into a nominal yaw angle, C, and a relative yaw angle, cr, as c \u00bc C\u00fe cr. In this case, the vector of generalized coordinates becomes qT \u00bc (C, y, Z , cr, x, f, xsw, xr, b, sf , xf ) T, where C, xr, and xf are cyclic" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003349_978-94-007-6558-0_44-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003349_978-94-007-6558-0_44-Figure2-1.png", "caption": "Fig. 2 The fluid pressure force and the contact forces acting on the inner gear", "texts": [ " After defining the gearing geometry of the gerotor pump gearing pair and after establishing a basic kinematic model, calculation of forces and torques, which act on the gears, can be done. The fluid pressure force which separates the suction inlet chambers from the pressure outlet chambers is a continuous force that can be represented by the equivalent concentrated pressure force Fp. The direction of vector Fp coincides with the centerline of the line segment AB that connects two contact points at the separation borderline between the suction inlet chambers and the pressure outlet chambers zones, as shown in Fig. 2. In accordance with this, the equivalent pressure force in the pump can be expressed in the following vector form: F\u00f0f\u00dep \u00bc Dpbkf AB\u00f0f\u00de: \u00f02\u00de The torque of the equivalent pressure force in respect to the instantaneous rotation center C, is expressed by the multiplication of vectors as: M\u00f0f\u00dep \u00bc CS\u00f0f\u00de F\u00f0f\u00dep ; \u00f03\u00de where S is the middle point of the vector AB (Fig. 2). Consequently the equivalent pressure force, Eq. (2), could be expressed in the following form: F\u00f0f\u00dep \u00bc Dpb y\u00f0f\u00deCB y\u00f0f\u00deCA if x\u00f0f\u00deCB x\u00f0f\u00deCA jf h i : \u00f04\u00de As a result, the torque of the equivalent pressure force in the pump in condensed form is defined as: M\u00f0f\u00dep \u00bc Dpb 2 CAj j2 CBj j2 kf : \u00f05\u00de The resultant of all contact forces that act on the inner gear can be calculated as sum of vectors Fn \u00bc Xz i\u00bc1 Fni; \u00f06\u00de that acts in the kinematic pole C. For the considered pump model, the equilibrium equation of forces could be written in the vectorial form as, Fp \u00fe Fn \u00fe F1 \u00bc 0: \u00f07\u00de Equilibrium equations of the torque in respect to point Ot could be written as MFp\u00f0Ot\u00de \u00feMFn\u00f0Ot\u00de \u00feM1 \u00bc 0; \u00f08\u00de or in respect to point Oa as, MFp\u00f0Oa\u00de \u00feMFn\u00f0Oa\u00de \u00bc 0: \u00f09\u00de The torques from the previous equations could be expressed by the multiplication of vectors in the form as: OtS Fp \u00fe OtC Fn \u00feM1 \u00bc 0 \u00f010\u00de and OaS Fp \u00fe OaC Fn \u00bc 0: \u00f011\u00de Based on the Eqs", " The normal force Fni that acts on the meshed teeth causes local tooth deformation and displacement of the contact point for the value wni in the direction of the force action. It also causes an angular displacement n which is assumed to be equal for all contact points, for selected angular position. If the deformation wni at the contact point Pi is greater than zero it is confirmed that contact is created at that point. If the deformation wni at the contact point Pi is negative or equal zero, the contact is not created and this contact point do not participate in the load distribution. Based on Fig. 2, the total torque about the gear center is equal to the sum of torques of normal forces that act at individual teeth pairs, MFn\u00f0Ot\u00de \u00bc Xq i\u00bcp MFni\u00f0Ot\u00de \u00bc k n Xq i\u00bcp e2 z 1\u00f0 \u00de2sin2 ani; \u00f015\u00de where p and q are the ordinal numbers used to identify the first and the last teeth of the external gear that are transmitting the load and k is the tooth stiffness that is considered to be constant The final form of the contact force Fni is obtained: Fni \u00bc MFn\u00f0Ot\u00de sin anj Pq j\u00bcp e z 1\u00f0 \u00desin2 anj \u00f016\u00de After the procedure of the gear pair identification that transmits the load and the calculation of the contact forces is possible to calculate the contact stresses" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002175_robot.2007.363146-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002175_robot.2007.363146-Figure2-1.png", "caption": "Fig. 2. Twists and wrenches of a RRR(RR) limb", "texts": [ " The other two joints, R4 and R5, intersect at a common point called rotation center. The intersection structure is denoted with parentheses (RR). Five R1 are chosen as actuators. The kinematical screw system[15;16] for one RRR(RR) limb consists of five joint screws ($i, i=1,2,3,4,5) corresponding to five joints, respectively. Since the rank of kinematical screw system is independent to the reference frame, we here assume Z-axis of the reference frame parallel the axis of R1, origin O locate at the rotation center for convenience, as shown in Fig. 2. Then, the kinematical screw system for one limb is $1=[0, 0, 1; y1, -x1, 0]T $2=[0, 0, 1; y2, -x2, 0]T $3=[0, 0, 1; y3, -x3, 0]T $4=[l4, m4, n4; 0, 0, 0]T $5=[l5, m5, n5; 0, 0, 0]T (5) where [x1, y1, 0], [x2, y2, 0] and [x3, y3, 0] are the coordinate of intersection points of axes of R1, R2, R3 with O-XY plane, respectively; [l4, m4, n4], [l5, m5, n5] denotes the direction cosine for axes of R4, R5. According to the screw theory[15;16], the constraint wrench of the limb is a screw which is reciprocal to the kinematical screw system in Eq", " Then, there are two independent constraint wrenches $ r1=[0, 0, 1; 0, 0, 0]T (17) $ r2= [0, 0, 0; 0, 1, 0]T (18) where $ r2 in Eq. (18) is a constraint couple which constrains the rotation around Y-axis. Such a singularity configuration can occurs only in one limb simultaneously. In this case, the end of the limb can not translate along Z-axis or rotate around any axis parallels to Y-axis. B. Actuation singularity When the actuator fixed on the joint R1 is locked, the kinematical screw system under the frame in Fig. 2 is $2=[0, 0, 1; y2, -x2, 0]T $3=[0, 0, 1; y3, -x3, 0]T $4=[l4, 0, n4; 0, 0, 0]T $5=[l5, 0, n5; 0, 0, 0]T (19) There are two independent wrenches, $r1 and $r2, at general configuration for a limb R2R3R4R5. The axis of $r2 for a limb is the intersection line of planes P23 and P45, shown in Fig. 2. In another way, according to screw theory[15;16], $r2=[lr2, mr2, nr2, pr2, qr2, rr2]T (20) where lr2= (1/\u03bb) (x3-x2)/(x2y3- x3y2); mr2= (1/\u03bb) (y3-y2)/(x2y3- x3y2); nr2= 0; pr2= (1/\u03bb) (m4n5-m5n4)/(l4m5-l5m4); qr2= (1/\u03bb) (l5n4-l4n5) /(l4m5-l5m4); rr2= (1/\u03bb); 2 2 2 2 rr ml +=\u03bb ; Assume the unit direction of normal vector for plane P45 is n45=[n45x,n45y,n45z] = [m4n5-m5n4, l5n4-l4n5, l4m5-l5m4], then pr2= (1/\u03bb) n45x / n45z (21) qr2= (1/\u03bb) n45y / n45z (22) $ r2=(1/(\u03bbn45z))[n45zlr2,n45zmr2,n45znr2, n45x,n45y,n45z]T (23) Thus, wrench system $r of the movable platform consists of six wrenches, $ r2 of five limbs and $r1, T)5( 2 )4( 2 )3( 2 )2( 2 )1( 21 ],,,,,[ rrrrrrr $$$$$$$ = (24) where )( 2 i r$ denotes the $r2 of the ith limb, i=1,2,3,4,5" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002441_detc2009-86358-Figure13-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002441_detc2009-86358-Figure13-1.png", "caption": "Figure 13. A Triple Planet Gear", "texts": [ "org/about-asme/terms-of-use 5 Copyright \u00a9 2009 by ASME )cos1(2/ \u03b8+\u22c5= yG (16) \u03b8tan)2/( \u22c5= GC (17) 22 1 2 1 )2/()( GCRR +\u2212= (18) Where, C is drop in the normal plane, G is total gage length and gage from center of crown is G/2 and dT is the clearance between space width of internal teeth and tooth thickness of external teeth. From the above equations, the following relationship can be derived. [ ] 2 mod1 sin/)cos1()cos1(2/2/ \u03b8\u03b8\u03b8 +\u22c5\u2212\u22c5+= TdTR (19) Example 1. A triple planet gear of a trailing edge actuator is shown in Fig. 13. The mating gears are all internal gears. The tangential load at both ends is calculated as 3,600 lbf. The relative deflection under this load is .0015 inch and the misalignment due to backlash and run-out from ring gear and planet gear is .0023 inch total. The face width on the end gear is 1.44 inch. The misalignment is .0008 in/in slope. The relative slope under the load at point B in Figure 13 is .0011 in/in. The total slope is .0019 in/in on point A. At point B, the slope is .0015/1.44 less .0011, and is equal to -.00006 in/in. The total slope is .0008-.00006= .00074 in/in on point B. After solving simultaneous equations, the crowning radius is 143 inch, and crowning center is .64 in from location B. This is a bias crown shown in Fig. 14. A contact stress value of 239 ksi is calculated. Compared to the base line design of crowning radius of 126 inch, crowning center is at the middle of the end tooth and contact stress of 254 ksi, the contact stress is 6% better" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001230_detc2005-84223-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001230_detc2005-84223-Figure5-1.png", "caption": "Figure 5: Various views of the 10 DOF demonstration EE path", "texts": [ "org/about-asme/terms-of-use Dow minimum distance and the average minimum distance between the manipulator and the obstacles were recorded to allow for comparisons of the effectiveness of the individual criteria. Ten of the twelve criteria make use of the influence coefficients derived in this paper. The criteria make use of minimum distance magnitudes, artificial force magnitudes (the forces are based on distances and can be reflected to various locations on the robot using influence coefficients), or their higher order properties. Figure 5 shows views of the demonstration environment. Comparisons (using the smallest minimum distance and the average minimum distance between the manipulator and the obstacles as measures of performance) between the twelve criteria tested were encouraging but mixed. No clear \u2018best\u2019 criterion was determined, but all of the obstacle avoidance criteria tested were shown to be (at least somewhat) suitable for use in manipulator obstacle avoidance. This example also showed that in some situations criteria that use higher-order properties (which are all based on the influence coefficients derived in this paper) can improve manipulator performance" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001489_s00170-006-0516-4-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001489_s00170-006-0516-4-Figure2-1.png", "caption": "Fig. 2 The 3-RPRU simulation mechanism", "texts": [ " The three identical RPRU limbs connect m with B by a universal joint U at point ai, a revolute joint R for connecting link ci and a driving rod ri, a prismatic joint P that includes a hydraulic cylinder and piston-rod, and a revolute joint R at point Ai for i=1, 2, and 3. 2.1 The simulation mechanism of the 3-RPRU parallel manipulator Based on the 3-RPRU parallel manipulator, a simulation mechanism of the 3-RPRU parallel manipulator is created in the 3D Sketch environment of Solidwork software, as shown in Fig. 2. Its creation processes are explained as follows. Step 1 Constitute an equilateral triangular \u0394A1A2A3 by using the polygon command, and set its central point O coincident with the coordination original point. Transform \u0394A1A2A3 into a plane by using the planar area command. Give each sideline of \u0394A1A2A3 an initial dimension Li=120 cm (i=1, 2, 3). Thus, an equivalent base B is constituted. Step 2 Constitute three lines ri (i=1, 2, 3), and connect them to form a closed equilateral triangle\u0394a1a2a3. Constitute a line c and connect its two ends to point a2 and sideline a1a3, then set c perpendicular to a1a3. Constitute a line d, and connect its two ends to a3 and c at O. Set d perpendicular to sideline a1a2. Give each sideline of \u0394a1a2a3 a dimension ri=80 cm (i=1, 2, 3). Thus, an equivalent moving platform m is constituted, and its central point O can be determined from the coincident point of c and d, as shown in Fig. 2. Step 3 Constitute three driving limbs RPRU for connecting m with B. The processes are explained as below: Constitute lines ci and ri (i=1, 2, 3), give ci a fixed dimension in length ci=40 cm, and give ri a driving dimension in length ri=80 cm. By using the point to point coincident command, connect the two ends of ri to the one end of ci and B at Ai. Connect the other end of ci to m at ai (i=1, 2, 3). Step 4 Constitute a prismatic joint P, two revolute joints R, and a universal joint U in the first driving limb", " \u2013 Set r1 perpendicular to c1 by using the perpendicular constraint command; thus, an equivalent revolute joint R for connecting l1with c1 is constituted. \u2013 Constitute an auxiliary line b1, connect it to m at a1, set b1 perpendicular to both c1 and a2a3 of m by using the perpendicular constraint command, and set b1 parallel to r1 by using the parallel constraint command; thus, an equivalent universal joint U for connecting c1 with m is constituted. Step 5 Similarly, the other two driving limbs can be constituted. In this way, a simulation mechanism of the 3-RPRU parallel manipulator is created, as shown in Fig. 2 2.2 DOF of spatial 3-RPRU parallel manipulator The number of degree of freedom (DOF) for the 3-RPRU parallel manipulator can be calculated by using the Kutzbach Grubler equation [11]: F \u00bc \u03bb\u00f0k n 1\u00de \u00fe Xn i\u00bc1 fi (1) where k is the number of links, n is the number of joints, \u03bb is the degrees of the space within which the mechanism operates for spatial motions (\u03bb=6); and fi is the degree of freedom of the ith joint. By inspecting the whole mechanism of Fig. 1, it is known that \u03bb=6 for a spatial mechanism", " The coordination of the three vertex points ai of \u0394a1a2a3 in o-xyz and O-XYZ are {xai, yai, zai} and {Xai, Yai, Zai}, respectively. The coordination of Ai of\u0394A1A2A3 inO-XYZ are {XAi, YAi, ZAi}. From results of computer simulation, it known that there are only three translations for the spatial 3-RPRU parallel manipulator. Therefore, a transformed equation is derived such that: Xai; Yai; Zai; 1f gT \u00bc 1 0 0 Xo 0 1 0 Yo 0 0 1 Zo 0 0 0 1 2 6664 3 7775 xai; yai; zai; 1f gT (2) After extending Eq. 2, we get: Xai \u00bc xai \u00fe Xo; Yai \u00bc yai \u00fe Yo; Zai \u00bc zai \u00fe Zo (3) From coordination in Fig. 2, the coordination of ai in oxyz and Ai in O-XYZ are obtained such that: xa1 ya1 za1 8>< >: 9>= >; \u00bc ffiffiffi 3 p r=2 r=2 0 8>< >: 9>= >;; ::: xa2 ya2 za2 8>< >: 9>= >; \u00bc 0 r 0 8>< >: 9>= >;; ::: xa3 ya3 za3 8>< >: 9>= >; \u00bc ffiffiffi 3 p r=2 r=2 0 8>< >: 9>= >; (4) XA1 YA1 ZA1 8>< >: 9>= >; \u00bc ffiffiffi 3 p R=2 R=2 0 8>< >: 9>= >;; :: XA2 YA2 ZA2 8>< >: 9>= >; \u00bc 0 R 0 8>< >: 9>= >;; :: XA3 YA3 ZA3 8>< >: 9>= >; \u00bc ffiffiffi 3 p R=2 R=2 0 8>< >: 9>= >; (5) Based on the feature of the 3-RPRU parallel mechanism, a constraint equation is obtained: ri \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0XAi xai Xo\u00de2 \u00fe \u00f0YAi yai Yo\u00de2 \u00fe \u00f0ZAi zai Zo\u00de2 c2 q ; i \u00bc 1; 2; 3 (6) Set R0=R\u2212r= (Li\u2212li)/tan60\u00b0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001489_s00170-006-0516-4-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001489_s00170-006-0516-4-Figure3-1.png", "caption": "Fig. 3 A 3D free surface of the 3-RPRU simulation parallel machine tool Fig. 4 The guiding plane of the tool path", "texts": [ " These results show that the 3-RPRU simulation mechanism is an effective mechanism of the actual 3-RPRU manipulator, and can be used to create a spatial 3-RPRU parallel machine tool. 3.1 The 3D free surface S Before creating the 3-RPRU simulation parallel machine tool, a 3D free surface S must be created by using the 3D modelling technique. The creating processes are explained below. 1. Modify the base B of the 3-RPRU simulation mechanism in the 2D sketching environment, constitute several datum planes Pi, and set all datum planes parallel to each other and perpendicular to B by adopting the reference plane command, as shown in Fig. 3. 2. Based on the prescript curve data or curve equation, constitute a spline curve ui on the ith datum plane Pi (i=1, 2, ... j) by using the spline command or data table, and set each spline curve to be above the moving platform m, as shown in Fig. 3. 3. Constitute S from all ui (i=1, 2, ... j) by using some special modelling techniques, such as loft, swept, extrude, and rotation commands. Here, S is constituted by adopting a loft modeling technique, as shown in Fig. 3. 3.2 The guiding plane P0 of tool path In order to create a reasonable tool path, a guiding plane P0 of T should be constituted. Its creation processes are explained below. 1. Modify B of the 3-RPRU simulation parallel machine tool, constitute a sketching datum plane by using the reference plane command, and set it parallel to B, as shown in Fig. 4. 2. In order to retain P0 close to and without intersecting with S, give the distance h from the sketching datum plane to B an initial driving dimension h=160 cm, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002412_icems.2009.5382984-Figure9-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002412_icems.2009.5382984-Figure9-1.png", "caption": "Fig. 9. Vibration mode of the motor at the natural frequencies", "texts": [ " The reasons of the difference between the measurement result and the analysis result are described as follows. First, the mechanical characteristic of stator case is different from the actual one. Second, the contact condition between the stator core and the case is unsuitableness. From these reasons, the natural frequency of the case should be measured, and the contact condition should be examined. The natural frequency of the motor is measured and is compared with the analysis result. The vibration mode and the natural frequencies are shown in Fig. 9 and Table IV The first natural frequency between the measured and the analyzed is widely differed. This is due to the contact condition between the rotor and the stator. However this can be ignored, because the vibration mode has little effect on the deformation of the case. TABLE III NATURAL FREQUENCY OF STATOR (a) 2118Hz (b) 2776Hz Fig. 8. Vibration modes of stator at the natural frequencies TABLE IV NATURAL FREQUENCY OF MOTOR Order Measured(Hz) Calculated(Hz) 1st 2300 2118 2 nd 2950 2766 Order Measured(Hz) Calculated(Hz) 1st 1100 1938 2 nd 2850 2967 0 2000 4000 6000 8000 10000 12000 0 2000 4000 6000 8000 10000 12000 Measured Natural Frequency (Hz) C al cu la te d N at ur al F re qu en cy (H z) Shaft RotorCore StatorCore The acceleration under the driving condition is measured by the acceleration sensors which mounted on the side of the stator case" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001775_iecon.2006.348122-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001775_iecon.2006.348122-Figure3-1.png", "caption": "Fig. 3. Rotor dimensions of two prototypes", "texts": [ " To explain this, analysis of neglecting the cage is compared with experimental values. And by using the model of considering the cage and experimental values, the effect of the cage is considered. Fig.2 shows the prototype of single-phase LSPM tested in this paper. The structure of this prototype is that the arc shaped PM is inserted in the rotor of multipurpose single-phase induction motor. So no optimization technique is done for this structure. Two types of rotors are constructed as shown in Fig.3. By comparing analysis of Rotor A with experimental value, the effect of squirrel cage in single-phase LSPM is examined. Furthermore, characteristics of Rotor A and Rotor B calculated by using FEM is shown to discuss the magnetic circuit. Table I shows the main design specification of the prototypes. Material of rotor cage is aluminum, and the number of conductor bar is 28. Concerning PM, arc 138deg magnet (per pole) magnetized parallel is mounted in the rotor. Material is Nd-Fe-B, and residual flux density and coercive force are Br =1" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002074_j.jappmathmech.2007.09.001-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002074_j.jappmathmech.2007.09.001-Figure2-1.png", "caption": "Fig. 2.", "texts": [ " In solving the problem it is assumed that the stress-strain state of the elastic half-space in the neighbourhood of each asperity is unaffected by the other asperities. This assumption holds if the distance between the asperities is fairly large.8 Consider the position of a rough cylinder for which the lowest asperity is situated symmetrically about the z axis (Fig. 1). Suppose the depth of penetration of this asperity into the elastic half-space, which coincides with the maximum depth of penetration of the whole cylinder, is equal to c, while the depth of penetration of a certain i-th asperity is ci (Fig. 2). From the triangles ABC and AOC (O is the centre of the cylinder, A is the vertex of the i-th asperity and C is the vertex of the central asperity) we have for the length of the section AC Assuming that l R, we obtain i = il/R. Then, the penetration of the i-th asperity is given by the relation (2.1) where i takes values from 1 to Nr on the right of the central asperity, and from 1 to Nl on the left of it. In all, in the interaction with the half-space there are Nr + Nl + 1 asperities. Hence, knowing ci we can consider the interaction of the i-th asperity with the half-space. We will introduce a local cylindrical system of coordinates (ri, i, zi) with origin at the centre of the contact area. Since the asperity has a spherical shape, the distributions of the pressures and elastic displacements in the half-space are symmetrical about the zi axis in the neighbourhood of an asperity (Fig. 2). Inside the contact area Ai, which is a circle of radius ai, the contact condition for a displacement u along the zi axis of the boundary of the elastic half-space is satisfied. By relation (1.1) this condition has the form (2.2) It follows from condition (1.2) and the symmetry of the problem that, in the region of adhesive interaction Bi, which is a ring ai < ri < bi, the pressure p on the boundary of the elastic half-space is determined by the adhesive attraction of the surfaces; outside the region of interaction the pressure is equal to zero: (2" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001735_epe.2005.219218-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001735_epe.2005.219218-Figure3-1.png", "caption": "Fig. 3: Voltage reference and inverter output voltage vectors", "texts": [ " k) MR2 i* (k) (11) L2 02d (k) O(k) = E {Pom(k) + cse(k)} Ts (12) EPE 2005 Dresden P.3 ENDO Ryo 2005 - Dresden ISBN: 90-75815-08-5 P.3 New Field-Weakening Control Considering Voltage Saturation for Vector Control System of Induction Motor where, p is a number of pole pair. Each current PI controller on the d-q rotating frame determines the voltage reference of the space voltage vector inverter V]d (k), vlq*(k). The voltage reference on a stationary reference frame vla*(k), vI (k) are calculate from Vld*(k), vlq*(k) as shown in (13). Figure 3 shows relation characteristics between the voltage reference vector vj*(k) and inverter voltage vectors as shown in figure 3 (a). The voltage reference vector is allocated by three voltage vectors including zero voltage vector V7. The inverter output voltage V1, V2 and V7 are realized by the inverter switching command Su, Sv and S. In the case of figure 3 (a), the voltage reference vector is expressed as shown in (14). via (k)] Fcos 0(k) - sin 0(k)]Fvl (k)] (13) v (k)j L sin 0(k) cos O(k) jLv (k) t ((k) t2(k) = VDC 1/2 ]t, (k)1 2VDC FUla U22 t1(k)1 (14) V1 +V2 s 3 TsL\u00b0 /2Lt2(k)_ 3 T. _u1 U2/J _t2(k)j where tl(k), t2(k) are voltage-supplying period, VDc is a dc link voltage value of the inverter. The voltagesupplying period tl(k) and t2(k) are calculated from voltage reference vlia(k) and vl]f*(k) as shown in (15) and (16). Dc v (k)U2,, VI* 1 t (k) = VDC (k)a (k) *(k)U (15) 2T u U=lU2a VDC v* kul v* (k)ul1, F2Tk)1 t2(kDC _(k)uja vl kuP(16) 2 Ts U1Uf Ul/JU2a VDC EPE 2005 Dresden P.4 ENDO Ryo 2005 - Dresden ISBN: 90-75815-08-5 P.4 New Field-Weakening Control Considering Voltage Saturation for Vector Control System of Induction Motor When the total time of tl(k) and t2(k) is longer than the sampling period T, the inverter cannot output the voltage corresponding to its reference due to the voltage saturation. In this case, both tl(k) and t2(k) are shortened as shown in (17). The amplitude of voltage vector is reduced up to the maximum voltage of the inverter as shown in figure 3 (b). The difference between the actual output voltage and the voltage reference on d-q rotating frame are calculated by (18) and (19). 71 (k) t, (k) TS 72(k)= t2(k) TS ~~~~~~~(17) t, (k) + t2 () t k2(k) t, (k) + t2 (k) (7 JVia ]Lv*(k)1 VDC FUIa U2 it,(k) - (k)1 (18) AJvI,8 ]vI)>k)j 3 TS LU, U2 ILt2(k)- 72(k)] LAvld(k)1 cosO sinO lFJvla(k)1 (19) zfvlq (k)] L- sin cos OIL0 Vl (k)j When the inverter output voltage is limited by voltage saturation, the state variables of current PI controllers do not corresponding to the actual output voltage" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001824_50041-2-Figure1.3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001824_50041-2-Figure1.3-1.png", "caption": "Figure 1.3.1 Internal and external forces in the flow domain . The internal forces cancel at all internal points due to the action = reaction rule and only the surface points remain for the internal forces contributions.", "texts": [ " What we only can see are the effects of forces, for instance the displacement of an object due to gravitational forces, or the displacement of a pointer on a measuring instrument under an electric potential difference indicating the presence of an electrostatic or electromagnetic force. In fact one of the more fundamental assumptions of modern physics is to consider that when we observe a certain effect, we assume the existence of a force behind it, as its cause. This is exactly what is considered in fluid mechanics: since a fluid can sustain internal deformations, a force, which is called the internal force of the fluid, must cause these deformations (see Figure 1.3.1). We refer you to your fluid mechanics courses for more details, and we will summarize here only the main properties. The most important is the definition of the internal force, acting on a surface element dS. In the general case, the internal force acting on this surface element depends both on its position and on its orientation, defined by the normal. Therefore it should be described mathematically by a tensor \u03c3, such that the local internal force vector is written as fi = \u03c3 \u00b7 n (1.3.4) where n denotes the unit normal vector to the surface element" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001259_jmes_jour_1972_014_043_02-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001259_jmes_jour_1972_014_043_02-Figure5-1.png", "caption": "Fig. 5. An example of synthesis", "texts": [ "4 - The configuration (group 3 0 2, Type 5) has been selected since it is one of the cases for which none of the elemental p.g.ts has a speed ratio equal to one of the overall speed ratios, and the numerical values for the ratios have been chosen so that power recirculation occurs in one \u2018gear\u2019: this has been done so that power-flow diagrams showing cases with and without power recirculation are obtained. Determination of the speed ratios of the elemental p.g.ts For convenience, the schematic diagram for the selected arrangement is shown in Fig. 5a. The required speed ratios are 0.5, 0.63 and 0.8, giving The connection matrix for this type is C = - 1 0 0 0 0 0 0 0 1 0 0 1 0 0 0 1 0 0 0 0 0 0 0 0 1 0 1 0 0 0 1 0 0 0 0 0 0 1 0 0 ~ 0 1 0 0 0 Substitution into equation ( 1 1) gives and evaluation of equation (14) gives -3+4.4 = 0-259 = 1+4-4 -3+4.4 2 - ~ = 0.412 - -1+4.4 -3+4.4 Q 3 = - - 9 $ - 4 4 - - -0.305 The configurations of the individual elemental trains may now be determined, and these are summarized in Table 4. A diagrammatic representation of a possible layout is shown in Fig. 5b. Having evaluated the proportions of the trains the next step is the calculation of the running speeds of the elements and the loads which they transmit. Journal Mechanical Engineering Science Vol14 No 5 1972 at UNIV CALIFORNIA SAN DIEGO on April 20, 2016jms.sagepub.comDownloaded from A = - - 0 0.26 0.6 1 1 1 -0.352 0 0.46 0 0.26 0.6 0.5 0.63 0.8 -0.352 0 0.46 0 0.26 0.6 -1.5 -0.85 0 - -0.352 0 0.46 - Speeds of the shafts in each gear Equation (17) is so that P B = ' I 1 1 0.5 0-63 0.8 B = [ 0 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002973_robio.2009.5420758-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002973_robio.2009.5420758-Figure3-1.png", "caption": "Fig. 3 Mechanical structure of SAYA", "texts": [ " In fact, a high correct recognition rate of 6 typical facial expressions that are possible to be recognized and express universally was achieved in previous research [12]. Fig. 2 shows samples of facial expressions that the face robot can express. Since the android robot SAYA equips the face robot on a dummy body as a head part, it looks like a real human as shown in Fig. 1(a). Furthermore, we applied the android robot SAYA to the reception in our university entrance by implementing speech dialogue function [13]. Fig. 1(c) and Fig. 3 show internal structure of the face robot. We use McKibben artificial muscle [16][17] for actuating the control points. Since it is small, light and flexible, it can be distributed to curved surface of the skull such as human muscles. In addition we form the facial skin with soft urethane resin to realize the texture of human facial skin. The face robot has an oculomotor mechanism that controls both pitch and yaw rotations of eyeballs by 2 DC motors. The two eyeballs move together since they are linked to each other", " Moreover we mounted a CCD camera to inside of the left-side eyeball. Thus, the eye direction is able to be controlled by recognizing the human skin color region from image of the CCD camera so that the face robot can pursue the visitor. In order to realize flexible neck motion like a human cervical spine, we adopted a coil spring for the head motion mechanism. Furthermore, the center of rotations for pitch rotation (\u201cPitch1\u201d) and yaw rotation were set in the base of the head. Movable positions and movable ranges were defined as shown in Fig. 3 by referring to anatomical knowledge. The forward and backward motions are realized by combination of the head rotation (\u201cPitch1\u201d) and the neck bending (\u201cPitch2\u201d). The roll-rotation, both pitch-rotations (\u201cPitch1\u201d and \u201cPitch2\u201d) and the yaw-rotation are also actuated by McKibben artificial muscle as shown in Fig. 4. We use electro-pneumatic regulators for controlling contraction of McKibben artificial muscle. Although various kinds of autonomous robots have been studied and developed a lot, intelligence of robots is lacked still to interact with human and act in daily lives automatically" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003076_elan.200904641-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003076_elan.200904641-Figure2-1.png", "caption": "Fig. 2. Schematic diagram of an ALP-based electrochemical biosensor using NPP and an avidin-modified ITO electrode.", "texts": [ "1 V compared to that in curve i of Figure 1b. However, there was no significant change in the peak current. These results show that there is almost no background current due to NPP and that the avidin/biotin/ APPA layer on the ITO electrodes does not significantly change electrocatalytic activities of the electrodes. More importantly, these results show that ALP product (NP) can be measured without electrochemical interference of the ALP substrate (NPP), even without modifying ITO electrodes with any electrocatalytic materials. Figure 2 shows a schematic diagram of an ALP-based electrochemical biosensor using NPP and an avidin-modified ITO electrode. To test the feasibility of the sensor, a sandwich-type immunosensor for detecting mouse IgG was designed. Biotinylated antimouse IgG was immobilized on avidin-, biotin-, and APPA-modified ITO electrodes to capture target mouse IgG [5, 6, 19]. ALP-conjugated antimouse IgG was attached after the mouse IgG was captured [5, 6, 19]. The ALPs bound on the electrode converted electrochemically inactive NPP into electrochemically active NP" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003358_978-1-4471-4426-7_2-Figure11-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003358_978-1-4471-4426-7_2-Figure11-1.png", "caption": "Fig. 11 True Demand Chain", "texts": [ " This implies that until now, industries are operating in the producer-centric framework and they do not necessary have to consider the basic needs or the basic expectations of the customer. But from now on, they have to get down to the basics of what customers really expect from them. S. Fukuda To describe this in terms of supply chain and demand chain, the traditional Supply Chain and Demand Chain concept was as shown in Figure 10. But Customerdriven Collaboration of the Producer and the Customer will lead to the true demand chain as shown in Figure 11, where customer\u2019s true or basic expectations will be realized as a product. 24 Concurrent Engineering in a New Perspective: Heading for Seamless Engineering 25 CE tomorrow will be heading toward seamless engineering and what we will be designing are not products but societies. There will be no walls between products and/or between industries and we will be fusing them to build up a smarter community and create a new lifestyle. S. Fukuda 10 References 1. http://www.terrafugia.com 2. http://en.wikipedia" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003713_1.3356876-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003713_1.3356876-Figure1-1.png", "caption": "Fig. 1. The three solid objects that make up our model are joined together by the S vectors. Sij is embedded in the body of the ith object and points to the joint that is shared with the jth object.", "texts": [ " In particular, we consider a two-dimensional articulated figure2 with two degrees of freedom. A formalism is developed for calculating the rotation of this model when it has zero angular momentum. It is shown that even if propellorlike motion 733 Am. J. Phys. 78 7 , July 2010 http://aapt.org/ajp Downloaded 28 Oct 2012 to 152.3.102.242. Redistribution subject to AAPT lic were forbidden, rotation is still possible. The formalism is then applied to a diver performing a zero angular momentum dive, and the rotation of the body during such a dive is calculated numerically. Figure 1 shows the model that we consider. It is confined to move in the x-y plane and is made up of three solid objects whose masses and moments of inertia about their centers of mass are given by mi and Ii. The center of mass position of the ith object is given by ri=xii+yij, and its orientation is given by i, the angle between some arbitrary direction axis embedded in the ith object and the +x direction. The objects are connected at joints that are defined by the vectors Sij i . The vector Sij points from the center of mass of object i to the joint it shares with object j. It is fixed in the body of object i, and hence its magnitude is constant, but its direction changes with i. If the three objects are unconnected, each object has two translational and one rotational degree of freedom, and thus the system has nine total degrees of freedom. To create an articulated figure, we join the objects together by imposing the conditions see Fig. 1 r1 + S12 = r2 + S21, 1a r1 + S13 = r3 + S31, 1b which reduce the number of degrees of freedom from nine to five. If we further require that the center of mass Rcm be fixed at the origin, then i=1 3 mi Rcm = i=1 3 miri = 0, 2 which reduces the number of degrees of freedom to three. Because we need only three variables to uniquely define the state of such a system, we could choose 1, 2, and 3 to describe the orientation of the objects. We will find that it is more convenient to choose the variables , , and , given by 733\u00a9 2010 American Association of Physics Teachers ense or copyright; see http://ajp", "6 To make the figure move realistically, the time span for the nine frames was taken to be 1.20 s which means each of the shown frames occurs at intervals of 0.15 s ; a time step of 0.01 s was used. The change in the figure\u2019s orientation from the first to the ninth frame is 1.82294 rad, in agreement with the result from Eq. 16 . Can a diver, without the help of any torque exerted by the diving platform on the feet, still execute a dive? This question has been addressed7 but is worth revisiting using the formalism we have developed. The model depicted in Fig. 1 can be modified to represent a crude model of the side view of a human figure. The i=1 piece is the torso and head, represented by a rod with a sphere connected at the end. The i=2 piece is the arms, which move in unison and are represented by a single rod. used to construct the simple model in Sec. III. 735S. John Di Bartolo ense or copyright; see http://ajp.aapt.org/authors/copyright_permission The i=3 piece is the legs, which we represent by two rods, one for the thighs and one for the calves, fused together at the knee", "2 s, the duration of a typical dive which means each of the shown frames occurs at intervals of 0.15 s ; a time step of 0.01 s was used. The change in the diver\u2019s orientation from the first frame when the diver is on the diving board to the ninth frame when the diver is in the water is 1.10489 rad, in agreement with the result from Eq. 18 . Even though the diver\u2019s angular momentum is zero, we see that is possible for the diver to experience rotation during a dive, in this case a clockwise rotation of 1.1 rad.10 We have shown that the articulated figure depicted in Fig. 1 can move its joints in such a way as to experience a rotation even though it has zero angular momentum. Unlike typical examples of rotation with zero angular momentum, the model moves in two dimensions, and the parts do not exhibit propellorlike motion. If we limit the number of degrees of freedom to one, the ability of the model to rotate itself would be lost. Such movement would correspond to either or remaining fixed, and the contour representing this movement would be Table I. Values used for the human body model in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002250_978-3-540-92841-6_559-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002250_978-3-540-92841-6_559-Figure6-1.png", "caption": "Fig. 6. Human Model", "texts": [ " 23 ___________________________________________ III. USER MODEL FOR WEARABLE ASSIST SYSTEM In this section, we introduce a control algorithm of the wearable walking support system as shown in Fig.5 based on the human model. Firstly, we derive the knee joint moment based on an approximated human model, and then the support joint moment, which should be generated by the actuator of the support device, is calculated. To control Wearable Walking Helper, we use an approximated human model as shown in Fig.6. Under the assumption that the human gait is approximated by the motion on the sagittal plane, we consider only Z - X plane. The human model consists of four links, that is, Foot Link, Shank Link, Thigh Link and Upper Body Link and these links compose a four-link open chain mechanism. To derive joint moments, we first set up Newton-Euler equations of each link. At the link i, Newton-Euler equations are derived as follows: where, fi-1,i, and fi,i+1 are reaction forces applying to the joint i and i + 1, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002074_j.jappmathmech.2007.09.001-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002074_j.jappmathmech.2007.09.001-Figure3-1.png", "caption": "Fig. 3.", "texts": [ " Interaction of the asperity with the half-space may also occur when there is no direct contact between the surfaces. In this case a negative adhesive pressure-p0 acts on the half-space in a circular area of radius bi. When solving this problem, contact condition (2.2) is not used, and it is assumed that ai = 0 in the remaining conditions. As a result, for the case when there is no contact, we obtain the relations (2.10) The dependence of the normal force qi, acting on the asperity from the side of the elastic half-space, on the depth of penetration ci has the form shown in Fig. 3. The heavy curve corresponds to contact of the surfaces and relations (2.8) and (2.9), while the thin curve is for the case when there is no contact and relations (2.10). It can be seen that the dependence of the force on the depth of penetration is described by a non-unique function. According to this relation, as the depth of penetration ci increases from -\u221e (as the asperity approaches the elastic half-space) the surfaces intermittently interact when ci = cr; when the depth of penetration increases further the interaction is described by curve 1", " We will apply the solution obtained above for the interaction of a single asperity with an elastic half-space to the problem of the rolling of a rough cylinder. For a specified maximum depth of penetration of the cylinder into the half-space c, which is identical with the depth of penetration of the central asperity, the depths of penetration of the remaining asperities ci are given by relation (2.1). The solution obtained above enables us, from these quantities, to determine the values of the normal forces qi, acting on each asperity from the side of the elastic half-space (as can be seen from Fig. 3, these forces can be positive or negative depending on the values of ci) and other characteristics of the contact interaction \u2013 the radii of the contact area ai and the regions of adhesive interaction bi. The solution for the i-th asperity will depend on the depth of penetration ci and on whether this asperity approaches the elastic half-space or moves away from it: if it approaches, the force acting on the asperity is determined by the function qr i (ci); if it is moving away from it, the force is determined by the function ql i(ci)", " The results show that taking adhesion into account leads to a nonunique dependence on the load of both the nominal and the actual contact areas, and also to the existence of contact in a certain region of negative loads on the cylinder. The solution of the problem when the adhesive interaction of the surfaces is ignored gives reduced values of both the nominal and actual contact areas. As follows from the solution of the problem considered in Section 2 for an individual asperity, the relation between the force acting on one asperity qi and the depth of penetration of this asperity ci (Fig. 3) is non-unique. It follows from this that for a cyclical approach of the asperity to a half-space and its separation from it a loss of energy occurs provided that the greatest depth of penetration of the asperity per cycle exceeds cr. The value of this energy loss corresponds to the area shown hatched in Fig. 3 and is given by the expression (4.1) The energy loss in a complete rotation of the cylinder is wN1, where N1 is the number of asperities in a section of the cylinder, for which the maximum depth of penetration into the half-space after a single rotation of the cylinder exceeds cr. Assuming that this energy loss is equal to the work of the moment of the resistance to rolling per single rotation of the cylinder 2 M, we obtain the following expression for the moment of the resistance (4.2) In the model of a rough cylinder considered in the previous sections, which has N similar asperities in a section, the number N1 is given by a step function (4" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002607_978-0-387-77747-4_6-Figure6.1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002607_978-0-387-77747-4_6-Figure6.1-1.png", "caption": "Fig. 6.1 Rotordynamic arrangement of radial bearings for large-scale gas turbine engines (left) and MEMS-based microengines (right). Note that the microengine blade-tip speed is equal to the bearing surface speed", "texts": [ "hapter 6 High-Speed Gas Bearings for Micro-Turbomachinery Zolt\u00e1n S. Spakovszky The mechanical design and architecture of high-speed rotating machinery, independent of size or scale, are strongly governed by the rotordynamic behavior of the spool and its bearing arrangement. Large-scale gas turbine engines yield multispool shaft constructions where the rolling contact bearings are close to the centerline of the engine supporting the shaft and disk assemblies as shown in Fig. 6.1 on the left. This architecture is governed by the achievable surface speed of the bearings, commonly characterized by the DN number. The DN number is defined as the Z.S. Spakovszky (B) Gas Turbine Laboratory, Massachusetts Institute of Technology, Bldg 31-265, 77 Massachusetts Avenue, Cambridge, MA 02139, USA 191J.H. Lang (ed.), Multi-Wafer Rotating MEMS Machines, MEMS Reference Shelf, DOI 10.1007/978-0-387-77747-4_6, C\u00a9 Springer Science+Business Media, LLC 2009 diameter of the shaft, D, in mm times the shaft speed, N, in rpm. Conventional ball bearings can reach DN numbers of up to 1\u20132 million mm-rpm. Since the turbomachinery is operating at much higher speeds than the bearing surface speed, the bearings must be located close to the shaft centerline as indicated by the solid red circles in Fig. 6.1 on the left. The architecture of MEMS-based micro-gas turbine engines is vastly different compared to large-scale engines. The powerMEMS devices are made out of multiple wafers stacked and bonded together to form the mechanical structure. Dictated by DRIE fabrication constraints and the inherently flat rotor architecture, the bearings are located on the outer periphery of the rotor. Since the bearing surfaces are running at the blade-tip speed, the bearings must be designed to operate at DN numbers of up to about 10 million mm-rpm, which is one order of magnitude higher than in conventional rolling element bearings", " In the light of the above requirements, there are three main levitation and support concepts to be considered for both journal (radial) and thrust (axial) bearings: rolling element bearings, electro-magnetic bearings, and fluid film bearings. As discussed earlier, the logical choice for MEMS turbomachinery applications is fluid film bearings. More specifically, gas (i.e., air) bearings have no temperature limit, yield high-load capacity, are relatively simple to manufacture, and allow oil-free operation. Gas bearings are also found in large-scale application such 30\u201370 kW micro-GTs, APUs, turbochargers, gyroscopic instruments, and dentist drills. Returning to Fig. 6.1, the gas journal bearings supporting the microturbomachinery are located at the outer periphery of the rotor. Given a typical rotor diameter of 4 mm and a blade-tip speed of about 500 m/s, they must be designed to operate at DN numbers of up to about 10 million mm-rpm, equivalent to a rotor shaft speed of 2.4 million rpm. Since these journal bearings are also very short compared to their relatively large bearing diameters, the bearing length to diameter ratio is typically 0.1. This inherently alters the driving flow mechanisms in the bearing journal and the rotordynamic stability of the rotor-bearing system, setting new challenges in the design, implementation, and operation of very high-speed gas bearings" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002236_2007-01-3740-Figure9-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002236_2007-01-3740-Figure9-1.png", "caption": "Figure 9b: T-CVT 2 Dimensions", "texts": [], "surrounding_texts": [ "Torotrak have well proven software modelling tools to simulate variable drive transmission performance. These tools include validated models of Variator behaviour derived from extensive test data acquired from many different Variator, transmission and vehicle tests [5]. To evaluate the performance of the T-CVT, Torotrak have created a generic A / B-class vehicle model comprising engine, 4 speed automatic transmission and torque converter models including torque converter lock up schedule and transmission shift maps. Starting with the durability of the T-CVT, a generic vehicle duty cycle representing a life of 160,000km was utilised and is shown in Figure 11 (only the positive wheel torque quadrant is shown). Estimated vehicle dutycyle Combining this dutycycle data with the optimised engine operating conditions produces the speed, torque and ratio of the Variator for each point of the dutycycle matrix. Hence the load and number of stress cycles are defined. From this information, a standard bearing life ninth power law is employed to calculate the consumed life fraction of the discs and rollers. The application of this rule is upon the basis of extensive experimental data produced from durability testing of real disc and roller components [6]. Finally all the life fractions are summed and the total consumed life fraction is calculated. Table 3 shows the predicted Variator life used over the 160,000km dutycycle. Variator ratio Input disc Output disc Roller -0.40 to \u20130.52 1.66 % 0.07 % 0.67 % -0.52 to \u20130.64 0.93 % 0.10 % 0.47 % -0.64 to \u20130.76 1.15 % 0.25 % 0.73 % -0.76 to \u20130.88 0.38 % 0.16 % 0.31 % -0.88 to \u20131.00 0.53 % 0.37 % 0.53 % -1.00 to \u20131.30 0.55 % 1.18 % 0.98 % -1.30 to \u20131.60 0.40 % 1.43 % 0.98 % -1.00 to \u20131.90 0.18 % 1.96 % 0.97 % -1.90 to \u20132.20 0.79 % 15.89 % 6.78 % -2.20 to \u20132.50 0.27 % 12.81 % 4.55 % Total 6.84% 34.22% 16.97% Table 3: Variator Component Life Fractions Hence for the applied dutycycle, only circa a third of the life of the transmission is consumed. Whilst calculating the Variator life the simulation tools also calculate the transmission efficiency for each of the dutycycle points. The overall transmission efficiency including the final drive over a typical dutycycle is shown in figure 12. It can be seen that for the majority of the operating conditions the overall transmission efficiency, including the final drive, is approximately 90%. This high efficiency, together with the ability to optimise the operation of the engine, produces excellent overall powertrain efficiency. Applying this transmission efficiency with the engine operation produces the fuel consumption of the powertrain. Due to the ability of the T-CVT to optimise the operation of the engine, a substantial fuel economy benefit is obtained. Comparing the T-CVT to a 4AT and belt CVT over the Japanese 10-15 fuel cycle produces the results in Figure 13." ] }, { "image_filename": "designv11_61_0000292_2004-01-0867-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000292_2004-01-0867-Figure8-1.png", "caption": "Fig. 8 Simplified geometry of the plate tab", "texts": [ " The predicted results are compared to the experimental results in Figure 6. The non-Newtonian, adiabatic calculation shows good correlation with the test results from a number of 10-second runs performed up to 100 RPM. It is clear from Equation (5) and Figure 7 that shear force increases rapidly when the plates approach each other and the gap, s, decreases. The viscous shear forces acting on side-1 and side-2, with corresponding gaps s1 and s2, form a counter-clockwise twisting moment, Mt v, (Equation 8), about the centroid of the base of the tab (Figure 8). As seen in Figure 4, the plate thickness ti, the circumferential angle 2\u03c8 and the radial length (r2-r1) define the tab geometry. Total twisting moment on one tab from the outer periphery up to a radius r is obtained after substituting for \u03bd from Equation (4), and for du/dy in terms of rotational speed \u03c9 and the plate gap s. 2 1 11 20 2 1 1 1( ) m mrm mi t v m B r tM r r s s \u03c8 \u03c8 \u03c1\u03bd \u03c9 dr d\u03c8 \u03b3 + ++ + \u2212 = \u2212 \u222b \u222b& (8) A representation of the tab as a simple rectangular prismatic bar shown in Figure 8 was used for the analysis with the mathematical model. Equation (9) (Young [9]) gives the deflection of a bar subjected to a twisting moment at its free end. t t M L K G \u03b4 = (9) 4 3 4 16where 3.36 1 3 1 ( , and as shown in Fig. 8) t b bK ab a a L a b = \u2212 \u2212 2 The net deflection at the tip is obtained by integrating the effect of the twisting moment over the radial length of the tab. The calculated tab deflection is shown in Figure 9 as a function of the plate gap, s2. This model was crosschecked for accuracy by a Finite Element Analysis (FEA) program using the exact geometry of the inner plate; the agreement was within acceptable limits. Assuming no edge effects, the tabs of the plate in the deflected state, may be modeled as a series of flat slider bearing elements", " Figure 9 shows that the plate tabs experience substantial axial forces, causing the leading edge of the tabs to be pushed on to the outer plates. This total force Pt may be calculated from Equation (11) as the sum of the forces on side-1 and side-2. An equivalent differential pressure ep may be defined in the contact zone on the plate, that will result in the same axial force on the tabs as given by Pt. ( )2 2 2 3 t e Pp r r\u03c0 = \u2212 (12) Let \u03b7 be the coefficient of friction between the plates. The resulting Coulomb frictional torque Tc is 3 3 2 3 2 ( ) 3c eT p r\u03c0\u03b7= \u2212 r (13) As seen in Figure 5 and Figure 8, the friction at the leading edge of the tab creates additional twisting moment t cM on the tab. 1 2 2 i t c P l tM n \u03b7 \u03b4\u2212 + = (14) P1-2 is the differential force between side-1 and side-2 and n is the number of tabs on the inner plate. Once Coulomb friction comes into effect, the plates are irrevocably in the STA mode, even if the twisting moment due to viscous shear torque, Mt v, reduces. Let a typical flow cell consist of the inner plate tabs and the adjacent outer plates. The pressure difference between the bottom side of the lower outer plate (side-4) and the slot space on the upper side of the inner plate (side-1) causes a net axial flow between the adjacent plate cells" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001400_00207170600852000-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001400_00207170600852000-Figure4-1.png", "caption": "Figure 4. Double inverted pendulum with unmodeled dynamics.", "texts": [ "5 for the natural length of the spring, g\u00bc 9.81 being the gravitational acceleration, b\u00bc 0.4 for the distance between the pendulum hinges. The parameters for the two spring-mass-damper systems attached to the pendulums are the following k1\u00bc 20, c1\u00bc 5, m1\u00bc 1.0, k2\u00bc 25, c2\u00bc 4.7, m2\u00bc 1.2. The system\u2019s equilibrium is off the vertical position due to the spring force offset. The objective is to stabilize the pendulums around the vertical equilibrium position. A pictorial representation of the system is provided in figure 4. It can be shown easily that the above system can be put in the normal form as in (1) by differentiating the outputs y1, y2 twice and choosing the states \u00bd 11, 12 >, \u00bd 21, 22 > as the states of the internal dynamics of subsystems 1 and 2. To show that assumption 2.1 is satisfied, we proceed along the same lines as for the coupled Van der Pol example. The equation for the internal dynamics for the first subsystem is given by _ 11 _ 12 \u00bc 12 1 m1 \u00bdk1\u00f0 11 r sin\u00f0y1\u00de\u00de \u00fe c1\u00f0 12 r cos\u00f0y1\u00de _y1\u00de \u00f052\u00de with ~v1 \u00bc vc1 v1 \u00bc \u00bd c11 11, c12 12 > \u00bc \u00bd ~ 11, ~ 12 > and e1 \u00bc yc1 y1 \u00bc \u00bdyc1 y1, _yc1 _y1 > \u00bc \u00bde1, _e1 >, we have, _~ 11 _~ 12 \" # \u00bc 0 1 k1 m1 c1 m1 2 4 3 5 ~ 11 ~ 12 \u00fe 0 k1r m1 \u00bdsin\u00f0yc1\u00de sin\u00f0y1\u00de \u00fe c1r m1 cos\u00f0 _yc1\u00de _yc1 cos\u00f0 _y1\u00de _y1\u00bd 2 4 3 5 \u00f053\u00de 0 5 10 15 20 25 30 35 40 45 50 \u22121" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001672_robot.2005.1570181-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001672_robot.2005.1570181-Figure3-1.png", "caption": "Fig. 3 The actual metallic powder feeder system", "texts": [ " 584 The above materials are mounted on a three-axis working table. . HARDWARE SYSTEM CONFIGURATION The metallic powder feeding system utilizes gravity, fluidization, pressure differential, vibrational energy and an inert carrier gas to deliver powder at consistent and controlled feed rate. The control and monitoring system and the hopper system are incorporated for the system. The control and monitoring system contains several units, controllers, indicators and an air vibrator\u2026etc. The actual metallic powder feeder system is shown in Fig. 3. The system can deliver powder at precisely controlled rates. The system accuracy and consistency provide more reliable and high quality for laser cladding. The powder delivery is generated by local fluidization that caused by the air vibrator. Controllable pressure differential between the hopper content and the carrier gas stream transports the fluidized powder out of the hopper and delivers it onto the substrate though the copper tube. The system adopts the PC-based multi-axis motion controller to control the three-axis motion of the motors" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001405_s0263574705002456-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001405_s0263574705002456-Figure1-1.png", "caption": "Fig. 1. An n-link snake robot model.", "texts": [ " Our discussions are organized around the following steps: Step 1: An n-link model of snake-like robots is constructed. Step 2: An effective form of Kane\u2019s equations of motion is demonstrated and the final equations are derived recursively. Step3 : Lagrange and Newton\u2019s formulations are presented. Step 4: The effectiveness of different approaches is compared in various aspects. http://journals.cambridge.org Downloaded: 19 Oct 2014 IP address: 129.81.226.78 In this research, we use an n-link model as a model of the snake like robot (see Figure 1). The center of mass (CM) of each link is considered to be in the middle of the link. Each link has a passive wheel, which does not have sideslip, so the CM motion of each link is restricted to be parallel to the link direction, which means there is a nonholonomic constraint in each link. In our model, x and y denote the position of the center of mass of the first link, \u03b8i (1 \u2264 i \u2264 n) is orientation of the i-th link and \u03d5i (1 \u2264 i \u2264 n \u2212 1) is the relative angle of the (i + 1)-th link to the i-th link" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001376_iros.2006.282371-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001376_iros.2006.282371-Figure2-1.png", "caption": "Fig. 2. (a) FC grasp: at least the intersection point sij1 lies inside the friction cone defined by Fk,r and F k, ; b) Non-FC grasp: there is not an intersection between the supporting lines of two primitive forces inside the third friction cone.", "texts": [ " For 2D objects and three-finger grasps, the necessary and sufficient condition presented in the following proposition was stated in [23]. Proposition 1: (from [23]) Three contact points pi, pj and pk allow a FC grasp if and only if: (a) the unitary primitive vectors that bound the friction cones at these points positively span the force space, and (b) at least one intersection point between the supporting lines of two primitive forces lies inside the double-side friction cone at the other contact point. Fig. 2a shows an example of three contact points that satisfy the necessary and sufficient condition in Proposition 1, allowing a FC grasp, and Fig. 2b shows an example of three contact points that do not satisfy the condition. From Proposition 1 the following two Lemmas can also be stated (illustrated in Fig. 3). Lemma 1: Consider two contact points pi and pj , and let sijm be the intersection point between the straight lines Fi,c and Fj,c (remember that c\u0302 can be either r\u0302 or \u0302 ). In order to obtain a FC grasp, the third point pk must lie in the intersection of the following two regions on the object boundary: \u2022 The region of points where the unitary vectors that bound the friction cone together with the unitary vectors of the primitive forces that determine sijm , positively span the force space", " The index C can be easily determined just checking a combination of any two contacts from each cell of the Tpn partition. Moreover, since the Tpn partition is always symmetric with respect the straight line i = j, it is only necessary to check a half of the space. The Tpn partition can be used for grasp planning since any combination of two contact points whose C = 0 allows a FC grasp. Let pi and pj be two contact points, C can be geometrically interpreted as the number of intersection points, sijm , between the supporting lines Fi,c, Fj,c that lies inside the friction cone of the third point pk (Fig. 2). Therefore, C can be used as a measure of the robustness of the FC grasp, obtaining the most robust FC grasps when C = 4 (this is the maximum possible value). In the next subsection the Tpn partition (especially the cells with maximum C) is used to obtain independent regions on the object boundary. B. Independent regions The independent regions were defined as regions on the object boundary such that a finger in each region ensures a FC grasp with independence of the exact contact point [7]. These regions are useful to provide robustness to the grasp in front of finger positioning errors, as well as for grasp planning" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000646_006-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000646_006-Figure4-1.png", "caption": "Figure 4. Free-body diagram for m5. T2 is the tension in the cord.", "texts": [], "surrounding_texts": [ "Let us first list the quantities with which we shall calculate. We have the three masses\u2014-m2, m3 and m5. The tension in the lower cord is T1 and that in the upper cord is T2. The force of the ceiling bracket on the upper pulley is P . The magnitudes of the accelerations of the three masses in the inertial laboratory frame are +a2, \u2212a3 and \u2212a5 respectively. (We take upward accelerations as positive and downward accelerations as negative.) The upward acceleration of the lower pulley is +a5. We shall use the symbol g for the acceleration due to gravity. There is a second frame of reference that we must also consider. This is the accelerated, noninertial frame in which the lower pulley is at rest. An observer at rest in this frame would conclude that m2 accelerates upward with an acceleration a and that m3 accelerates downward with an equal and opposite acceleration \u2212a. The magnitudes of these accelerations must be equal in the pulley rest frame; otherwise the cord would snap. The acceleration +a2 of m2 in the inertial frame equals the vector sum of its acceleration in the pulley frame and the acceleration +a5 of the pulley in the inertial frame: +a2 = a + a5 (1) and similarly for +a3, \u2212a3 = \u2212a + a5. (2) Now consider the appropriate free-body diagrams for m2, m3, m5 and the lower pulley (figures 2\u2013 5). From these diagrams we can construct the net force acting on each body and equate it with the product of the mass with the acceleration. For m2 T1 \u2212 m2g = m2a2 = m2(a + a5). (3) Figure 2. Free-body diagram for m2. T1 is the tension in the cord. T1 m2g m2 For m3 T1 \u2212 m3g = m3(\u2212a3) = m3(\u2212a + a5). (4) For m5 T2 \u2212 m5g = m5(\u2212a5). (5) Even though the lower pulley is massless, it still obeys Newton\u2019s second law. Since its mass is zero, the net force acting on it is also zero. T2 \u2212 2T1 = 0. (6) This establishes that T2 = 2T1. (7) Equations (3), (4), (5) and (7) contain four unknowns: \u2212a, a5, T1 and T2. These are readily 290 P H Y S I C S E D U C A T I O N May 2004 solved and yield the results: a = (10/49)g = 2.00 m s\u22122 a5 = g/49 = 0.20 m s\u22122 T1 = [(120/49) kg]g = 24 N, T2 = 2T1 = 48 N. From equations (1) and (2) we find the accelerations a2 and a3 in the inertial laboratory frame as a2 = 2.20 m s\u22122 and a3 = \u22121.80 m s\u22122. At this point we have solved the problem as set. However, the problem has considerably more pedagogic value than just the above solution. The further topics we shall treat include inertial forces, d\u2019Alembert\u2019s principle, the breakdown of Newton\u2019s laws in non-inertial frames, centre of mass motion and the answer to the question of imbalance between equal masses. Inertial forces, d\u2019Alembert\u2019s principle and accelerated frames Rearranging equations (3) and (4) we obtain T1 \u2212 m2g \u2212 m2a5 = m2a (8) and T1 \u2212 m3g \u2212 m3a5 = m3(\u2212a). (9) Although these equations look as though they express Newton\u2019s second law, they do not. The left side of each contains two real forces and a \u2018fictitious\u2019 force. These two fictitious forces, \u2212m2a5 and \u2212m3a5, are essentially reversed effective forces that enable us to describe the motion of the 2 and 3 kg masses in the frame in which the pulley is at rest. We interpret these as inertial forces, forces that arise when we deal with accelerated frames. This artifice, first introduced by d\u2019Alembert in the eighteenth century, is the origin of the centrifugal force. In uniform circular motion we write Fnet = \u2212mv2/r. (10) If, following d\u2019Alembert, we write Fnet + mv2/r = 0 (11) we reduce the problem to one of equilibrium by introducing the fictitious inertial force, +mv2/r , known as the centrifugal force. To see the effect of inertial forces let us return to equations (3) and (4). These apply to figures 2 and 3. The net force is the vector sum of the tension and the weight in each case. It equals the mass times the acceleration in the laboratory frame. However, the acceleration in the laboratory frame may also be written as the vector sum of the acceleration in the non-inertial pulley frame (a for m2 and \u2212a for m3) plus the acceleration a5 of the pulley itself. This is how we arrive at equations (8) and (9). In words we can state equation (8) as follows. The net force on m2 (i.e. T1 \u2212 m2g) plus the inertial force (\u2212m2a5) equals the mass times May 2004 P H Y S I C S E D U C A T I O N 291 the acceleration in the non-inertial frame. It is only in the inertial frame of the laboratory that the net force equals the mass m2 times its acceleration in the laboratory frame a2 (a2 = a + a5). In other words, Newton\u2019s second law does not hold when applied to non-inertial frames. To describe motion in non-inertial frames we must introduce inertial forces. It is unfortunate that d\u2019Alembert\u2019s principle seems to have fallen into neglect in recent years. It is an extremely useful idea in classical mechanics that illuminates the meaning of inertial and noninertial frames. Though we shan\u2019t discuss it here, it is also important in the development of the concept of virtual work. For a fuller exposition of the principle see Goldstein [2] and Sommerfeld [3]." ] }, { "image_filename": "designv11_61_0003668_speedam.2010.5545157-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003668_speedam.2010.5545157-Figure1-1.png", "caption": "Fig. 1 FEM model of test motor", "texts": [ "00 \u00a92010 IEEE SPEEDAM 2010 International Symposium on Power Electronics, Electrical Drives, Automation and Motion phasors and instantaneous symmetrical components analysis and other methods [10, 11 & 12\u202618]. III. FINITE ELEMENT ANALYSIS A proper FEM model of the test motor has been built by using Flux 2D analysis package. In this study, because a two-dimensional analysis software is used, the inductance and resistance values of the end ring of the motor are submitted by an additional model. For achieving this purpose, conventional calculation methods and commercial software are used [19, 20]. The FEM model of the motor is shown in Fig.1. According to these calculations, for example, the inductance of the segment end ring between two rotor bars is 113.367\u00d710-11 H and the resistance is 1.3677\u00d710-6 . In the model, it is considered that the broken bars are off completely. The 2D electromagnetic field computation model of machines in (x, y) Cartesian coordinates is based on the magnetic vector potential formulation characterized by the partial differential equation: s 1 A J j A (1) where A[0, 0, A(x, y)] is the magnetic vector potential [Wb/m], Js[0, 0, Js(x, y)] is the current density [A/m2] in the stator slots, \u03bc is the magnetic permeability and \u03c3 is the electric conductivity" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002542_00029890.2009.11920919-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002542_00029890.2009.11920919-Figure8-1.png", "caption": "Figure 8. (a) Elliptical sector OE0 E determined by trammel inclined at angle \u03b1. (b)\u2013(d) Proof that the area of this sector is ab\u03b1/2.", "texts": [ " As a and b vary, keeping a + b constant, the corresponding ellipses are tangent to the astroid. This property, illustrated in Figure 7b, is described as follows: The envelope of a family of ellipses with a + b constant is an astroid. February 2009] THE TRAMMEL OF ARCHIMEDES 119 A generalization of Theorem 1 to flexible trammels is given in Section 10, leading to a common envelope for a flexible trammel and the curves traced by its points. 6. AREA OF AN ELLIPTICAL SECTOR. The trammel is shown in Figure 8 as a segment AB of length a + b in the first quadrant inclined at an angle \u03b1 with the x axis. The ellipse is traced by point E , where AE has length a and EB has length b. As noted earlier, \u03b1 is the eccentric angle of the ellipse. Let S(\u03b1) denote the area of the shaded elliptical sector OE0 E . Here, E0 denotes the position of E when the trammel is horizontal. Then we have the following simple result which we shall deduce without calculus: Theorem 2. The elliptical sector OE0 E has area S(\u03b1) = 1 2 ab\u03b1. (9) Proof. The elliptical sector can be obtained from a circular sector with central angle \u03b2 and radius b (Figure 8b) by horizontal dilation by the factor a/b (Figure 8c). The circular sector has area b2\u03b2/2, so the dilated sector has area a/b times as much or ab\u03b2/2. But Figure 8d shows that \u03b2 = \u03b1 because both are base angles of congruent right triangles having hypotenuses of equal length b and equal altitudes. Therefore, this simple geometric argument gives (9). When a = b the ellipse is a circle of radius b and (9) gives the area of a circular sector in terms of the eccentric angle, which now equals the central angle. The right member of (9) is linear in \u03b1, so the area of the sector of the ellipse between any two values of \u03b1, say 0 < \u03b11 < \u03b12 \u2264 2\u03c0 , is ab(\u03b12 \u2212 \u03b11)/2. Thus the area of a more general elliptical sector, such as that shown in Figure 9a, is given by: Area of general elliptical sector = 1 2 ab\u03d5, where \u03d5 = \u03b12 \u2212 \u03b11 denotes the angle between the two trammel positions, which is also the change in eccentric angle" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001573_acc.1995.531386-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001573_acc.1995.531386-Figure1-1.png", "caption": "Figure 1. Single Rigid Body with a Centroidal BodyReference Frame", "texts": [ " Thus, if one can ensure the single-body equations of motion satisfy the skew-symmetric property, the generalized equations of motion, by nature, also satisfy the skewsymmetric property. This is the key to satisfying the skewsymmetric property in the recursive Newton-Euler formulation for multi-link manipulators. . 2. The Newton-Euler For~ulat ion For SingleBody Dynamics Consider a single rigid body with a body reference frame XIy'-z' attached to its center of gravity, called a centroidal body reference frame, of which origin is located at r, from the global inertia reference frame, as shown in Figure 1. According to the DAlembert principle's, the virtual work of the single-body system vanishes when the system is in dynamic equilibrium; that is, 6rcT[mkc -mg , -f]+&rT[~,ci,+61J,o-n]=~ (4a) or &:[M,v, + ccv, + g, - z,] = 0 (4b) where 6rc is the virtue displacement of the center of gravity &r = 6AAT is the virtual rotation of the rigid body, and A is the direction cosine transformation matrix [4] &, = [ 21, called virtual displacement (W v, = [ 21 is the generalized centroidal velocity (4g) g, is the gravity acceleration Tilde I-' is a vector cross-product operator ( i b = a x b)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001439_n01-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001439_n01-Figure4-1.png", "caption": "Figure 4. Experimental set-up of the active gearbox vibration control system: (1) test gearbox; (2) slave gearbox; (3) belt driver; (4) clutch and flywheel; (5) torque transducer; (6) speed sensor; (7) piezoelectric stack actuator; (8) adjustable preload nut.", "texts": [ " With the analytical tools described thus far, it is now possible to select components for the experimental active gearbox vibration system described earlier. The following section will discuss the components selected for the particular experimental effort as well as provide a discussion of specific measurement results. In this section, the experimental studies performed to verify the theory described in previous sections are discussed. The test bed for these experiments is the power recirculation gearbox system described earlier. A photograph of the experimental system is provided in figure 4. Note that the actuator support is equipped with a nut such that the preload on the actuator can be easily adjusted. Before the results are analyzed, though, one must first determine the effective stiffness of the structure. To obtain the effective stiffness of the host structure as noted above, the driving point frequency response function is needed. Furthermore, to estimate the required actuation force, the cross-point frequency response function is needed. These two frequency response functions can be determined by performing a standard modal test without the actuator connected to the host structure" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000283_2005-01-1591-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000283_2005-01-1591-Figure2-1.png", "caption": "Figure 2 Yaw plane dynamical model", "texts": [ " Alternatively when the estimated yaw rate is invalid, a different control strategy using feed-forward control replaces the yaw-rate feedback control during the time when the estimated yaw rate is invalid. In the following section, we discuss the ETSC control structure. Section 3 describes various ETSC algorithms. In Section 4, we compare the performance of ETSC to the performance of a two channel ESC under similar vehicle test scenarios. In this section, we present brief descriptions of the ETSC algorithms. Details can be found in the Appendices. The estimation of the vehicle yaw rate can be accomplished using both wheel speeds and vehicle speed information as follows: By referencing to Fig. 2, Vm is the instantaneous velocity of the center of gravity, and Vi, i=1,4, is the forward velocity measured at wheel i. The forward velocity Vi is related to the instantaneous velocity of the center of gravity Vm by the following relationship: 4,1,)1( =\u2212+= itVV i mi \u03c8D (2) where t is half the track of the vehicle and D\u03c8 is the yaw rate of the vehicle. The yaw rate can be determined from equation (2) using the two un-driven wheels 3 and 4, .3 4 \u03c8 \u03c8 tVV tVV m m \u2212= += Therefore t VV 2 34 \u2212=\u03c8 . (3) Assume that un-driven wheels are not braking then 4,3i,rV iii == \u03c9 " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001607_robot.2005.1570128-Figure9-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001607_robot.2005.1570128-Figure9-1.png", "caption": "Fig. 9. Camera and world coordinate frame", "texts": [ " Firstly, the region of rope and load is described by the point of W ri = (W Xi, W Yi, W Zi)T in the world coordinate frame. Then, using the points of rope and load in camera coordinate frame Cri, the points in world coordinate frame is given by W ri = W rC0 + W RC Cri (16) where W RC is rotational matrix from camera coordinate frame to world coordinate frame, and W rC0 is origin of camera in the word coordinate frame. The points of rope and load in camera coordinate frame Cri is then written by Cri = (W RC)T(W ri \u2212W rC0) (17) Figure 9 shows the relationship between world coordinate frame and camera coordinate frame. The screen coordinates (u\u0302i, v\u0302i) corresponding to the camera coordinate frame is ( u\u0302i v\u0302i ) = ( Cxi Czi f Cyi Czi f ) (18) where (Cxi, C yi, C yi)T =C ri and f is focal length of the camera. Using the point (u\u0302i, v\u0302i) in the camera coordinate frame, the pixel value of the screen (ui, vi) is calculated by ui = int ( u\u0302i u\u0302i max ) ui max (19) vi = int ( v\u0302i v\u0302i max ) vi max (20) where ui max, vi max are maximum values of the screen\u2019s pixel along u, v direction, u\u0302i max, v\u0302i max are actual size of the image sensor, and the symbol int(*) means a function of taking floor of (*)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001008_kem.291-292.163-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001008_kem.291-292.163-Figure8-1.png", "caption": "Fig. 8 Grinding results calculated by simulation", "texts": [ " In the case of the constant journal rotation speed method, the undeformed chip size varies and it is larger near C=0 deg. On the centrally, the chip size is controlled to maintain constant value in the controlled speed ratio grinding method. From the previous figure, in the controlled speed ratio grinding method, the grinding speed ratio is smaller than that in the constant journal rotation method, so the journal rotational speed is set to slower. Therefore, the chip thickness is thinner than in the constant journal rotational method, and the chip size is controlled. Figure 8 shows the grinding results obtained by the simulation. The results are calculated under deferent control method of the journal rotation, one is the controlled speed ratio grinding method (below), and the other is the constant journal rotation method (upper). The normal grinding force ps (left), the surface roughness (Ry)s (center) and the total amount of the residual stock d0+d1+dr (right) are shown respectively, and also illustrated around the cross section of a pin . Where, d0 is a wheel wear, dr is a residual stock due to the elastic and plastic deformation of a grain" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001958_s0022-0728(69)80332-3-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001958_s0022-0728(69)80332-3-Figure3-1.png", "caption": "Fig. 3. Dependence of limiting current on pH (curve (a)) shown together with the composit ion of the soln. in terms of the percentages of the various dissociated forms of EGTA (curves as indicated), at different pH-values. Data from Table 1.", "texts": [ " The assumption made in the derivation is that the complex HgZ 2- is the predominant species produced by the electrode reaction and the equations will therefore fail if other species such as hydrogen or hydroxo complexes are formed in appreciable proportions at low and high pH-values. Table 1 summarises the results obtained in a study of the dependence of the EGTA wave on pH in the pH-range, 2.3-13.0. Constant ionic strength buffers, maintained at an ionic strength of 0.5, were used and potentials were measured using a three-electrode system as described in the experimental section. The small variation of the limiting currents with pH is presumably due to the slight kinetic nature of the wave, as described earlier. Figure 3 shows this variation together with the composition of the solution in terms of the percentages of various dissociated forms of EGTA at different pH-values. The stability constant of the mercury-EGTA chelate, HgZ 2 , had so far only been determined potentiometrically and the values of 10 z32\u00b0 at 20 \u00b0 and /~=0.1 J. Electroanal. Chem., 21 (1969) 541-546 obtained by Schwarzenbach et aI.13; 1023.8 at 25 \u00b0 and p=0.1 obtained by Reilley and coworkers 14'4, and 1023\"12 at 20 \u00b0 and #=0.1 obtained by Mackey et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001202_135065005x33900-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001202_135065005x33900-Figure1-1.png", "caption": "Fig. 1 Geometry of a hydrodynamic journal bearing groove", "texts": [ " The solution of the equations involved in the THD lubrication problem is obtained by the finite element Corresponding author: Universite\u0301 Hassan II \u2013 Mohammedia, Faculte\u0301 des Sciences et Techniques, BP 146, 20650 Mohammedia, Morocco. J02804 # IMechE 2005 Proc. IMechE Vol. 219 Part J: J. Engineering Tribology at NORTH CAROLINA STATE UNIV on April 18, 2015pij.sagepub.comDownloaded from method. This model allows to determine the effects of the feeding pressure and the runner velocity on the THD behaviour of the lubricant in the groove of HD journal bearing and to emphasize the dominant phenomena in the feeding process. Figure 1 shows the geometry of an axial groove of an HD journal bearing. An incompressible, viscous fluid fills the space between the stationary bearing and the rotating journal. The bearing has a radius Rb and length Lb, while the journal rotates with a surface velocity Uj. The circumferential width is supposed to be small versus shaft radius and the surface curvature effects are not considered. Moreover, the geometry introduces the assumption that the groove axial length is much larger than its circumferential width or depth" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003612_s11044-010-9190-2-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003612_s11044-010-9190-2-Figure1-1.png", "caption": "Fig. 1 Soft mounted electrical machine (here, 2-pole induction motor) with sleeve bearings", "texts": [ "eywords Rotordynamics \u00b7 Sleeve bearing \u00b7 Rotor eccentricity \u00b7 Foundation \u00b7 Electrical machine Electrical machines are often used in the industry as drive applications in different plants. Because of plant specific requirements, the foundation of the electrical machine is often only designable as a soft foundation, e.g. as a steel frame foundation (Fig. 1), or by using soft rubber elements under the machine feet. Large electrical machines with power ratings \u22651 MW, operating at high speed (n \u2265 3000 rpm), are often designed with sleeve bearings because of the high circumferential velocity of the shaft journals. Additionally, a flexible shaft design is often used, so that the U. Werner ( ) Industry, Drive Technologies, Large Drives, Industry Development, Siemens AG, Vogelweiherstra\u00dfe 1-15, 90441 Nuremberg, Germany e-mail: werner.ulrich@siemens.com motor is operated above the first bending critical speed of the rotor", " as\u2212w,\u03ba = r\u0302+ s\u2212w,\u03ba + r\u0302\u2212 s\u2212w,\u03ba; bs\u2212w,\u03ba = \u2223\u2223r\u0302+ s\u2212w,\u03ba \u2212 r\u0302\u2212 s\u2212w,\u03ba \u2223\u2223; \u03c8s\u2212w,\u03ba = ( \u03b1+ s\u2212w,\u03ba +\u03b1\u2212 s\u2212w,\u03ba )/ 2 (44) Again, the solutions can be superposed for all three kinds of dynamic eccentricity: r s\u2212w = \u2211 \u03ba=u,m,a r+ s\u2212w,\u03ba \u00b7 ej (\u03a9\u00b7t+\u03d5\u03ba ) + r\u2212 s\u2212w,\u03ba \u00b7 e\u2212j (\u03a9\u00b7t+\u03d5\u03ba ) (45) The same derivation can be applied to deduce the mathematical description of the relative orbit between the bearing housing centre B and the shaft journal point V . Only the indexes have to be changed in the formulas (35)\u2013(45) according to (46). s \u2212 w \u2192 b \u2212 v; s \u2192 b; w \u2192 v (46) In this section, the orbital movements of a soft mounted 2-pole induction motor (Fig. 1; Table 1) are analyzed exemplarily, caused by the different kinds of dynamic eccentricity: \u2013 rotor mass eccentricity \u00eau \u2013 magnetic eccentricity \u00eam \u2013 bent rotor deflection \u00e2. The absolute orbits of the centre of stator mass S, the shaft centre point W , the centre of shaft journal V, and the centre of bearing housing B , and the relative orbits between the centre of stator mass S, and shaft centre point W, and between the centre of shaft journal V, and the centre of bearing housing B are calculated" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000973_978-3-540-44415-2_15-Figure19-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000973_978-3-540-44415-2_15-Figure19-1.png", "caption": "Fig. 19. Max kicking range and lowest hip height", "texts": [ "18, different kicking directions and different kicking points on the ball may result in different kicking effects. Kicking in a horizontal direction is not the optimal way. But accurate kicking requires accurate information on both the ball and robot localization. To simplify the problem, we only consider kicking in horizontal direction in this paper. 3.1.2 Max Kicking Range and Effective Kicking Range To achieve a longer kicking range, the humanoid robot\u2019s hip should be kept as low as possible, because at the max kicking range point, the kicking leg stretches straight (Fig. 19). The lower the Hip Height (HH) is, the wider the max kicking range is. But the HH is subjected to the constraint that the CG should lie in the supporting foot, because if HH is lowered, the CG moves forward. So, there exists a threshold value for the Lowest Hip Height (LHH). Based on the above analysis, the kicking leg speed is zero at the maximum kicking range point. If kicking the ball at that point, there is no momentum for the ball. We need to define another range, effective kicking range (see Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001154_sice.2006.315039-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001154_sice.2006.315039-Figure6-1.png", "caption": "Fig. 6 Wheeled vehicle", "texts": [ " -r0 Upper bound1 (11) -18 -18 MPC methods typically determine control input based on finite horizon open-loop control optimization problems. Our optimization problem at k to determine vx[k] in (10) is N-1 min Zv: T=O subject to (T+l := AclT + BlvT, ( <- (T <- (, u < -K,4T (12) (o := (x[kJ13) + VT K, (14) 3. MULTI-VEHICLE FORMATION 3.1 Problem formulation We consider a group of n unicycles indexed by z= , n: \u00b1i= vi cos Oi, Yi = vi sin Oi, Oi = wu, (19) where vi and wi are the linear and angular velocities of the vehicle i respectively, and (xi, yi, Oi) denotes the measurable coordinate with respect to a global frame (see Fig. 6). -18 where u,a, 77 are the upper and lower bounds given in (7)-(8) and (3). The first element v0 of the optimal solution is applied at each time step k. While traditional MPC methods solve online the optimization problem as described above, recent methods[11][12] can solve offline the problem above as the following piecewise affine feedback low: vx[k] Fr(x[k] + Gr, iJf x[k] C Pr (15) Pr = { CR2 Hrf < Er}, r = 1, ..., Np, (16) where Np is the number of the polytope regions. In order to take account of disturbances and modeling errors, we adopt a robust MPC approach[4], which deals with the additive uncertainty wx[T] C W as (X[T + 1] = Ac1ix[T] + Blvx[T] + WX [T]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001376_iros.2006.282371-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001376_iros.2006.282371-Figure1-1.png", "caption": "Fig. 1. Friction contact, where fi is the applied force, fn i and ft i are its normal and tangent components with directions ni and ti, respectively, ri and i are the unitary primitive vectors, and Fi,r , F i,l , F i,l and Fi,t are the supporting lines of the these force components.", "texts": [ " Given Bd and the inward normal direction \u03b8i at each point qi, it is assumed that if pj \u2208 [qi, qi+1] then \u03b8j \u2208 [\u03b8i, \u03b8i+1] (the obtention of Bd and \u03b8i is outside of the scope of this paper, and it can be done as in [20]). Let f i be the force exerted by a finger on the object, either on a point pi or qi, and let fn i and f t i be its components normal and tangent to the object boundary whose directions are given by the unitary vectors n\u0302i = (cos \u03b8i sin \u03b8i)T and \u0302ti = (\u2212 sin \u03b8i cos \u03b8i)T , respectively (Fig. 1). Based on the Coulomb model of friction, the finger slippage on the object boundary is avoided if: \u03bc\u2016fn i \u2016 \u2265 \u2016f t i\u2016 (1) being \u03bc the friction coefficient. Geometrically, equation (1) constraints the force applied by the finger to lie inside a friction cone centered on the direction normal to the object boundary and limited by the called primitive forces, f r i and f i , whose directions are given by the following unitary primitive vectors: r\u0302i = [cos(\u03b8i \u2212 \u03d5) sin(\u03b8i \u2212 \u03d5)] (2) \u0302 i = [cos(\u03b8i + \u03d5) sin(\u03b8i + \u03d5)] (3) with \u03d5 = arctan\u03bc" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000251_s1474-6670(17)63640-1-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000251_s1474-6670(17)63640-1-Figure1-1.png", "caption": "Fig. 1 F4-E aircraft with additional hori zontal canards.", "texts": [], "surrounding_texts": [ "Keywords. Aircraft control; Pole region assignment; Robust control; Parameter space method.\nINTRODUCTION\nRedundancy management in control systems is usually viewed separately from the control algorithm. The control system is designed under the assumption, that sensors do not fail. Then redundancy management has to pro vide the required measurements with only very short interruptions by failures of in dividual sensors. If the plant is for exam ple an unstable aircraft, this means that failure detection is vital for stabilization, it has to operate fast and this requirement is in conflict with the requirement of low probability of false alarms.\nIn this paper a hierarchical concept is pro posed. Its basic level is a fixed gain con trol system, which is designed such, that pole region requirements are robust with respect to changing flight conditions and component failures. All more sophisticated tasks like failure detection and redundancy management, plant parameter identification and controller parameter adaptation or gain scheduling are assigned to higher levels, if they are required for best performance. The higher levels process more information and are operating in a slower time scale than the basic level. Since the higher levels are not vital for stabilization they can make their decisions without panic haste.\nThis paper deals with the design of the ro bust basic level control system. The parti cular example is an F4-E aircraft, which is destabilized by horizontal canards, see\norder low pass with transfer function\n14/(s + 14), its state variable is 0 , the e deviation of the elevator deflection from its trim position. 0 ~s not fed back, be e cause this would require an estimate of the trim position.\nIn a previous study Franklin (1980, 1981) assumed measurement of normal acceleration Nz and pitch rate q and the linearized state equations were written in sensor coordinates with the state vector xT [N q 0 ]. z e Thus X A x + b u", "1178 J. Ackennann\nData for four typical flight conditions were taken from Berger, Hess and Anderson (1973) and are given in the appendix. The eigenvalue locations of the short period mode are given Ln Table I.\nThe aircraft is unstable in subsonic flight and unsufficiently damped in supersonic flight, such that adequate handling proper ties must be provided by the control system. Note that in stationary flight the elevator and canard are not used independently. The commanded deflections are coupled as 6ecom = u, 6ccom = -0.7u, where the factor -0.7 was chosen for minimum drag. Thus the short period mode stabilization is a single input problem.\nThe required closed loop eigenvalue locations are given by military specifications for flying qualities of piloted airplanes (1969). For the short period mode described by\ns2 + 2s W s + w2 0 (2) sp sp sp\nthe restricted range of damping s and natural frequency W is sp\nsp\n0.35 ~ s sp ~ 1.3 (3)\nwhere wa and wb depend on the flight condi\ntion and are given in the appendix for the four conditions considered here.\nFig. 2 shows the nominal region r., eq. (3) J together with the open loop eigenvalues for a subsonic flight condition j. Damping greater than one in eq.(3) corresponds to two real eigenvalues. Eq.(3) would admit some real pairs of poles with one of them outside the region r .. In the following no use is made\nJ of this possibility. For all real pairs in side r . condition (3) is satisfied. We reJ quire, that the closed loop short period poles of each flight condition j = 1, 2, 3, 4 are loca ted in the respective region r ..\nJ\nThe military specifications do not contain requirements for the location of additional closed loop poles originating from actuator or feedback dynamics. Quick response is essen tial for a fighter, therefore the non short\nperiod eigenvalues should not unnecessaril y slow the dynamic response. In order to keep them fast enough and separate from short pe riod eigenvalues an additional regi on t o the left of r. is prescribed. The dampin g reJ quirement s ~ 0.35 is kept from eq.(3) and a natural frequency range wb ~ W ~ wd ' wd = 70 rad/sec is chosen in order Eo main-\nThe assumed type of sensor failure is that the nominal gain v = 1 is reduced to some value 0 ~ v < 1. As far as eigenvalue loca tion is concerned, only this multiplicative error is important. There may be an additive bias or noise tenn, which should be removed by a failure detection system at a higher hierarchical level.\nThe objective of this paper is to design the basic level control system such that the pole region requirements of Fig. 2 are robust with respect to changing flight conditions and sensor failures. This is an example for the application of a novel parameter space design technique and generalized D-decomposition, see Ackermann and Kaesbauer (1980, 1981). It will be reviewed briefly in the following paragraph. In application to the example it is then shown, how robustness with respect to changing flight conditions can be achieved by appropriate choice of kN and k in an z q output feedback control law\nu - [k Nz k q 0] ~ (4)\nFor robustness with respect to sensor fai lures Franklin (1980, 1981) studied a con figuration with two gyros and one accelero meter and dynamic feedback. It showed the disadvantage of using the accelerometer. Therefore here a different solution with three gyros and dynamic feedback is given. For this solution the responses in C:: for a pilot step input are given, where\n(N + 12.43q)/C z '\"\n(5 )" ] }, { "image_filename": "designv11_61_0002109_s11071-007-9228-z-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002109_s11071-007-9228-z-Figure3-1.png", "caption": "Fig. 3 (a) Symmetric rotor. (b) Aerodynamic three tilting pads bearing", "texts": [ " An algorithm of calculation of laminar gas flow in bearing in Fortran is based on the work of Lund [8, 9]. The used program for calculation of dynamic characteristics of bearings at different revolutions takes into account inertia properties of tilting pads. The vertical load from the weight was Fst = 38 N. This program gives discrete values of elements of full stiffness and damping matrices [3, 5]. Successful running of prototype [1] proves the correctness of calculation by the program developed for design of aerodynamic bearings. A photo of a tilting pad bearing is included in Figure 3b. In spite of that, the method of calculation was elaborated for constant revolutions, the experiments show that bearing properties do not change essentially with the slow increase or decrease of rotor angular velocity. The TECHLAB description of linear and frequencydependent stiffness and damping properties was applied for solution of rotor motion. Discrete values must be replaced for numerical solution by continuous functions of angular velocities \u03c9 (s\u22121). These values vary relatively very strong at different revolutions", " The substitutive functions for the following solution must be therefore selected as a combination of monotone polynomials c + d\u03c9 or c + d\u03c9 + f \u03c92 and of functions describing real or imaginary components of 1-DOF system response: K = c + d\u03c9 + a 2 \u2212 \u03c92 ( 2 \u2212 \u03c92)2 + b2\u03c92 , K = c + d\u03c9 + f \u03c92 + a b\u03c9 ( 2 \u2212 \u03c92)2 + b2\u03c92 . (1) Good agreement of analytical continuous functions with discrete points (circles) can be reached by appropriate selection of parameters c, d, f, 2, a, b. The functions of stiffness Kxx (\u03c9) and damping Bxx (\u03c9) are shown in Figs. 1 and 2 as examples. The experimental rotor (Fig. 3a) is symmetric, with total mass m = 7.6 kg. Inertia moment to the y- or xaxis is I = 0.10024 kg m2. The distance between the centers of bearings is l = 0.32 m. The inertia properties defined by mass m and moment of inertia I can also be replaced by effects of three masses [2], two of them m1, m2 situated in the centers of bearings and the third mass m3 in the center of gravity T. The centrifugal force me\u03c92 acts at the distance a to the right from the center T. The rotor is supported on two identical three-pad aerodynamic bearings (diameter d = 50 mm, clearance \u03b4 = 0.05 mm), which, at sufficiently high revolutions, do not need any supply of air pressure, as the surrounding air is drawn into bearing and forms a load-bearing lubricated film. The simple structure of aerodynamic bearing is shown in Fig. 3b. Springer Fig. 1 Stiffness Kxx versus frequency \u03c9 Fig. 2 Damping Bxx versus frequency \u03c9 Matrices K and B are used for calculations of responses of rotor at unbalance excitation, ascertained by the following equations: (m1 + m3/4)x\u03081 + m3/4x\u03082 + Kxx (\u03c9)x1 + Bxx (\u03c9)x\u03071 + Kxy(\u03c9)y1+Bxy(\u03c9)y\u03071 = (1/2\u2212a/ l)me\u03c92 cos \u03c9t, Springer (m2 + m3/4)x\u03082 + m3/4x\u03081 + Kxx (\u03c9)x2 + Bxx (\u03c9)x\u03072 + Kxy(\u03c9)y2 + Bxy(\u03c9)y\u03072 = (1/2+a/ l)me\u03c92 cos \u03c9t, (m1 + m3/4)y\u03081 + m3/4y\u03082 + Kyy(\u03c9)y1 + Byy(\u03c9)y\u03071 + Kyx (\u03c9)x1 + Byx (\u03c9)x\u03071 = (1/2 \u2212 a/ l)me\u03c92 sin \u03c9t (m2 + m3/4)y\u03082 + m3/4y\u03081 + Kyy(\u03c9)y2 + Byy(\u03c9)y\u03072 + Kyx (\u03c9)x2+Byx (\u03c9)x\u03072 = (1/2 + a/ l)me\u03c92 sin \u03c9t, (2) where x1 and x2 are measured from the equilibrium positions, given by constant load mg" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002320_978-3-540-77457-0_2-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002320_978-3-540-77457-0_2-Figure8-1.png", "caption": "Fig. 8. Simulation of posture control of grasped object", "texts": [ " (7) Let mobj be the mass of the object and Iobj be the moment of inertia of the object around its center of gravity. Let Ifinger be the moment of inertia of the finger around its rotational joint. Assuming that mass transfer due to the deformation of each fingertip is negligible, the kinetic energy of the system can then be formulated as T = 1 2 mobj(x\u03072 obj + y\u03072 obj) + 1 2 Iobj\u03b8\u0307 2 obj + 1 2 Ifinger\u03b8\u0307 2 l + 1 2 Ifinger\u03b8\u0307 2 r . (8) From eqs.(7) and (8), we can formulate the Lagrange equations of motion of a pair of fingers pinching a rigid object. We then introduce viscosity terms to equations. Figure 8 shows a simulation of posture control of an object pinched by a pair of fingers. We used the identified Young\u2019s modules in the simulation. Figure 8-(a) shows the contact between the fingers and the object without fingertip deformation. Figure 8-(b) shows the initial grasping, where both fingertips have the same deformation. Both fingers rotate counterclockwise in Figure 8-(c), and the object rotates clockwise. Both fingers rotate clockwise in Figure 8-(d), and the object rotates counterclockwise. The simulation result based on the parallel distributed model agrees with the observation shown in Figure 2. Figure 9 shows a comparison between simulation and experimental results in posture control of a pinched object. Figure 9-(a) shows the relationship between orientation angle \u03b8obj and coordinate xobj of a pinched object. As shown in the figure, both results agree with each other, but the difference between the results increases as the orientation angle becomes larger" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003868_sii.2013.6776699-Figure9-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003868_sii.2013.6776699-Figure9-1.png", "caption": "Fig. 9. Kinematic of the mobile robot.", "texts": [ " The Mamdani fuzzy model was applied for the proposed navigation control methodology. Two-side Gaussian membership function were used to fuzzify the sensor inputs and outputs control variables to mobile robot as \ud835\udf07\ud835\udc34(\ud835\udc65) = { \ud835\udc52\ud835\udc65\ud835\udc5d [\u2212 1 2 ( \ud835\udc65\u2212\ud835\udf0e1 \ud835\udf0e1 ) 2 ] \ud835\udc65 \u2264 \ud835\udc501 1 \ud835\udc501 \u2264 \ud835\udc65 \u2264 \ud835\udc502 \ud835\udc52\ud835\udc65\ud835\udc5d [\u2212 1 2 ( \ud835\udc65\u2212\ud835\udf0e2 \ud835\udf0e2 ) 2 ] \ud835\udc502 \u2264 \ud835\udc65 (4) Where \u03c31, c1, \u03c32 and c2 is the width and the central of the Gaussian for the left and right side respectively. The output to the mobile robot is the linear velocity and turning angle . The inputs from sensors are organized as follows and shown in Figure 9. For maze scenario, three input membership functions as shown in Figure 10 are used as 1) \ud835\udc6c\ud835\udc95\ud835\udc82\ud835\udc93\ud835\udc88\ud835\udc86\ud835\udc95 \ud835\udf3d (target direction) in the universe of disclosure [-90, 90] contains 3 fuzzy sets {L, F, R} denote left, front and right respectively, 2) \ud835\udc6c\ud835\udc90\ud835\udc83\ud835\udc94\ud835\udc95\ud835\udc82\ud835\udc84\ud835\udc8d\ud835\udc86 \ud835\udf3d (obstacle direction) in the universe of disclosure [-90, 90] contains 7 fuzzy sets {L, SL, ML, Z, SR, MR, R} denote left, small left, medium left, zero, small right, medium right and right respectively and 3) \ud835\udc6c\ud835\udc90\ud835\udc83\ud835\udc94\ud835\udc95\ud835\udc82\ud835\udc84\ud835\udc8d\ud835\udc86 \ud835\udc85 (distance to obstacle) in the universe of disclosure [0 mm, 600 mm] contains 3 fuzzy sets {Z, N, F} denote zero, near and far respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001583_icma.2006.257688-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001583_icma.2006.257688-Figure1-1.png", "caption": "Fig. 1 (a) Rhomb grid, (b) a partition example of virtual rhomb grid for sensor field, (c) sensor deployment based on virtual rhomb grids Fig. 2 -redundant rhomb grid The above is ideal and theoretical, that is, the performance", "texts": [ "1 a sensor\u2019s radio ability is omnidirectional, that is, its coverage range is a disk whose radius is r and whose area is D (D= r2). Hyp.2 in a sensor field, all sensors\u2019 radio power is 1-4244-0466-5/06/$20.00 \u00a92006 IEEE uniform, that is, the radio radius r of all sensors is equal. Hyp.3 in a sensor field, all sensors are in the same plane. Hyp.4 the initial deployment is random. Hyp.5 every node has the ability to know its own location by some method such as GPS [19] or other methods [20]-[24]. Hyp.6 each node can be mobile. To illustrate in Figure 1 (a), each sensor associates itself with one of the vertices of a rhomb grid. The rhomb grid both makes enough use of sensing and communication range and ensures full coverage and full connectivity of sensor field [25]. Furthermore, the minimum number of the nodes in sensor field covered fully and seamlessly is determined by the equation [25] (1) 23 3 2 2 2= = (1) 3 3rr F F FN \u03b4 = Where F is the area of the sensor field 3 3 2 2 r\u03b4 is the effective area of each sensor, the r is the radius of sensing or communication of sensors. Note that the MSDVRG algorithm assumes that the number of the sensors n is not less than N. The condition is required not by the MSDVRG algorithm but by no \u201choles\u201d (full and seamless coverage of sensing and communication) in the sensor field. The sensor field is partitioned into \u201cvirtual rhomb grids\u201d (illustration in Figure 1 (b)). Figure 1 (c) shows the form of sensor deployment based on the virtual rhomb grids. In the form of sensor deployment, each sensor locates at one of the vertices of virtual rhomb grid. In this way, there are no \u201choles\u201d in the sensor field, namely, full and seamless coverage. Furthermore, the sensor number of the sensor deployment is minimum. parameters of all the sensors are the same and all the sensors must be in the same plain. But the practical situations is not often so. Therefore, the paper presents -redundant Movement-assisted Sensor Deployment based on Virtual Rhomb Grid ( MSDVRG). III. -REDUNDANT MOVEMENT-assisted SENSOR DEPLOYMENT based on VIRTUAL RHOMB GRID ( MSDVRG) To illustrate the same as Figure 2, lines are marked out at /2 from the vertices of the original rhomb. In this way, the new Rhomb ABCD is shaped in the original rhomb, and its edge length is 3r \u03b5\u2212 . The sensor field is partitioned into the new rhomb in the form of the same illustration as Figure 1 (b). Thus the form has the redundancy compared to the form of the original rhomb grids. The paper names the algorithm MSDVRG for this. Figure 3 shows the data structure of the MSDVRG algorithm, and Figure 4 shows the implementation details in pseudocode form. For a given number of sensors, MSDVRG attempts to move each sensor to the desired vertex of the virtual rhomb grids. The desired vertex is closest from the sensor. The cluster head or base station is responsible for executing the MSDVRG algorithm and managing the one-time movement of sensors to the desired locations" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003730_icias.2012.6306067-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003730_icias.2012.6306067-Figure5-1.png", "caption": "Figure 5. Actual robot used for the expe", "texts": [ " An additional PID controller is added into the system to compensate for position difference. The controller is inserted in between the two controllers like shown in Fig. 4. This type of compensation is coined in [10] where only an integrator was used. But for better control, a full PID is used for this purpose. IV. EXPERIMENTAL RESULTS For the experiment, a robot was fabricated from aluminum bars, direct current motors and quadrature encoders. The total weight of the robot is 8.6 kg with length of 0.5m and wheels distance of 0.7m. The robot is shown in Fig. 5. The robot will move on vinyl flooring as required by the ABU ROBOCON contest. The cascade controller is implemented in a microcontroller using the UTM team\u2019s motor control board where the cascade control, encoder interface and motor actuation are programmed in a dsPIC33FJ128MC802 microcontroller. The feedback was obtained from quadrature encoders attached to the ground and the actuators are 41 Watt direct current motors. [ 504 ] riment. graph graph Firstly, the inner loop of the velocity controller, was tuned" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001672_robot.2005.1570181-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001672_robot.2005.1570181-Figure7-1.png", "caption": "Fig. 7 The relation to various parameters for the proposed system", "texts": [ " OPERATION PARAMETERS OF THE RT PROCESSES The physical properties of nickel-based alloy we adopted are good for the proposed RT system. Its considerably lower melting point, 940 , is suitable for lower laser power and the hardness of Rockwell C 58-63 is applicable to create a mold. The appropriate powder selection, there are many parameters need to be considered, such as laser power, traverse speed, layer thickness, spot size of laser, feed rate and offset distance of tool paths [10]. Several process parameters of the proposed RT system are shown in Fig. 7. \u2206 \u03c7 was defined to be the distance between the center of the spot size and the intersection point of the axis of powder stream and substrate surface. \u2206 \u03c7 was 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. The height and width of the single clad is very important for the effect of RT performance. Therefore we make many single clads by the RT system to find out the suitable process parameters (see Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001607_robot.2005.1570128-Figure10-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001607_robot.2005.1570128-Figure10-1.png", "caption": "Fig. 10. Experimental gantry crane", "texts": [ " The control method is the same as presented in Section 4, where the desired trajectory is modified (smoothed) with a lag system, then the rope length and trolley position trajectories are calculated using the inverse dynamics calculation. Finally the trajectories of rope length and trolley position are used to control the suspended load using an appropriate servo controller for trolley and pulley. VI. AUTOMATIC STOP EXPERIMENT Using the proposed control method based on the inverse dynamics and emergency stop control method in real time, a crane work is experimented considering two cases that no obstacle in the work space and man is entering in the work space. The size of experimental crane is shown in Fig.10. In the first experiment, the desired trajectory of the suspended load is given by control stick as desired velocity. The suspended load is moved in x \u2212 y plane. The actual path of the load for this case is shown in Fig.11. By this experiment, we confirm that the suspended load is controlled without sway for the case of no obstacles using the proposed control method. In the figure, pd = (xd, yd) is desired trajectory by the control stick operation, p\u0302d = (x\u0302d, y\u0302d) is the modified trajectory using a lag system, and p = (x, y) is actually measured trajectory of the suspended load" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003094_iros.2009.5354487-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003094_iros.2009.5354487-Figure3-1.png", "caption": "Fig. 3. Model of planar fully-actuated compass-like biped robot.", "texts": [ " This result strongly depends on the special property of the simplest walking model. The value of \u03b5 is determined only by the hip angle at impact regardless of the swing-leg\u2019s angular velocity just before impact in this model. In other words, SLR does not affect the impact dynamics of the simplest walking model. In the following sections, we will analyze the energy-loss coefficient in general compass-like biped models whose leg mass cannot be neglected. Especially, the effects of SLR and mass-distribution are investigated. Fig. 3 shows the model of a planar, fully-actuated, compass-like biped robot with flat feet. Two joint torques, u1 and u2, can be exerted at the ankle joint and hip joint. Let \u03b8 = [ \u03b81 \u03b82 ]T be the generalized coordinate vector, where \u03b81 and \u03b82 are the angular positions of the stance and swing legs with respect to vertical. The dynamic equation then becomes M(\u03b8)\u03b8\u0308 + C(\u03b8, \u03b8\u0307)\u03b8\u0307 + g(\u03b8) = Su = [ 1 1 0 \u22121 ] [ u1 u2 ] . (3) These matrices are described in detail elsewhere [2]. If we assume inelastic collisions for the stance-leg exchange and set suitable values for the physical parameters, the robot can exhibit passive dynamic walking on a gentle slope" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001618_robot.2005.1570438-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001618_robot.2005.1570438-Figure2-1.png", "caption": "Fig. 2. Analytical model of a two-link planar casting manipulator. The first link is rigid and actuated, whereas the second is flexible (like a string) and presents no actuation. A gripper is used to catch the target object and a braking mechanism is employed to control its ballistic flight.", "texts": [ " The analytical model of a two-link casting manipulator is obtained in Section II. Section III describes the vision system as far as concerned with the algorithms used to extract the target position and to calibrate the camera. Successive Section IV presents the control algorithm, and Section V describes the experimental setup and then shows the results of the experiments. Finally, the closing Section VI summarizes the work achievements and discusses about possible future developments. 0-7803-8914-X/05/$20.00 \u00a92005 IEEE. 2191 Consider the casting manipulator depicted in figure 2 which consists of a two-link planar manipulator. The first link L1 is rigid, whereas the second one, L2, is composed of a flexible material (like a string) and a gripper mounted at the end of it. Both joints are rotational, but the first one is actuated by the torque \u03c4 , whilst no actuation is present at the second one. The robot is then provided with a braking mechanism, preventing the second link to unroll. As far as concerned with the nomenclature, let I1 and m1 be the inertia and the mass of link L1, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003748_1.4025208-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003748_1.4025208-Figure6-1.png", "caption": "Fig. 6 Solid model of the engine assembly", "texts": [ " The nonlinear stiffness behavior is approximated by a third order polynomial function [9] Fi \u00bc X j Aij s3 i;j \u00fe X j Bij si;j (1) 2.4 Cylinder Bore Distortion Analysis. This section describes a finite element method analysis, which was used to predict a cylinder block bore distortion. The model includes a cylinder head, valve seat inserts, a head gasket, a cylinder block, and head bolts. In addition, a cam carrier, cam carrier bolts, injectors, and injector clamps are added to the model in order to obtain realistic results as shown in Fig. 6. Soft springs are used to facilitate the convergence. The analysis simulates assembly of these components considering their interactions. Bore distortion is postprocessed based on displacement results obtained from the analysis. Cylinder bore distortion results consist of roundness, cylindricity, and Fourier coefficients calculated at different depth of cylinder bores. These results are used to solve problems related to piston slap, oil consumption, and piston blow-by. Bore distortion analysis is a nonlinear static analysis solved in ABAQUS" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002197_6.2007-2727-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002197_6.2007-2727-Figure5-1.png", "caption": "Figure 5. Force Subjected to Aircraft during Turning", "texts": [ " ( ) ( ) ( ) ( ) 0 ( ) ( ) ( ) 0 ( ) t t t t l t t t t u u d X X d Y Y Y Y d v v d Y Y d X X X X d \u03b8 \u2212 \u2212 \u2212 \u0394\u23a7\u2212 \u2212 \u2260\u23aa \u2212\u23aa= \u23a8 \u2212 \u2212 \u2212 \u0394\u23aa \u2212 \u2260 \u23aa \u2212\u23a9 & (15) where ( )( ) ( )( )t t t tX X u u Y Y v vd d \u2212 \u2212 + \u2212 \u2212 \u0394 = (16) From Eqs. (12) and (15), the required lateral acceleration yca can be calculated using airspeed aV . ( ) ( ) ( ) ( ) 0 ( ) ( ) ( ) 0 ( ) yc a v t t a pn t t t t a pn t t a V u u d X X dV k Y Y Y Y d v v d Y Y dV k X X X X d \u03c8= \u2212 \u2212 \u2212 \u0394\u23a7\u2212 \u2212 \u2260\u23aa \u2212\u23aa= \u23a8 \u2212 \u2212 \u2212 \u0394\u23aa \u2212 \u2260 \u23aa \u2212\u23a9 & (17) The required rolling angle command to achieve the lateral acceleration in Eq. (17) is calculated from the equilibrium of force. Fig.5 shows the force subjected to the UAV during steady turning at an angular rate of v\u03c8& . The required lateral acceleration yca is now broken down to the rolling angle command c\u03c6 . 1tan yc c a g \u03c6 \u2212= (18) The rolling angle command calculated in Eq. (18) will be given to the attitude control loop. b.) Longitudinal Control For longitudinal control, a simple altitude hold system is used. The block diagram of this system is shown in Fig.6. The pitch angle command is calculated to retain the desired altitude" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002384_memsys.2009.4805354-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002384_memsys.2009.4805354-Figure1-1.png", "caption": "Figure 1: Schematic of the wireless implantable biochip system.", "texts": [ " The flexible polymer tube lab-chip integrated with microsensors has been proposed and applied for developing smart microcatheter [8]. The main problems in the biosensor chip fabrication include: (1) the incompatibility between the biomaterials and the MEMS processes and (2) system miniaurization. The miniaturized and packaged implantable biochip systems of multi-functions are still rare. In this paper a multi-analyte implantable biochip system has been developed. It consists of a biocompatible package, a control and wireless module, a battery, an inductive power coupling module, and a PDMS microchannel (Figure 1). The PMMA package is coated by 5gm-thick parylene-C for biocompatibility. It has circular shape to avoid the tissue injury during implantation. The control module is to drive the DEP micropump and to supply the working voltage to the electrochemical electrodes. Its wireless sub-module includes a microcontroller unit (MCU), an amplifier, and an RF transmission section for signal transmission and process. The battery supplies a 3.7V voltage for the biochip system and can be recharged via inductive coupling" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002441_detc2009-86358-Figure11-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002441_detc2009-86358-Figure11-1.png", "caption": "Figure 11. Gear Crowning with Misalignment and Deflection", "texts": [], "surrounding_texts": [ "The contact stress in a spur involute gear set is usually calculated at the Lowest Point of Single Tooth Contact (LPSTC) of the pinion. The transverse radii of curvature of the gear tooth profiles at this contact point is defined as in AGMA Standards [2] 11 2 1 2 11 /2)( NRRR bbO \u22c5\u2212\u2212= \u03c0\u03c1 (1) 12 sin \u03c1\u03c6\u03c1 mopdC \u22c5= (2) where \u03c11 is the transverse radius of curvature of pinion at LPSTC, \u03c12 is the transverse radius of curvature of gear at Highest Point of Single Tooth Contact (HPSTC), RO1 is the outside diameter of the pinion, Rb1 is the base circle radius of the pinion, N1 is the number of teeth of the pinion, Cd is the center distance of the gear set, \u03c6op is the operating pressure angle, and -/+ is for external and internal gear meshes respectively. The contact stress in a spur gear set with no crowning and no misalignment is defined in AGMA Standards [2] IFd W CS t pc \u22c5\u22c5 = (3) Where, d I op \u22c5\u00b1 = ) 11 ( cos 21 \u03c1\u03c1 \u03c6 (4) ) 11 ( 1 2 2 2 1 2 1 EE C p \u03bd\u03bd \u03c0 \u2212 + \u2212 \u22c5 = (5) where Sc is the contact stress, Cp is elastic coefficient, Wt is tangential load, d is the operating pitch diameter of pinion, F is net face width, I is geometry factor for pitting resistance, \u03bd1 and \u03bd2 are Poisson\u2019s ratio for pinion and gear respectively, and E1 and E2 are modulus of elasticity for pinion and gear respectively. If a gear set is crowned, usually the crown is on the pinion. The contact stress calculated from equation 3 considers only the contact stress without crowning. AGMA does not have an Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 4 Copyright \u00a9 2009 by ASME equation for the contact stress with crowning. We propose using equations from Roark and Young [3], namely, the contact stress for the general case of two bodies in contact. The shape of the instantaneous contact area is an ellipse, and the contact stress is calculated by the following equation. ba P c \u22c5\u22c5 \u22c5 = \u03c0 \u03c3 5.1 (6) where op tW P \u03c6cos = (7) 3/1)( ED CKPa \u22c5\u22c5\u22c5= \u03b1 (8) 3/1)( ED CKPb \u22c5\u22c5\u22c5= \u03b2 (9) d RR K D \u22c5+++ = ) 1111 ( 5.1 2121 \u03c1\u03c1 (10) ) 11 ( 2 2 2 1 2 1 EE CE \u03bd\u03bd \u2212 + \u2212 = (11) where \u03c3c is contact stress for crowned gear pair, P is the normal load, a is the semi length of the instantaneous contact ellipse in the face width direction, b is the semi length of the instantaneous contact ellipse in the profile direction, \u03b1 and \u03b2 are geometrical coefficients [3], R1 is the crowning radius of the pinion, and R2 would be infinite if the gear is not crowned. Gear crowning is specified using the following equation, as shown in Fig. 10. 2 1 2 1 2 1 )( LDropRR +\u2212= (12) Figure 10. Gear Crowning Where, R1 is crowning radius, Drop is the drop over the distance L1 which is from center of crown to the end of the tooth. The drop should include the deflection and misalignment. The general rule for a good crown design should be that the contact pattern within the tooth face at the maximum misalignment condition. 11 LaR \u2264+\u22c5\u03b8 (13) Where, \u03b8 is the angular displacement from the combination of misalignment and deflection. Depending on the amount of misalignment and deflection, with all the equations above, one can optimize the contact so that the contact stress is within the material allowable, and the contact ellipse is stabilized within the tooth face boundary without edge loading. Depending on the application, the optimization between the variables a, R1, \u03b8, and L1 can greatly influence the wear life. Some gears have more deflection than misalignment, as in the case of the compound planet gears shown in Fig. 5, and the contact area is wide. Some gears are tilted by deflection, so the center of the crown is not at the center of the tooth face (called bias crown). Some gears have more misalignment than deflection, as in the case of crowned splines, and contact ellipse is small compared to the face width. Here, the center of the crown is at the center of the tooth face (called full crown). A crowned spline has one more limitation: The tooth thickness has to be modified from the standard because the minimum effective clearance is zero. The tooth thickness Tmod is dependent on R1 and \u03b8. From Figure 12, the following equations can be derived. \u03b8cos/)2/2/( mod dTTX += (14) \u03b8tan/Xy = (15) L1 R1 Drop Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 5 Copyright \u00a9 2009 by ASME )cos1(2/ \u03b8+\u22c5= yG (16) \u03b8tan)2/( \u22c5= GC (17) 22 1 2 1 )2/()( GCRR +\u2212= (18) Where, C is drop in the normal plane, G is total gage length and gage from center of crown is G/2 and dT is the clearance between space width of internal teeth and tooth thickness of external teeth. From the above equations, the following relationship can be derived. [ ] 2 mod1 sin/)cos1()cos1(2/2/ \u03b8\u03b8\u03b8 +\u22c5\u2212\u22c5+= TdTR (19)" ] }, { "image_filename": "designv11_61_0003713_1.3356876-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003713_1.3356876-Figure2-1.png", "caption": "Fig. 2. Left The articulated figure with each element\u2019s angle defined with respect to the positive x direction. Right The same figure with the center element defined with respect to the positive x direction, but with the other two elements defined with respect to the center element.", "texts": [ " In much the same way that the center of mass of a zero linear momentum system cannot move, it appears that these restrictions allow us to define a rotational analog to the center of mass. Without \u201cpropellorlike\u201d motion and the freedom to move in three dimensions, a system that returns to its original relative configuration would appear to return to its original orientation. The purpose of this paper is to show that it is possible for the limited system we have described to experience an orientation change without an external torque acting on it. In particular, we consider a two-dimensional articulated figure2 with two degrees of freedom. A formalism is developed for calculating the rotation of this model when it has zero angular momentum. It is shown that even if propellorlike motion 733 Am. J. Phys. 78 7 , July 2010 http://aapt.org/ajp Downloaded 28 Oct 2012 to 152.3.102.242. Redistribution subject to AAPT lic were forbidden, rotation is still possible. The formalism is then applied to a diver performing a zero angular momentum dive, and the rotation of the body during such a dive is calculated numerically", " 9 becomes = C G d + G d = S G \u2212 G d d , 11 where we have used Green\u2019s theorem and S is the flat surface bound by the closed contour C. Let\u2019s define , the integrand in the surface integral in Eq. 11 , as = G \u2212 G . 12 If we can show that is positive at all points on a certain region of the phase space, then a closed counterclockwise contour in this phase space will yield a positive , and a closed clockwise contour will yield a negative . If were shown to be negative, then the signs for in these cases would be reversed. We now apply the formalism developed in Sec. II to the model depicted in Fig. 2. Each piece is a rod of zero thickness with mass m and length and the angles , , and , defined as in Eq. 3 see Fig. 2 . From Eqs. 1 and 2 , the x-coordinates are found to be x1 = \u2212 6 cos + + cos + , 13a x2 = 6 3 cos + 2 cos + \u2212 cos + , 13b x3 = \u2212 6 3 cos + cos + \u2212 2 cos + . 13c The y-coordinates are similar, with the cosines replaced by sines. The vector G can be found using Eqs. 7 and 10 , giving 734S. John Di Bartolo ense or copyright; see http://ajp.aapt.org/authors/copyright_permission G = \u2212 3 \u2212 3 cos + cos \u2212 13 + 6 cos \u2212 2 cos \u2212 \u2212 6 cos , \u2212 3 + cos \u2212 + 3 cos 13 + 6 cos \u2212 2 cos \u2212 \u2212 6 cos . 14 We use Eq. 14 for G and find see Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001898_s0005117907080024-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001898_s0005117907080024-Figure1-1.png", "caption": "Fig. 1. Fig. 2.", "texts": [ " STABILITY MARGINS OF THE D-SYSTEMS Determination of the stability margins of system (1), (2) requires the following perturbed transfer functions w1(s) = kw(s), w2(s) = e\u2212j\u03b8w(s), (34) where k and \u03b8 are positive numbers. By the stability margin in absolute magnitude [11] is meant the positive number L such that for k > L or k < L\u22121 the perturbed system loses stability and the stability margin in phase is the greatest positive number \u03d5 such that for \u03b8 \u2208 [0, \u03d5] the perturbed system is stable. Since the stability margins in absolute magnitude and phase are determined independently, plausible is the situation shown Fig. 1 where the locus of the amplitude-phase characteristic of the open-loop system (the Nyquist plot) passes arbitrarily near the unit-radius circle and the real axis. For the system with such locus, the stability margins are L = \u221e, and \u03d5 = 45\u25e6. If the perturbed system obeys the transfer function w3(s) = ke\u2212j\u03b8w(s), then it loses stability under arbitrarily small deviations of the number k from unity and \u03b8 from zero. To avoid such situation, we make use the radius of the stability margins [15]. Definition 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001978_aero.2007.352757-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001978_aero.2007.352757-Figure3-1.png", "caption": "Figure 3. The I- I geometrical formation scheme", "texts": [ " Here, helicopter 2 follows 1 through an I- a scheme. Helicopter 3 follows both 1 and 2 and helicopter 4 follows 1 and 3 through I- I scheme, respectively. The I- a control development has been done previously by the authors. This scheme has been reported in [9]. In this paper, we only report the development of the I- I low-level controller. The I- I Control Scheme In order to control the distances of the helicopter constrained with two neighbors, we are going to formulate the geometry shown in Fig. 3 in the following form. Four geometrical parameters 113, 123, 3123 and 03 can be used to determine the relative configuration of the helicopter 3 (follower) with respect to two other neighboring leaders named as 1 and 2 in the three-dimensional (3D) space. A control point p is defined as a point located on the z axis of the follower body coordinate system (Z3). Definition of the control point guarantees the sensitivity of the controller outputs (the formation parameters) to the follower helicopter's roll and pitch motions" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002503_sav-2009-0470-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002503_sav-2009-0470-Figure1-1.png", "caption": "Fig. 1. Symmetric scheme of the bow and arrow system.", "texts": [ " The asymmetry of the system in its main plane is small (3\u20134%), therefore by taking into account a common precision that is acceptable for engineering calculations, it is possible to use in modeling hypothetic forms from the symmetrical scheme of the chain (Appendix). A geometrical linear model of the Euler-Bernoulli beam was considered to be acceptable because; i) an arrow length is \u223c100 time bigger than the diameter of its aluminum alloy cylinder tube and ii) bend and buckle deformation amplitude was concluded to be in the range of 5 % of the arrow length using high speed video footage [12]. Geometric and force equations for the neutral position of the chain are (Fig. 1): h + l cos \u03b8 = s cos \u03b3; l sin \u03b8 + s sin \u03b3 = \u03beA; cs (s\u2212 s0) l sin (\u03b8 + \u03b3) = cl (\u03b8 + \u03d5) , (1) where cl is a virtual stiffness of a limb located in the hinge at the end of a riser; c s is a stiffness of a string branch; s0 is a half-length of the string; \u03d5 is an angle of the virtual limb at the free position (i.e. without a string); \u03beO\u03b7 is \u2018fixed\u2019 to the archer hand immovable system of coordinates. The sense of the rest symbols is clear from the scheme (see Fig. 1). For a small displacement from the neutral position, related equations regarding the upper (U) and lower (L) parts of the kinematical chain are: h + l cos ( \u03b8U/L \u00b1 \u03ba ) = sU/L cos\u03b3U/L \u00b1 \u03b7A; \u00b1h\u03ba + l sin ( \u03b8U/L \u00b1 \u03ba ) + sU/L sin \u03b3U/L = \u03beA; cs ( sU/L \u2212 s0 ) l sin ( \u03b8U/L + \u03b3U/L ) = cl ( \u03b8U/L + \u03d5 ) , (2) where \u03ba is the angular displacement of the riser. We mark in the twin subdivides (U/L) the upper letter concerns the upper part of the chain and opposite. The geometric equations for the neutral Eq. (1) and displaced Eq", " The main determinant of the obtained system of linear (relative the amplitudes) equations should be equal to zero:\u2223\u2223\u2223\u2223\u2223\u2223 b\u03ba\u03ba b\u03ba\u03c4 b\u03ba\u03b7 \u2217 b\u03c4\u03c4 b\u03c4\u03b7 \u2217 \u2217 b\u03b7\u03b7 \u2223\u2223\u2223\u2223\u2223\u2223 = 0, (9) where b\u03ba\u03ba = cs (\u03beA cos \u03b3)2 \u2212 \u03c92I\u03ba; b\u03ba\u03b7 = cs\u03beA cos2 \u03b3; b\u03ba\u03c4 = csl sin (\u03b8 + \u03b3) \u03beA cos \u03b3 \u2212 \u03c92I\u03c4 ; b\u03c4\u03c4 = cs [l sin (\u03b8 + \u03b3)]2 + cl \u2212 \u03c92Il; b\u03b7\u03b7 = cs cos2 \u03b3 \u2212 1 2 mA\u03c92. We can make a previous analysis of the chain vibration by assuming the arrow to be a rigid shaft. This assumption is based on the arrow\u2019s elastic deformations being significantly small in comparison with the virtual displacement which the limbs transferred to the arrow tail A (see Fig. 1) [15]. We also pin inertial properties of the arrow to the tail point A. We get virtual mass of the arrow from the system of equations \u2013 equal linear momentum and zero angular momentum relatively the pinned point A: la\u222b 0 \u00b5(z)\u03b7\u2032a(z)dz = maA\u03b7\u2032A; la\u222b 0 \u00b5(z)\u03b7\u2032a(z)zdz = 0, (10) where \u00b5(z) is the distributed mass of the arrow; la is its length; \u03b7a(z) = \u03b7A + \u03c8z is the displacement of the arrow axis as a rigid shift. Substituting the last form into Eq. (10), we get the following expression for the virtual mass: maA = mar2 a r2 A +r2 a , where rA = la\u222b 0 \u00b5(z)zdz ma is a distance between the tail and the center of mass of the arrow; ma is the total mass of the arrow; ra = \u221a\u221a\u221a\u221a la\u222b 0 \u00b5(z)(z\u2212rA)2dz ma is the radius of inertia of the arrow", " Because this very clear reason, the arrow is situated a little above of the hand holding the bow in the middle of the handle. Hence, the bow and arrow system is asymmetric in the vertical plane. Taking into account a common precision that is acceptable for engineering calculations, lets show why we use an equal symmetrical scheme. Because string elasticity is too small and does not matter in the problem of bow and arrow vibration (see Section 3), we can ignore string strength. A distance between the noke points of the two schemes (symmetrical and asymmetrical) in the relevant neutral position is (see Fig. 1): \u03b7A0 = \u00b1h\u00b1 l cos \u03b8asym U/L \u2213 (s0 \u2213 \u2206s0) cos \u03b3asym U/L , (A1) where \u2206s0 is a half of the difference in the length of the lower and upper branches of the string. The rest nomenclatures have been introduced in the main part of the paper. Relevant distance from the handle to the noke point is: \u03beA = l sin \u03b8asym U/L + (s0 \u2213 \u2206s0) sinasym U/L . (A2) Like it is in the section, let\u2019s connect together geometric equations for the neutral (1) and displaced (A1), (A2) positions of the chain: \u03b8asym U/L = \u03b8 + \u2206\u03b8asym U/L ; \u03b3asym U/L = \u03b3 + \u2206\u03b3asym U/L " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001470_sice.2006.315455-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001470_sice.2006.315455-Figure2-1.png", "caption": "Fig. 2 Definition of posture of a mobile robot.", "texts": [ "1 Structure of mobile robot We introduce the structure of the mobile robot (AMIGOBOT: ACTIVEMEDIA) which we use in this paper. Fig. 1 shows the robot. This robot is driven by two differential wheels, and the robot has a traveling wheel to make a run stability. Furthermore, the robot is equipped This work is supported in part by TATEISI Science and Technology Foundation. with a camera and 8 ultrasonic sensors. 2.2 Kinematics of mobile robot We consider a kinematics of the mobile robot as shown in Fig. 2. In Fig. 2, (Xw, Yw) is a world coordinate system and (XR, YR) is a robot coordinate system. The world coordinate system is without translation and rotation. On the other hand, the robot coordinate system is with translation and rotation, depending on a motion of the robot. In addition an initial position of the robot corresponds with a origin of the world coordinate system. In this paper, we consider a mobile robot driven by two differential wheels. Assuming that the robot moves on the planner surface without slipping, the linear velocity VR and angular velocity WR at the center of the robot can be written as EWR 1a _a Or 0UJRj 0 b b 200 where 0r and QI denote the rotational velocities of the right and left wheels, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002953_ijsurfse.2009.026607-Figure16-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002953_ijsurfse.2009.026607-Figure16-1.png", "caption": "Figure 16 Engagement in the different coordinate system with the x-axis corresponding to the line of action", "texts": [ " To validate the measurement, TE should be calculated and compared with each other using the exact analytical solutions. Here, the equations, which can calculate TE curves with gear eccentricity, are derived shortly. Figure 15 shows an involute gear pair, which has the radial eccentricities of each gear in two dimensions. After rotating each gear about the centre of rotation F by some angles A , two dashed base circles move to the corresponding solid base circles, respectively. To derive the analytical equations more easily, we can redraw this relation. Figure 16 shows gear engagement in the different Cartesian coordinate; the x-axis is always corresponding to the line of action. For the engagement of an involute gear pair, it is easier to understand movement of the gear pair through the line of action. From Figure 16, rotational angle A of each gear is written as follows: 1 0 1A \u03b3 \u03b3 \u03b8= \u2212 + (5) 2 0 2 .A \u03b3 \u03b3 \u03b8= \u2212 + (6) The coordinates of instant centre of rotation 1 1 1( , )F x y\u2032 \u2032 \u2032 and 2 2 2( , )F x y\u2032 \u2032 \u2032 are expressed as follows: 1 0 1 1 1 1 0 1cos e cos( )c bx l r\u03b6 \u03b8 \u03c6 \u03b3 \u03b8\u2032 = + + \u2212 + (7) 1 1 1 1 0 1e sin( )by r \u03c6 \u03b3 \u03b8\u2032 = \u2212 \u2212 + (8) 2 2 2 2 2 0 2e cos( )bx r \u03b8 \u03c6 \u03b3 \u03b8\u2032 = + \u2212 \u2212 (9) 2 2 2 2 0 2[ e sin( )]by r \u03c6 \u03b3 \u03b8\u2032 = \u2212 + \u2212 \u2212 (10) where e1 and e2 are the amount of eccentricity, \u03c61 and \u03c62 are the phase angle of the eccentricity of a driver and a follower, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003215_iecon.2008.4758119-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003215_iecon.2008.4758119-Figure5-1.png", "caption": "Fig. 5. Fault #4, corrosion.", "texts": [ " Then, we have investigated a fault which has not be considered in previous researches, i.e. a deformation of the protective shield (Fig. 4). This fault can be produced by errors during the assembly and can be considered as a cyclic fault, even if it does not produce effect like air-gap eccentricity. So, it is expected not to show particular changes in the current spectrum. Finally, we have produced a corrosion of the bearing, which can be caused by humidity of the environment and can be considered as a generalized roughness (non-cyclic fault, see Fig. 5). The experimental set-up consists of a 2.2 kW three-phase induction motor with two pole pairs, fed by the mains (400 V, 50 Hz) and coupled with a brake [13]. This type of motor is normally used in pumping systems for domestic appliances. The motor has 24 stator slots, 18 rotor bars and 0.5 mm airgap length. The value of the load torque can be imposed and measured by means of the control unit and the visual display unit of the brake. A powermeter collects, for each phase of the motor, the rms value of current and voltage and the measurement of the active power", " So, it has been interesting to check the level up to that the thickness of lubricant is able to protect the structure of the bearings and, in case of failure of this protection, what are the repercussions on the working of the motor. To produce fault #4, the bearing has been plunged for some weeks in water, so as to rapidly simulate the effect of the presence of a humid environment, and then to spoil the layer of grease and to arouse the corrosion. At the end of the tests, it is noticeable that, even in a so hostile environment, the damage of the bearing does not harm its working. In fact, as one can see from Fig. 5, the external structure has been attacked by the corrosion process, but internally the lubricant carries out its task, by obstructing the entrance of the products of the corrosion in the internal structure of the bearing. Therefore, the small differences obtained in the current spectrum compared to the healthy case can be attributed not only to the fact that this is a non-cyclic fault, but also to the weak effect produced on the bearing by this fault. All the considered faults in the bearings have produced a decrease in the efficiency of the motor: the extent of this fall reaches nearly 4% in low load condition and 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002564_s10015-008-0560-2-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002564_s10015-008-0560-2-Figure7-1.png", "caption": "Fig. 7. Locus of the center of gravity", "texts": [ " We shall take the parameters of the system as shown in Table 1. A response of the manipulator from the initial position (\u03b81i = \u2212\u03c0/2, \u03b82i = 0, \u03b83i = 0) to the fi nal position [\u03b81(tf) = 1.3, \u03b82(tf) = 3.0, \u03b83(tf) = 1.1 rad] is shown in Fig. 6, under the condition that working time tf = 1.2 s. In this case, joint 1 is passive, and joint 2, 3 are actuated. In Fig. 6, it is shown that link 1 arrived at desired position \u03b81(tf). And the motion is caused by inertia force of pendulum movement of link 2 and 3. Figure 7 shows a locus of the center of gravity about the motion of Fig. 6. The dynamic equation of the manipulator is rearranged for \u03c41 = 0, and A A A A A 11 21 31 1 2 3 14 11 0 0 1 0 1 0 0 1 0 . . . . \u2212 \u2212 \u23a1 \u23a3 \u23a2 \u23a2 \u23a4 \u23a6 \u23a5 \u23a5 \u23a1 \u23a3 \u23a2 \u23a2 \u23a2 \u23a4 \u23a6 \u23a5 \u23a5 \u23a5 = \u2212 \u03b8 \u03c4 \u03c4 2 2 13 3 24 22 2 23 3 34 32 2 33 3 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u2212 \u2212 \u2212 \u2212 \u2212 \u23a1 \u23a3 \u23a2 \u23a2 \u23a2 \u23a4A A A A A A A \u23a6 \u23a5 \u23a5 \u23a5 (1\u2019) The angular acceleration of joint 2, 3 are given in iterative dynamic programming method, and the angular acceleration of joint 1 and torques of joint 2, 3 are calculated simultaneously" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003411_s11771-012-0984-7-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003411_s11771-012-0984-7-Figure5-1.png", "caption": "Fig. 5 Photograph of spindle with automatic variable preload device", "texts": [ " Thus, it is expected that the device will reduce manufacturing and operating costs because it makes possible to remove hydraulic systems or electric equipments used in the conventional variable preload device. A spindle is fabricated for installing an automatic variable preload device and evaluating its performance. Figure 3 shows the conceptual drawing for fabricating a spindle in which four angular contact ball bearings are used. Two sets of bearings are installed at both the front and rear of the spindle, respectively. Auxiliary bearings, which support the shaft, and main bearings, are installed at the rear and front sides, respectively. Figure 4 shows the spindle fabricated in this work. Figure 5 shows the installation of an automatic variable preload device on the spindle. J. Cent. South Univ. (2012) 19: 150\u2212154 152 For the performance evaluation of the spindle, the vibration and noise according to the installation of an automatic variable preload device were compared and analyzed. A device used to measure the vibration of the system was Vibrometer (VL-8000) by HOFFMANN that presents a resolution of 0.1\u2212200 mm/s, a vibration range of 0.05 mm/s, and an uncertainty of measurement of 0.03 mm/s with its equipment resolution" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002175_robot.2007.363146-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002175_robot.2007.363146-Figure6-1.png", "caption": "Fig. 6. The second case of actuation singularity", "texts": [ " T 222 )( 2 ]0,0,0,,,[ i r i r i r i r nml=$ (25) namely, \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 = 000000 000000 000000 1 0 0 5 2 4 2 3 2 2 2 1 2 5 2 4 2 3 2 2 2 1 2 5 2 4 2 3 2 2 2 1 2 rrrrr rrrrr rrrrr r nnnnn mmmmm lllll $ (26) where ],,[ 222 i r i r i r nml denotes the direction cosine of axis )( 2 i r$ , i=1,2,3,4,5. Obviously, rank of wrench system, $r is three. In this case, the movable platform can still rotate about the rotation center even after locking five actuators. In other words, there are three uncontrollable rotation DoF. This actuation singularity can be passed by choosing different joints as actuators. For example, six wrenches in Eq.(24) are not dependent if choosing three R1 and two R2 as actuators. 2) Second case: As shown in Fig.6, the plane P45 of five limbs parallel to the base plane. Let the reference frame be the same with Fig. 5. In this case, P45 of five limbs will be the same plane. And $r2 of five limbs will be also in the plane. According to Grassmann geometry[5], the rank of coplanar linear vectors (five $r2) is three. Thus, the number of linear independent constraint wrenches (five $r2 and $r1) is four. According to screw theory, from Eq.(23), the last three entries of the $ r2 are the direction cosine of the normal vector for P45" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003437_9781118516072.ch4-Figure4.2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003437_9781118516072.ch4-Figure4.2-1.png", "caption": "Figure 4.2. Nacelle of a wind turbine with (a) gearbox and high speed (Courtesy of General", "texts": [ " An anemometer and a wind vane located on the nacelle roof are used to provide the data necessary for the guidance control system to trigger or stop the wind turbine according to wind speed. Some manufacturers have tried to suppress the gearbox by introducing the \u201cdirect attack\u201d system, or to reduce it. This requires a special electrical generator capable of running at the same speed with turbine rotor speed, which means that the generator must be designed with a large number of poles pairs [1]. Figure 4.2 shows the arrangement of the components in the nacelle for two types of wind turbines. Figure 4.2a shows the conventional drive train design in the form of a geared transmission with a high-speed generator. Figure 4.2b, by contrast, shows the gearless variant with the generator being driven directly from the turbine. A wind turbine can be equipped with an induction generator or a synchronous generator. In terms of the rotational speed, in general, wind turbine systems can be classified into two types: fixed speed and variable speed. The largest machines tend to operate at variable speed, whereas smaller and simpler turbines are of fixed speed. THE AERODYNAMIC PROFILE OF WIND TURBINE\u2019S BLADES. Figure 4.3 shows the general aspect of a horizontal axis wind turbine as well as the blades aerodynamic profile" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002984_s00170-008-1573-7-Figure10-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002984_s00170-008-1573-7-Figure10-1.png", "caption": "Fig. 10 Pocket machining example", "texts": [ " The hexapod operator can select an appropriate curve from the atlas based on the required machining area and depth of machining. The percentage of benchmark stiffness of the selected curve will govern the selection of machining parameters such as depth of cut, feed rate, and cutting speed. In order to illustrate the use of atlas in identification of optimal work piece location based on a selected benchmark stiffness value, pocket machining by hexapod machine tool as shown in Fig. 9 is considered. Figure 10 shows a simple pocket machining example considered for illustration purpose. The selection of workspace location is primarily influenced by the required machining area. The maximum machining path length \u2018x\u2019 and the maximum width \u2018y\u2019 of the machining area perpendicular to x, are identified. The required machining area can be treated as an area enclosed within a rectangle of length x and width y. This area of the rectangle gives the minimum required area of the workspace slice. But the workspace slice should provide ample clearance for tool approach and over travel" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002900_s11465-009-0064-3-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002900_s11465-009-0064-3-Figure2-1.png", "caption": "Fig. 2 Assembly diagram of axle and bearing housing", "texts": [ " (1), (2),\u2026, and (10), we have (shown in the appendix for the detailed deduction process): \u00bdH \u00f03\u00de \u00bc \u00bdTT \u00f01\u00de ab \u00bdTR\u00f01\u00de ab \u2013 \u00bdTT \u00f02\u00de cb \u2013 \u00bdTR\u00f02\u00de cb 2 4 3 5\u00bdHB \u2013 1 0 @ \u2013 \u00bdTT \u00f01\u00de ba \u00bdTT \u00f02\u00de bc \u2013 \u00bdRT \u00f01\u00de ba \u00bdRT \u00f02\u00de bc 2 4 3 5 \u00fe \u00bdTT \u00f01\u00de aa 0 0 \u00bdTT \u00f02\u00de cc 2 4 3 5 1 A, (11) \u00bdHB \u00bc \u00bdH J \u00fe \u00bdTT \u00f01\u00de bb \u00fe \u00bdTT \u00f02\u00de bb \u00bdTR\u00f01\u00de bb \u00fe \u00bdTR\u00f02\u00de bb \u00bdRT \u00f01\u00de bb \u00fe \u00bdRT \u00f02\u00de bb \u00bdRR\u00f01\u00de bb \u00fe \u00bdRR\u00f02\u00de bb 2 4 3 5, (12) where \u00bdH J in Eqs. (11) and (12) is the matrix of the joint region to be identified. 3 Identification of dynamic stiffness matrix of bearing joint Suppose that the axle is substructure 1, the bearing housing is substructure 2, and the assembly of the axle, bearing, and bearing housing is the whole structure 3. The assembly structure is shown in Fig. 2. The bearing is simplified into a five-dimensional stiffness and damping matrix [2,3,7], as shown in Eqs. (13) and (14). \u00bdK \u00bc kxx kxy kxz kx y kx z kyx kyy kyz ky y kz z kzx kzy kzz kz y kz z k yx k yy k yz k y y k y z k zx k zy k zz k z y k z z 2 66666664 3 77777775 , (13) \u00bdC \u00bc cxx cxy cxz cx y cx z cyx cyy cyz cy y cy z czx czy czz cz y cz z c yx c yy c yz c y y c y z c zx c zy c zz c z y c z z 2 666664 3 777775 , (14) where x, y, z, y, and z denotes the x-axle direction, the y-axle direction, the z-axle direction, the y-axle rotational direction, and the z-axle rotational direction, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002412_icems.2009.5382984-Figure12-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002412_icems.2009.5382984-Figure12-1.png", "caption": "Fig. 12. Distribution of acceleration (Radial component) Unit: m/sec2", "texts": [ " Also the structural analysis is performed at the same sensor position. Each result is shown in Figs. 10 and 11. The measured acceleration is shown as the gain towards to the ground vibration in order to eliminate the ground vibration. The largest acceleration around 3000Hz is confirmed in both the measurement and the analysis result. This vibration is due to the resonance phenomenon between the natural frequency and the electromagnetic force. The distribution of the radial component of the acceleration at 3040Hz is shown Fig. 12. However this acceleration is not large, in spite of the vibration mode of the stator case is greatly deformed. This is because the measured point is located as a node in the un-effect deformation area. The effect of the measuring point location was verified by using the analysis result. We were able to confirm a huge acceleration by rotating the measuring point 45 degrees at 3040 Hz and a large amount of electromagnetic vibrations are caused by the resonance phenomenon. The result is shown Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000484_elan.200503301-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000484_elan.200503301-Figure1-1.png", "caption": "Fig. 1. Schematic of the epoxy-PVC glucose sensor.", "texts": [ " To avoid pinhole formation during solvent evaporation, 5 \u2013 10% surfactant agent (Brij 30) may be selectively added to the polymer solutions. The resulting sensing membrane was dried first at 24 1 8C for 24 hours. As needed, a PU membrane was coated upon this incompletely-cured epoxy-membrane with a 2% PU THF solution. Finally, the two ends of the sensing element were sealed with electrically-insulating sealant and then the entire sensor was placed in an oven to dry at 80 8C for 60 minutes.A schematic of theEpoxy-PVCglucose sensor is shown in Figure 1. Newly prepared glucose biosensors were conditioned for at least 2 hours in a 5 mM glucose/PBS solution and then continuously polarized at \u00fe0.7 V vs. SCE until the stable background current was reached. The sensors for long-term observation were always stored in the 5 mM glucose/ phosphate buffer saline (PBS) at 24 1 8C and the stock solution was renewed every 2 days after measurements. The response time was determined as the time needed for the sensorLs current to reach 95%of themaximumcurrentwhen the glucose concentration changed from 5 mM to 15 mM" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003643_ictee.2012.6208635-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003643_ictee.2012.6208635-Figure4-1.png", "caption": "Fig. 4 Dr", "texts": [ " MODERN TOOL FOR TEACHING E DRAWING The modern computer aided teaching to Drawing includes \u2022 2D animations developed on Flash \u2022 Videos of drafting process recorde tion visual presenter \u2022 Animated power point presentation \u2022 Three dimensional models develop Fig. 2 shows a modern classroom de ing Engineering Drawing \u2022 Since while teaching the instructor the students very good concentratio achieved. \u2022 High degree of automation makes t from routine work and therefore he individually on each and every stud Fig. 3 shows the Document camera drafting videos. We have used Elmo P30S record the drafting videos. Fig. 4 shows a drafting video CHING ENGINEER- CED enerally have folconcepts of Engi- ents Engineering three dimensional livering lecture, or ensional diagrams them in hand. ccurately drafting ts on board .But it vertical board and g instruments then e drafting process, struction between students when he es his grip on the NGINEERING ol for Engineering d by high resolu- s ed on Auto CAD ed for Teaching veloped for teach- is facing towards n level is he instructor free can concentrate ent. deos used for making visual presenter to screen shot of a Main advantages of Drafting vid \u2022 The use of Drafting vid instructor to explain th ds the students while teaching eo are eos makes it very easy for the e construction method ecording of drafting process afting Video \u2022 Drafting videos can be played at desired faster speed to save the time of students \u2022 Faster doubt clearing because of replay facility at desired portion of video In general basic concepts of special curves are best explained by 2D-animations made on Flash " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000074_i2003-10034-6-Figure10-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000074_i2003-10034-6-Figure10-1.png", "caption": "Fig. 10. (a) A sketch of the smectic layer morphology in a thin film. Initially the nematic director no is planar. Emerging smectic layers experience a pull towards out-of-plane orientation of their layer normal n, measured by the tilt angle \u03b8, assumed modulated along the y-axis. In equilibrium, the layers would prefer to be parallel to the substrate, the state labelled {eq.}.", "texts": [ " On the other hand, the texture decorated by the stripes is distinctly nematic, since a smectic cannot sustain the bend deformations apparent in a number of images. As we have briefly discussed in the introduction, the principal effect taking place on cooling a thin polymer film from its nematic into the smectic phase is the competition between the planar alignment of the director in the plane of the polymer film and the tendency of the emerging smectic layers to become parallel to the substrate themselves (with the layer normal perpendicular to the substrate). Figure 10 illustrates the point. As the smectic layers first form, with an initially weak smecticorder parameter, their initial orientation should be vertical (in the so-called bookshelf geometry) because the layer normal (the initial director no) is well aligned in the film plane (cf. Fig. 1). As the material is cooled down, the smectic order increases and the layers experience a strong pull to align in the plane of the substrate, with their normal n perpendicular to the plane. A large literature exists on this subject, only partially covered by references [12\u2013 14]", " Assuming an initially small angle of director deflection \u03b8, the corresponding free-energy density takes the form, in the leading approximation, f0 = 1 2K(\u2207\u03b8)2 + (w \u2212 \u03b3s\u03c8 2)\u03b82, (1) where the Frank elastic term (with K \u223c 10\u221211 J/m, as in low-molar mass nematics, see [9]) penalises any nonuniformity in the director distribution. So far, however, our arguments do not lead to any director modulation: the free-energy density f0 is lowered by a uniform increase in the tilt angle \u03b8, as long as \u03b3s\u03c8 2 > w. Given sufficient time to equilibrate, the smectic liquid would certainly reach the conformation with layers aligned in the film plane, labelled {eq.} in Figure 10. In fact, this is essentially what we find on very slow cooling of our samples from the nematic phase, when a rather homogeneous texture is observed. Note that since in the smectic phase the backbone chain anisotropy becomes oblate, with the aspect ratio of gyration r = R\u2016/R\u22a5 < 1 (cf. Fig. 1(a) and imagine rotating the oblate \u201cdisk\u201d until the director n is vertical), the thickness of the polymer film should reduce according to h(\u03b8)\u2212 h0 = \u2212[1 \u2212 r](1 \u2212 cos \u03b8) \u2248 \u2212 1 2 [1 \u2212 r]\u03b82, (2) where h0 is the initial thickness (in the planar nematic phase) and the last expression is an expansion at \u03b8 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000928_gt2004-53708-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000928_gt2004-53708-Figure3-1.png", "caption": "Fig. 3. Determination of film thickness using capacitive sensors", "texts": [ "org/about-asme/terms-of-use Dow uid film thickness is based on the dependence of the capacity on the specific dielectric coefficient. In general, the dielectric coefficient depends on material properties. Therefore, the measuring device is influenced by the volume fraction of the liquid located inside the probe volume. Measuring the capacity, the mean film thickness can be estimated [11]. An application inside the bearing chamber implies a sensor that can be adapted to the geometry without lapping into the chamber. This demand is fulfilled by the sensor shown in Fig. 3. Ground and shield electrode are located as concentric rings around the measuring electrode. Measuring and shield electrode are on the same electric potential, thus the homogeneity of the electric field in the short range of the sensor is given. Increasing the distance of the film surface from the sensor, the electric field becomes more and more inhomogeneous, so that the capacity C and in consequence the film thickness is determined by use of calibration functions. Therefore, the sensors were calibrated at static conditions down to a minimum film thickness of h\u0304 f \u2248 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000209_68.5.441-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000209_68.5.441-Figure1-1.png", "caption": "FIG. 1. Coordinate system and vortex systems. Thick lines denote bound vortices, and thin lines represent wake vorticity. Curved arrows indicate typical directions of vortex rotation, when positive lift is generated. The coordinate system x and y are non-dimensionalized by use of the root chord cmax and the semi-span s/2, respectively.", "texts": [ " One is uniform flow relative to the bird, corresponding to its forward flight. Another is velocity induced by distributed vorticity. When vorticity distribution is known, Biot\u2013Savart\u2019s relation, an inverted version of (2.2), gives induced velocity u: u(r) = 1 4\u03c0 \u222b \u03c9(r\u2032) \u00d7 (r \u2212 r\u2032) |r \u2212 r\u2032|3 dV (r\u2032), (2.3) where r and r\u2032 denote the position vectors of u and \u03c9, respectively. Throughout this study I shall use x and y coordinates non-dimensionalized by use of the root chord cmax and the semi-span s/2, respectively. Figure 1 shows the coordinate system and the vortex systems that stand for birds in a flock. We take x and y coordinates following aeronautical convention as shown in Fig. 1, but z should be a relative altitude. Every bird, with its bound vortex parallel to the y direction, is assumed to fly toward the negative-x direction at the same altitude, z = 0. The reason for the proposed layout, where every bird flies at the same altitude, is given as follows. The i th bird\u2019s drag, a force parallel to the flight direction, is affected most strongly by the neighbouring j th bird, if a z-component of velocity induced by the j th bird becomes the largest. The equation (2.3) shows this occurs if both i th and j th birds are at the same altitude for the fixed x and y components. The situation depicted in Fig. 1 and (2.1) leads to the fact that there is no transverse force to change y-directional intervals between birds, because there is no y-directional component in induced velocity nor z-directional component of the vorticity. Thus the transverse intervals remain unchanged in the present problem and so this study treats transverse intervals as given quantities. It is also pointed out that, in the wake, velocity vectors are parallel to the wake vorticity, and hence the wake will not generate forces. The study does not consider the aerodynamic moments, because their effects are known to be of negligible order (Hummel & Bock, 1981)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003949_j.ast.2011.09.010-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003949_j.ast.2011.09.010-Figure5-1.png", "caption": "Fig. 5. Selected flexible modes for the tuned 27 000 lb model: (left) 1st flexible mode at 6.36 Hz (Boeing: 6.36 Hz; 1st lateral \u2013 aft pylon lateral), (middle) 2nd flexible mode at 6.81 Hz (Boeing: 7.52 Hz, 1st vertical \u2013 aft pylon longitudinal), (right) 3rd flexible mode at 11.62 Hz (Boeing: 12.89 Hz, 2nd lateral \u2013 fwd pylon lateral).", "texts": [ " Apart from the concentrated masses for the rotors, engines etc, the remaining mass are uniformly distributed over the airframe. The 27 000 lb OWE mass was transformed to other mass levels by scaling the concentrated masses along the fuselage centre line. The contributions from heavy mass items at these points were not affected. The modal analysis of the unsupported fuselage model results in a series of resonance frequencies and modes. As the validation will concentrate on the 3/rev loads, only the lowest resonance frequencies are of interest. The results for the first three flexible modes are presented in Fig. 5. Table 2 presents the tuned frequencies compared to the model as presented in Kvaternik et al. [10]. The aeromechanics model already contains rotor mass at the hub location. It was determined that the hub mass in the fuselage model has to be reduced from 2100 to 400 lb in order to prevent duplication (Rhoads [13]) This mainly affects the resonance frequencies (see Table 3 for examples). Mode shapes are very similar for the three aircraft weights. The fuselage model mesh has been refined at the crown of frame 482 thus providing enough detail to capture the actual loca- tion of strain gauges in this region" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002821_j.ijsolstr.2008.05.021-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002821_j.ijsolstr.2008.05.021-Figure1-1.png", "caption": "Fig. 1. Cumulative microslip in a conrod big end system.", "texts": [ " The obtained analytical results are shown to be in agreement with the finite element computations. 2008 Elsevier Ltd. All rights reserved. Conrod assemblies of reciprocating engines are submitted to cyclic thermo-mechanical loadings. Slip mechanisms on the contact interface, such as the cumulative microslip phenomenon (Antoni, 2005; Antoni and Ligier, 2006;Antoni et al., 2007; Antoni and Nguyen, 2008), are often observed. Such unbounded cumulated microslips at each cycle lead to a global rotation of the bearing shell (see Fig. 1). Although the cumulative microslip problem is actually a risk, it has not been well considered in practical analyses of engineers and common elementary tools are proved to be deficient in most cases. Because of the phenomenon\u2019s recurrent appearance, it is important for designers to have some efficient estimates on its occurrence. Since a finite element simulation of the system behaviour requires long and costly computations, it is interesting to have some simplified estimates, especially in the form of analytical criteria ensuring the non-rotation of bearing shells" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003725_s12206-011-0419-7-Figure12-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003725_s12206-011-0419-7-Figure12-1.png", "caption": "Fig. 12. Schematic diagram of the whirling motion.", "texts": [ "13 1 1 1 1 1 1 1 1 1 1 1 1 cos( ) cos( ) cos( ) cos( ) x x fx y y fy x y f F t f F t x X t y Y t \u03c9 \u03d5 \u03c9 \u03d5 \u03c9 \u03d5 \u03c9 \u03d5 = + = + = + = + , 2 2 2 2 2 2 2 2 2 2 2 2 cos( ) cos( ) cos( ) cos( ). x x fx y y fy x y f F t f F t x X t y Y t \u03c9 \u03d5 \u03c9 \u03d5 \u03c9 \u03d5 \u03c9 \u03d5 = + = + = + = + (13) Substituting these variables into Eq. (12), we have 1 1 1 1 1 1 1 1 2 2 2 2 2 2 2 2 sin( ) sin( ) sin( ) sin( ) . x fx x y fy y x fx x y fy y W F X F Y F X F Y \u03c0 \u03d5 \u03d5 \u03d5 \u03d5 \u03c0 \u03d5 \u03d5 \u03d5 \u03d5 = \u2212 + \u2212 + \u2212 + \u2212 (14) Evaluation of Eq. (14) reveals that the work W done by the seal force is the sinusoidal function of four phase differences between the seal force and vibration. As shown in Fig. 12, the tangential component force F\u03c4 is orthogonal to the radial deflection of the rotor relative to the cylinder. When the phase difference \u2206\u03c6 is 90\u00b0, the tangential force F\u03c4 is at its maximum, which is the main factor in rotor whirl instabilities. Consequently, the work done by the total seal force is at a maximum, which is likely to cause rotor instability. The phase difference decreases gradually with increasing inlet pressure, as shown in Fig. 13. The damping effect of the seal force has increasing influence on the vibration under increasing inlet pressure" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000089_rspa.2002.1105-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000089_rspa.2002.1105-Figure7-1.png", "caption": "Figure 7. The Admiralty splice.", "texts": [ " This is because the distance required for the subrope to reach full load is large, and depends on the friction coefficient and the tightness of the twist (pitch). Because of the loss in load capacity and the long splice length required, the Admiralty and Liverpool splices are preferred. Various references (Air Cadets of Canada 1941; British Ropes 1982; Budworth 1983; Cordage Group 1977; Davis & Van der Water 1946; Day 1947, 1953; Graumont & Hensel 1952; Hasluck 1912; Jarman 1984; Klust 1983) give rules for the splice length for the long splice; these are summarized in table 1. The geometry of the Admiralty splice is shown in figure 7; the rope is at the bottom and the splice is evolved by progressing up the figure. The shaded region at the bottom is an R subrope and this is twisted in the clockwise (Z) direction about the axis of the rope and progressing into the splice; as it encounters the splice S subropes, it is woven over and under them in succession, developing a braided structure, in this case a 12-component braid. The S subropes come from the top in an anticlockwise (S) direction and a typical component is shown (shaded braid)", " The path can be estimated using the following assumed equation, r = rm + 1 2\u2206r sin(n\u03c8), where r is the radial position of the subrope, \u03c8 is its angular position and n is the number of rope subropes. The subropes coming from the eye move in the opposite direction and out of phase with these. The rate of change of angular position with the axial station (d\u03c8/dz) cannot be assumed constant here, whereas in the rope, it has been justifiably assumed constant. (b) The direction cosine assumption Refer to figure 7 at the crotch, and consider an axial load on the structure. For an established splice, at any station, all the subropes from the rope and eye are at the same (\u00b1) angle to the axis: those from the eye at a positive angle and those from the rope at the same but negative angle. Thus, in the established part of the splice they all equally contribute to the splice load. Now, consider a neighbouring station: the splice load is the same and again the subropes contribute equally to the same load. Since in the developed region the subrope load does not vary, and since the Proc" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003845_s00170-013-5162-z-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003845_s00170-013-5162-z-Figure7-1.png", "caption": "Fig. 7 Setup of the circular laser hardening modeling", "texts": [ " The absorption coefficient was set equal to 0.45; this value is coherent with the literature [13]. Moreover, the convection coefficient with the environment was set constant too and equal to 0.2 W/mm2 \u00b0C. The time per step (tps) selection was evaluated imposing the constraint to monitor at least ten times each node inside the laser spot. Equation 2 and Fig. 6 define the constraints adopted. tps \u00bc \u0394\u03b1\u22c5tcycle 2\u03c0 where \u0394\u03b1 \u00bc \u03b1spot=10 \u03b1spot \u00bc 2arcsin lspot D tcycle \u00bc 60=n \u00f03\u00de A scheme of the simulation setup is shown in Fig. 7. Under the mentioned constraints, the experimental campaign reported in Table 1 was simulated. The computational time was set equal to the average hardening time. The maximum temperature reached at the end of simulation was acquired. Table 2 reports all the results for each set of process parameters. Table 2 underlines two important considerations: & The software is able to predict with a good accuracy (maximum error equal to 10 %) the final maximum temperature. & The maximum temperature is overestimated for the bars having a diameter equal to 10 mm while on the contrary for higher-diameter values, the maximum temperature is underestimated" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000225_irds.2002.1041467-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000225_irds.2002.1041467-Figure1-1.png", "caption": "Figure 1: WMR coordinates", "texts": [ " The WMR starting from a configuration (position and orientation) ( ~ ~ , y ~ , q 5 ~ ) ~ is required to reach a desired goal configuration (zgr y g , dg)T moving around a known, nominal reference trajectory ( X r , k , Y , , k , @ T , k ) T . k represents the discrete time index. The actually executed, modified trajectory is determined by ( Z k , y k , d k ) T . The WMR generates its actions by processing measurement data originating from one or more beacons. The vehicle motion is described by the model [11] z k + 7JkATcoddk + d J k ) Y k f v k A T s i n ( d k + d J k ) + qg,k , d k + *SZndJk Y AT . 1 (3 (;:::) dk+l = ( where X k and Y k denote the WMR position coordinates with respect to a b e d frame (Fig. 1) , and d k represents the orientation angle with respect to the x axis. They form the state vector i k = ( Z k , Y k , $ + k ) T . L is the wheel base (the distance between the front steering wheel and the axis of the driving wheels), AT is the sampling time, and qk = (q,,k,q2,k, q4,k) T represents the process noise. The WMR is controlled through a demanded velocity zik and a direction of travel $ k , i.e. the control vector is U k = ( V k r & ) T . Due to physical constraints, both the velocity vk, and the angle dJk cannot exceed boundary values, namely U k E [o,vmaz], dJk E [-$maz,$moz] ($ma= 5 I )" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001139_iros.1992.587374-Figure14-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001139_iros.1992.587374-Figure14-1.png", "caption": "Fig. 14: Dynamic environment with a moving box.", "texts": [ "ry to achieve a smooth p&li by interpolation, such a new planning results in a short, decrease of robot's speed only (see section 4). IV. RESULTS The discussed algorithms have been realized on a 80386 based real-t,ime computer system (16 MHz) for two different robots (ICUKA 160 and MANUTEC 1-3). A few siiiiulat,ion resulk which illustrate the performance of the syst,eiri are given in this sedion. To 1nea.siire the computation time for upda.tirig the cspace model dynamically the environment shown in fig. 14 is used. The marked box moves along a zigzag-course between the other obstacles, which a.re stat,ionary. For one update of the c-space model using the OCMEM-algorithm introduced for quasi-st>ationary obstacles a cornput,at,ion time of about 1.7 seconds is needed. By utilization of the OCMEM algorit-lim for dyna- 45 1 mic obstacles this t ime is reduced to less the 400 msec. in average. Thus a real-time update of t,he e-space model is reached even on a 80386 processor for comparatively slow object motion" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003686_20110828-6-it-1002.03420-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003686_20110828-6-it-1002.03420-Figure1-1.png", "caption": "Fig. 1 : Compass-Type Biped Robot", "texts": [ " By right/left discrete external forces, a continuous external force F c : TQ \u2192 T \u2217Q can be discretized as F d+(qk, qk+1)=(1\u2212\u03b1)hF c ( (1\u2212\u03b1)qk+\u03b1qk+1, qk+1\u2212qk h ) , F d\u2212(qk, qk+1)=\u03b1hF c ( (1\u2212\u03b1)qk + \u03b1qk+1, qk+1\u2212qk h ) . (9) Therefore, by calculating variations for (7), we obtain the discrete Euler-Lagrange equations with discrete external forces: D1L d(qk, qk+1) + D2L d(qk\u22121, qk) + F d+(qk, qk+1) + F d\u2212(qk\u22121, qk) = 0, k = 1, \u00b7 \u00b7 \u00b7 , N \u2212 1. (10) In this subsection, we first give a problem setting of the compass-type biped robot. In this paper, we consider a simple compass-type biped robot which consists of two rigid bars (Leg 1 and 2) and a joint without rotational friction (Waist) as shown in Fig. 1. In Fig. 1, Leg 1 is called the supporting leg which connects to ground and Leg 2 is called the swing leg which is ungrounded. Moreover, for the sake of simplicity, we give the following assumptions; (i) the supporting leg does not slip at the contact point with the ground, (ii) the swing leg hits the ground with completely inelastic collision, (iii) the compass-type biped robot is supported by two legs for just a moment, (iv) the length of the swing leg gets smaller by infinitely small when the swing leg and the supporting leg pass each other", " In addition, we assume that Leg 1 is the swing leg and Leg 2 is the supporting leg in odd-numbered swing phases, and Leg 1 is the supporting leg and Leg 2 is the swing leg in even-numbered swing phases. This subsection presents a continuous-time model of the compass-type biped robot which is called the continuous compass-type biped robot (CCBR) via normal continuous mechanics based on the problem setting shown in the previous subsection [10]. Let \u03b8(i), \u03c6(i) be angles of Leg 1 and Leg 2 in the i-th swing phase, respectively. We derive both the swing phase and the impact phase for the case where Leg 1 is the swing leg and Leg 2 is the supporting leg as shown in Fig. 1. Hence, it is noted that for the case where Leg 1 is the supporting leg and Leg 2 is the swing leg, we can easily obtain the model by changing \u03b8(i) for \u03c6(i) in the both models. First, we consider the model of the swing phase of the CCBR. We substitute the continuous Lagrangian (11) into the continuous Euler-Lagrange equations. Then, we assume that a torque to the waist can be controlled and is denoted by v(i). Adding the torque v(i) to the righthand side of the Euler-Lagrange equations, we obtain the CCBR\u2019s model of the swing phase as (12)", " We define the notations for the DCBR: h: a sampling time, \u03b1: a division ratio in discrete mechanics, k = 1, 2, \u00b7 \u00b7 \u00b7 , N : a time step, i = 1, 2, \u00b7 \u00b7 \u00b7 , L: the order of swing phases, \u03b8 (i) k , \u03c6 (i) k : the angles of Leg 1 and 2 at the time step k in the i-th swing phase, respectively, u (i) k : the control input at the time step k in the i-th swing phase as a discrete torque for the swing leg. We first derive the swing phase model of the DCBR for the case where Leg 1 is the swing leg and Leg 2 is the supporting leg as shown in Fig. 1. Note that for the case where Leg 1 is the supporting leg and Leg 2 is the swing leg, we can easily have the model by changing \u03b8 (i) k for \u03c6 (i) k . Cal- culate the discrete Lagrangian Ld \u03b1(\u03b8(i) k , \u03b8 (i+1) k , \u03c6 (i) k , \u03c6 (i+1) k ) from (11) as (2) and substitute it into the discrete EulerLagrange equations (6). Moreover, adding the control input to the left-hand side of the discrete Euler-Lagrange equations, we obtain the swing phase model as f1(\u03b8 (i) k\u22121, \u03b8 (i) k , \u03b8 (i) k+1, \u03c6 (i) k\u22121, \u03c6 (i) k , \u03c6 (i) k+1, u (i) k ) = 0, (14) f2(\u03b8 (i) k\u22121, \u03b8 (i) k , \u03b8 (i) k+1, \u03c6 (i) k\u22121, \u03c6 (i) k , \u03c6 (i) k+1, u (i) k ) = 0, (15) where functions g1 and g2 are defined as (16) and (17), respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002885_wnwec.2009.5335810-Figure14-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002885_wnwec.2009.5335810-Figure14-1.png", "caption": "Figure 14. FEA modeling when considering thermal-mechanical coupling", "texts": [ " However, the maximum bending deformation of the output shaft is 0.0249 mm if bearing clearance is considered. Same modeling is used to analyze the thermal-mechanical coupling here. The differences are that temperature restraint and convective exchange coefficient should be set before the thermal-mechanical coupling is analyzed. Temperature constraint was applied in the heat source, such as: bearings, wheel and pinion. FEA modeling of the gear pair after considering thermal-mechanical coupling is shown in Fig. 14. As shown in Fig. 14, the red section denotes the torque applied in the spline; the blue section denotes the displacement constraints applied in the bearing. The temperature constraints were applied in the yellow section and the gradient temperature field was applied in the contact area of wheel and pinion. The atmosphere temperature is 25 . Other thermodynamics coefficients are as below: the thermal conductivity is 1.35E-3W/(m\u00b7K) and the thermal expansion coefficient is 1.12E-5/K. The convective exchange coefficient is set as Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000702_tro.2005.844677-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000702_tro.2005.844677-Figure2-1.png", "caption": "Fig. 2. Rigid-body and body and inertial coordinate frames.", "texts": [ " The twist vector of a body is understood here as a set of scalar variables that comprise the necessary and sufficient amount of information to determine the velocity field in the body. For a rigid body, the three components of the body angular velocity and those of the velocity vector of a landmark point of the body provide this information. We define below two different twist vectors for a rigid body: 1) vector , called the twist and 2) vector , termed the Cartesian twist (1) where , , and , shown in Fig. 2, are the position vector of the local-frame origin, its absolute velocity, and the frame angular velocity, respectively, all expressed in the local frame. Furthermore, according to the convention set forth in Section I, is the element-wise time derivative of vector ; it is, in general, different from . In fact, we have (2) Hence, the twist and the Cartesian twist are related through (3) in which is the cross-product matrix1 of vector , while and are the 3 3 identity and zero matrices, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000014_memsys.2002.984233-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000014_memsys.2002.984233-Figure5-1.png", "caption": "Figure 5: Photomicrograph of inclined SU-8 structure.", "texts": [ " However, if the SU-8 used is not of sufficiently high viscosity, the ramp upon curing may possess a concave curvature. To minimize this curvature and ensure that the ramp is as flat as possible, multiple microdispersions may be necessary to smooth out the concavity. After curing, the sidewalls of the ramp were photodefined in the SU-8, and the photoresist guides were removed, leaving behind just the ramp and the plateau that had served as its far-wall. At this point, the inclined structure fabrication is completed (Figure 4e). Figure 5 shows a close view of an SU-8 inclined ramp structure. To continue with fabrication of magnetic switches, thick positive photoresist (Clariant AZ PLP 100 XT) was spun over the sample surface to act as a sacrificial layer of 30 um thickness. This sacrificial layer PR was patterned to expose areas on the SU-8 plateau upper surface on which the cantilever would be anchored. Next a seed layer of Cu was conformally deposited over the entire surface and a second photoresist layer deposited over this seed layer, which was then patterned to form the mold for the cantilever beam" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000297_s0076-6879(04)81045-0-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000297_s0076-6879(04)81045-0-Figure1-1.png", "caption": "Fig. 1. Schematic of the catheter-type Clarke-style amperometric oxygen sensor.", "texts": [ " This article describes the procedure employed to construct functional NO-releasing catheter-type amperometric oxygen sensors, including coating the sensor with NO-release materials, assembly of the sensor itself, and in vivo evaluation of the analytical performance and hemocompatibility of the device (details of the surgical protocol used for animal studies are described elsewhere11). The following procedure describes the specific fabrication of an intravascular catheter-type sensor (for use in animal studies only) that is introduced into an artery via a 14-gauge, 1.16-in. angiocath. The sensor (see Fig. 1) is fixed into a four-way stopcock to allow a means to attach the sensor to the catheter securely and to introduce a saline drip to prevent blood from pooling in the catheter. The length and diameter of the sensor can be adjusted by selecting tubing of the appropriate diameters and adjusting lengths of wires and the sensor body to accommodate a wide variety of sizes and specific methods of introducing the sensor into the blood vessel or tubing to be monitored (e.g., sensors could also be used for on-line PO2 measurement in perfusion fluids in extracorporeal circuitry, etc", " Nuthakki, R. E. Callahan, C. J. Shanley, J. K. litis, J. Elmore, S. I. Merz, and M. E. Meyerhoff, J. Med. Chem. 46, 5153\u20135161 (2003). Fi silico with Silastic medical grade tubing (0.55 mm i.d. 0.94 mm o.d.) (Helix Medical, Inc., Carpinteria, CA) Dow Corning RTV-3140 silicone rubber (World Precision Instruments, Inc., Sarasota, FL) 26-gauge steel wire (hangers for dip coating sensor sleeves) 1. Cut 26-mm lengths of Silastic medical grade tubing (0.51 mm i.d. 0.94 mm o.d.) to make sensor sleeves (see Fig. 1). One end of the tubing is filled with a SR plug by putting a small bead of Dow Corning 3140 RTV-SR in a weighing dish and gently dipping one end of the tubing into the SR several times (see Fig. 2A). Continue dipping until about 2 mm of the end of the tubing is filled with SR. It is important to remove excess SR from the outside of the tubing so that the sensor sleeves will be smooth. Allow SR plug to cure overnight under ambient conditions. g. 2. Illustration of process for (A) forming a sensor sleeve by plugging one end of ne rubber tubing with silicone rubber and (B) arrangement for dip coating sensor sleeves NO-release material" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000746_12.619143-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000746_12.619143-Figure6-1.png", "caption": "Fig. 6 Schematic of the radial dimension changes that occur in the lens mounting of Fig. 1 with increasing temperature.", "texts": [ " (19) The value for CR is then added to the denominator ofEq. (4) to derive a new K3 including this effect. 3.3 Effect of shoulder deflection The shoulder also acts as a continuous ring flange so Eqs. (16) through (19) can be used to determine its deflection Ax under preload and its axial compliance C. One needs to substitute t for tR and apply the appropriate values for a and b for the shoulder. The compliance Cs is added to the denominator ofEq. (4) to derive a new K3 including this effect. 3.4 Effect of radial dimension changes of the lens and cell Figure 6 approximates schematically how the interfaces between the retainer and the lens and between the shoulder and the lens change when the temperature rises by a AT of 1\u00b0F. The locations of the interfaces move radially outward by Ay because of differential expansion with UM > aG. Since the lens surfaces are inclined by the angles p and p2 at the interfaces, those interfaces move axially toward each other by Ax1 + Ax2. The following relationships apply: (j 9O\u00b0\u2014arcsin(y/R1) (20) Ay (UM _aG) yc (21) Ax =\u2014Aye / tan q" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002089_1464419jmbd49-Figure11-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002089_1464419jmbd49-Figure11-1.png", "caption": "Fig. 11 Main dimensions for the motorcycle", "texts": [ "comDownloaded from Kiy integral control gain of rider Kpy proportional control gain of rider Km engine constant Kn spring constant of tyre Ksh linear stiffness coefficient of monoshock lBE 0 undeformed monoshock length lBC, lCD, lBE distances between points on the rear suspension system M mass matrix Mn aligning tyre torque Msw swing arm torque q vector of generalized coordinates Q force vector rsp sprocket radius R polar radius during cornering R0 nominal radius of road curve Reff,r effective rolling radius of rear wheel Rw deformed wheel radius sf extension of front suspension slong, slat, sn longitudinal, lateral, and \u2018normal\u2019 slips stot total tangential slip in tyre contact S transformation matrix between rates of change of generalized coordinates and generalized velocities t time Tm engine torque Tb steering torque u vector of generalized velocities vrel relative velocity between pressure point and surrounding air V0 nominal forward velocity of motorcycle Vw longitudinal component of wheel velocity x vector of state variables xB, xC, xD x-coordinates of suspension points xpp, zpp coordinates of pressure point with respect to rear frame X, Y, Z coordinates of the reference point for rear frame y difference between polar radius R and nominal curve radius R0 z vector of user-defined state variables zB, zC, zD z-coordinates of suspension points b steering angle gw wheel camber angle 1 steering head angle _elong, _elat, vn longitudinal and lateral slip velocities and normal spin rate en normal tyre deflection ~en smoothed tyre deflection en0 smoothing parameter for tyre deflection l vector of system parameters mw coefficient of sliding friction at high speed jeq ratio of cornering to camber stiffness jlong, jlat slip coefficients jn tyre yaw damping coefficient r transverse crown radius of tyre rair density of air slong, slat, sn relaxation lengths slong,0, slong,1 coefficients for the longitudinal relaxation lengths f roll angle of rear frame x pitch angle of rear frame xf , xr rotation angle of front and rear wheels xsw rotation angle of swing arm c yaw angle of rear frame cr relative yaw angle during cornering C polar angle during cornering of motorcycle v0 angular velocity vector of wheel Subscripts ss value at steady state w for generic wheel, it can denote either front or rear wheel Superscripts ^ a hat over a variable denotes a maximal value . a dot over a variable denotes a total derivative with respect to time t APPENDIX 2 MOTORCYCLE DATA In this appendix, the default parameter values of the motorcycle model in section 2 are listed. Figure 11 gives the main dimensions, while the symbols for the parameters with their values and a description are listed in Tables 1 to 3. Note that the effective rolling radius Reff,r in equation (16) is calculated from the deformed radius, the nominal load distribution between the wheels, the normal stiffness and the longitudinal slip parameter jlong. Proc. IMechE Vol. 221 Part K: J. Multi-body Dynamics JMBD49 # IMechE 2007 at The University of Iowa Libraries on March 16, 2015pik.sagepub.comDownloaded from a 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001100_11816171_15-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001100_11816171_15-Figure1-1.png", "caption": "Fig. 1. The laboratory set-up TRMS", "texts": [ " The Two Rotor MIMO System (TRMS) is a laboratory set-up deigned for control experiments. In certain aspects, its behavior resembles that of a helicopter. From the control point of view it exemplifies a high order nonlinear system with significant cross-couplings. The approach to control problems connected with the TRMS proposed in this paper involves some theoretical knowledge of laws of physics and some heuristic dependencies difficult to express in analytical form. A schematic diagram of the laboratory set-up is shown in Fig. 1. The TRMS consists of a beam pivoted on its base in such a way that it can rotate freely both in the horizontal and vertical planes. At both ends of the beam, the rotors (the main and tail rotors) are driven by DC motors. A counterbalance arm with a weight at its end is fixed to the beam at the pivot. The state of the beam is described by four process variables: horizontal and vertical angles measured by position sensors fitted at the pivot and two corresponding angular velocities. Two additional states variables are the angular velocity of the rotors measured by tachogenerators coupled with the driving DC motors. In a normal helicopter, the aerodynamic force is controlled by changing the angle of attack. The laboratory set-up from Fig. 1 is so constructed that the angle of attack is fixed. The aero dynamic force is controlled by varying the speed of the rotors. Therefore, the control inputs are supply voltage of DC motors. A change in the voltage value results in a change of the rotation speed of the propeller which results in a change of the corresponding position of the beam. A system performance index is used for fitness function in the RGA. It is an optimization criterion for parameters tuning of control system, which is suitable for the RGA" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002878_acc.2008.4587030-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002878_acc.2008.4587030-Figure1-1.png", "caption": "Fig. 1. The LVLH and Body Frame", "texts": [ " This dynamics is forced by the control defect between the applied and the saturated values. In essence, the net effect is to have modified the x\u0308r values in the overall control law to facilitate the actuator saturation. To illustrate the effectiveness of the proposed adaptation laws in presence of control saturation, we consider the problem of spacecraft rendezvous. Let us consider the chaser spacecraft motion relative to the target spacecraft, in the Local-Vertical-Local-Horizontal (LVLH) frame as shown in Fig. 1. The LVLH reference frame is attached to the center of mass of target space vehicle with X-axis pointing radially outward of its orbit, Y -direction perpendicular to X along its direction of motion and Z completes the right handed co-ordinate system. Usually in rendezvous and docking problems, the trajectory of target spacecraft is described in the LVLH coordinate system, and this frame is taken as the reference target trajectory for the chaser spacecraft. The relative dynamics between two spacecrafts is governed by fully non-linear Clohessy-Wiltshire equations, given as follows [19]: x\u0308\u2212 2\u03b8\u0307y\u0307 \u2212 \u03b8\u0308y \u2212 \u03b8\u03072x = \u2212\u03bc(rc + x) \u03c13/2 + \u03bc r2 c + Fx m y\u0308 + 2\u03b8\u0307x\u0307 + \u03b8\u0308x\u2212 \u03b8\u03072y + \u03bc r3 c y = \u2212 \u03bcy \u03c13/2 + \u03bc r3 c y + Fy m z\u0308 + \u03bc r3 c z = \u2212 \u03bcz \u03c13/2 + \u03bc r3 c z + Fz m (14a) r\u0308c = rc\u03b8\u0307 2 \u2212 \u03bc r2 c \u03b8\u0308 = \u22122 r\u0307c\u03b8\u0307 rc \u03c1 = \u221a (rc + x)2 + y2 + z2 where x, y, z represents the relative position of chaser spacecraft w" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000796_s00170-003-1681-3-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000796_s00170-003-1681-3-Figure4-1.png", "caption": "Fig. 4 The platform and the inverse screws", "texts": [ " 3 The inverse screws applied to the platform the direction \u00f0 cos b1 cos a1 cos b1 sin a1 sin b1 \u00de, which is just the axis of rod P1B1 shown in Fig. 3. The same is with the other two branches of the platform. So, the inverse screws (constraint forces) applied to the platform are three pure forces along the individual axis of the rods and three pure moments of couples, two of which are perpendicular in the direction of the z-axis. In order to analyze the singularity of the platform, we create an absolute Cartesian coordinate at the geometry centre of the platform shown in Fig. 4. For the two pure moments of couple that are perpendicular to the direction of the z-axis will prevent the platform from rolling around the x and y axes unless $r1 P 1B1 , $r1 P3B3 in a line and $r1 P2B2 perpendicular to the xy- plane, which is impossible in reality. Similarly, $r1 P 2B2 is neither parallel to nor superposed with any of $r1 P 1B1 and $r1 P3B3 . That is, three pure moments of couple applied to the platform can be spanned in the three linear normal spaces. Therefore, they can constrain the three orthogonal revolutions and the platform can only execute translational movements along three orthogonal directions", " So, we obtain one singularity constraint: b \u00bc 0 or b \u00bc p 2 \u00f04\u00de If the platform is stable and b 6\u00bc 0; b 6\u00bc p 2, the fol- lowing equations must always hold: AFinv \u00bc F ; where A \u00bc cos a1 cos a2 cos a3 sin a1 sin a2 sin a3 1 1 1 2 4 3 5 Finv \u00bc $r2 P1B1 $r2 P2B2 $r2 P3B3 T ; F \u00bc Fx= cos b Fy= cos b Fz= sin b\u00bd T ; Fx, Fy, Fz are any external forces. Because $r1 P1B1 , $r1 P3B3 will not be in a line, the singularity criterion is that the determinant of the coefficient matrices A is zero. Set |A|=0, the following equation can be gained. sin a3 a2\u00f0 \u00de \u00fe sin a2 a1\u00f0 \u00de \u00fe sin a1 a3\u00f0 \u00de \u00bc 0 \u00f05\u00de Beside the constraint of Eq. 2, the mechanism has geometry constraints, shown as Fig. 2 and Fig. 4. The geometry constraints can be expressed as: l cos b cos a1 \u00bc l cos b cos a3 l cos b cos a1 l cosb cos a2 \u00bc w h That is, cos a1 \u00bc cos a3 cos a1 cos a2 \u00bc w h l cos b ( \u00f06\u00de From equation cosa1=cosa3 and the geometry character of the mechanism, we can gain a3 \u00bc 2p a1 cos a1 cos a2 w h l cos b \u00bc 0 sin a1 cos a2 \u00fe 3 cos a1 sin a1 \u00bc 0 8 < : \u00f07\u00de Assume x1 \u00bc cos a1 x2 \u00bc cos a2 ; According to Fig. 2, we can gain: 1 > x1 > 0 1 6 x2 6 0 \u00f08\u00de Considering the geometry character, we obtain the following equations: x1 x2 w h l cos b \u00bc 0 x2 3x1\u00f0 \u00de ffiffiffiffiffiffiffiffiffiffiffiffi 1 x21 q \u00bc 0 8 < : \u00f09\u00de Therefore, the criteria of the singularity is: h w l cosb 2x1 ffiffiffiffiffiffiffiffiffiffiffiffi 1 x21 q \u00bc 0; that is, x1 \u00bc h w 2l cos b or x1 \u00bc 1 or x1 \u00bc 1 So we obtain: a1 \u00bc arccos h w 2l cos b or a1 \u00bc p or a1 \u00bc 0: In the following, we will discuss the three cases: 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003233_pime_conf_1965_180_321_02-Figure24.1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003233_pime_conf_1965_180_321_02-Figure24.1-1.png", "caption": "Fig. 24.1 7 . Correlation of experimental eccentricity ratio with attitude locus for", "texts": [ " Oil feed pressure. Pressure. Volumetric flow rate. Flow rate due to pressure. Flow rate due to velocity. Side leakage from load-carrying wedge. Recirculated flow rate. Radius of bearing. Sommerfeld number = - - * Specific heat. Temperature. Shaft surface speed. Applied load. Eccentricity ratio. Viscosity. Effective viscosity of oil feed. Coefficient of friction. Density. Attitude angle. r lUb w c \u201ci2 Subscripts 1, 2 , . . . Circumferential position in bearing (see Fig. 24.10). The bearing test rig is shown in Fig. 24.1. The test shaft, driven by a 75-hp motor by means of a 40: 1 step-up gearbox, is located in three bearings. Counting from the gearbox outwards these are the supporting bearing, the loading bearing, and the test bearing, Fig. 24.2. The loading bearing applies a load vertically downwards by Vol 180 Pt 3K at UNIV NEBRASKA LIBRARIES on May 25, 2016pcp.sagepub.comDownloaded from Vo1180 Pt 3K Proc Instn Mech Engrs I96546 at UNIV NEBRASKA LIBRARIES on May 25, 2016pcp.sagepub.comDownloaded from 78 K. G" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003486_978-3-642-28572-1_24-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003486_978-3-642-28572-1_24-Figure4-1.png", "caption": "Fig. 4 This figure shows samples of the learned forehands. Note that this figure only illustrates the learned meta-parameter function in this context but cannot show timing and velocity and it requires a careful observer to note the important configuration differences resulting from the meta-parameters.", "texts": [ " As cost function we employ the metric distance between the center of the paddle and the center of the ball at the hitting time. The policy is evaluated every 50 episodes with 25 ball launches picked randomly at the beginning of the learning. We initialize the behavior with five successful strokes observed from another player. After initializing the meta-parameter function with only these five initial examples, the robot misses about 95% of the balls as shown in Figure 3. Trials are only used to update the policy if the robot has successfully hit the ball. Figure 4 illustrates different positions of the ball the policy is capable of dealing with after the learning. Figure 3 illustrates the costs over all episodes. Preliminary results suggest that the resulting policy performs well both in simulation and for the real system. We are currently in the process of executing this experiment also on the real Barrett WAM. In this experiment, we built a movement library for the table tennis stroke consisting of 300 motor primitives sampled from successful movements" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001008_kem.291-292.163-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001008_kem.291-292.163-Figure2-1.png", "caption": "Fig. 2 A model of crankshaft pin grinding mechanism Fig. 3 An analysis model of crankshaft pin grinding", "texts": [ " However the grinding mechanism is not investigated well. Therefore, in order to make the grinding mechanism clear, the substantial surface speed of both the workpiece and the grinding wheel are investigated. Then, the grinding results such as grinding force, surface roughness and residual stock are estimated by simulation method, and effectiveness for precision under the controlled speed ratio grinding method is compared to the one under the constant journal rotation speed grinding method. Analysis of the crankshaft pin grinding mechanism Fig. 2 shows a model of grinding mechanism of the pin. When the journal rotates, then the pin separates from the wheel surface. In order to Fig. 1 Crankshaft pin grinding with wheel head oscillating type CNC crankshaft pin grinder 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-13/04/15,08:26:42) maintain the grinding condition, the wheel head moves horizontally" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001598_acc.2006.1656432-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001598_acc.2006.1656432-Figure2-1.png", "caption": "Fig. 2. Example of desired interception geometry between the Nyquist plots", "texts": [ " This restriction will be met if T (j\u03c9d) = p + jq, or using (5): j\u03c9db1 + b0 (j\u03c9d)3 + a2(j\u03c9d)2 + a1(j\u03c9d) + a0 = p + jq (6) Thus b0 + jb1\u03c9d = p(a0 \u2212 a2\u03c9 2 d) \u2212 q(a1\u03c9d \u2212 \u03c92 d) + (7) j [ q(a0 \u2212 a2\u03c9 2 d) + p(a1\u03c9d \u2212 \u03c93 d) ] Introducing the new variables: x = a0 \u2212 a2\u03c9 2 d (8) y = a1 \u2212 \u03c92 d (9) Eq. (8) yields the design equations: b0 = px \u2212 q\u03c9dy (10) b1 = py + qx \u03c9d (11) Conditions on T (s) parameters are now established to satisfy system stability. To study the stability of a limit cycling system in which the DF is a complex function, the Loeb Criterion will be used [1] in order to enforce stable oscillation. The Loeb criterion permits studying the case in which T (j\u03c9) intersects \u22121/N(A,\u03c9) in any point in the complex plane. Fig.2 represents a possible desired intersection between T (j\u03c9) and \u22121/N(A,\u03c9) in the complex plane. The oscillation condition is written as: C(j\u03c9)G(j\u03c9) = \u2212 1 N(A,\u03c9) (12) Since the intersection does not lie on the real axis, Eq. (12) can be written in the form: X(A,\u03c9) + jY (A,\u03c9) = 0 (13) Consider the nominal periodic solution: A = Aoe j\u03c9ot (14) If a perturbation is introduced in the nominal solution such that the nominal amplitude Ao is lead to Ao + \u0394A and the nominal frequency \u03c9o to \u03c9o + \u0394\u03c9 + j\u0394\u03c3, the perturbed solution can be written as: A = (Ao + \u0394A)ej(\u03c9o+\u0394\u03c9+j\u0394\u03c3)t (15) Considering this perturbation, Eq", " (20) for the parameters of the shaping function T (s), the restriction to be satisfied is: (K4\u03c9 3q + (K3K4 \u2212 K1K2)\u03c9 4)y4 + (21) (2K2\u03c9 5p \u2212 K4(4q\u03c95 + (\u22122a2\u03c9 2p \u2212 2q\u03c93)\u03c92))y3 + (K2(\u22124p\u03c95a2 2 + (2a2 2p\u03c93 + 2qa2\u03c9 4)\u03c92) \u2212 K4(\u22125q\u03c95a2 2 + (3qa2 2\u03c9 3 \u2212 2a2p\u03c94)\u03c92) + 2(K3K4 \u2212 K1K2)a 2 2\u03c9 6)y2 + (2K2\u03c9 7pa2 2 \u2212 K4(4qa2 2\u03c9 7 + (\u22122a2\u03c9 2p \u2212 2\u03c93q)a2 2\u03c9 4))y + (K2((2a2 2p\u03c93 + 2qa2\u03c9 4)a2 2\u03c9 4 \u2212 4a4 2\u03c9 7p) + (K3K4 \u2212 K1K2)a 4 2\u03c9 8 \u2212 K4((3qa2 2\u03c9 3 \u2212 2a2\u03c9 4p)a2 2\u03c9 4 \u2212 4qa4 2\u03c9 7)) \u2265 0 Amplitude and frequency of the limit cycle shall be robust against modelling uncertainty. This restriction will be met if T (j\u03c9) crosses \u22121/N(A,\u03c9) perpendicularly [8]. Calling \u03b8 the angle between T (j\u03c9) and \u22121 N(a,\u03c9) at the point (p, q), see Fig. 2, the projection of T (j\u03c9) in the direction of the tangent of \u22121 N(A,\u03c9) at (Ad, \u03c9d) can be written as: proj\u03b8T (j\u03c9) = Re(T (j\u03c9)) cos(\u03b8)+Im(T (j\u03c9)) sin(\u03b8) (22) Then, following the same development in [8], the necessary condition to have a robust limit cycle is T (j\u03c9) and \u22121 N(a,\u03c9) be perpendicular to each other. This condition is met if: \u2202proj\u03b8T (j\u03c9) \u2202\u03c9 = 0 (23) Computing the derivative and making the necessary substitutions in order to have an expression in variables y and a2, the robustness requirement is written as: sqy2 + (2\u03c9dsa2p \u2212 2\u03c92 dcp \u2212 2\u03c92 dsq)y + 2\u03c92 dca2 2p + (24) \u03c92 dsqa2 2 \u2212 2c\u03c93 da2q + 2sa2\u03c9 3 dp = 0 where, s = sin\u03b8 and c = cos\u03b8" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002130_iciea.2007.4318411-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002130_iciea.2007.4318411-Figure3-1.png", "caption": "Fig. 3. Geometric characteristics of motor.", "texts": [ "00 c\u00a92007 IEEE The control is called \u201cposition sensorless\u201d, due to absence of speed and position sensors. It is based on the detection of rotor position which is measured by back-EMF waveform detection in the unexcited phase winding [3]. B. Magnetic, Electric and Geometric Characteristics The used BDCM has three stator winding, 180 W as nominal power and four poles. It was projected to work up to 418.9 rad/s (4000 rpm), continuously without turn off, differently from others motors applied to household refrigeration systems. Geometric and magnetic properties (Fig. 3) of motor were defined, in order to make the analysis of motor magnetic fields, according to constructive characteristics specified in [4]. The steel plates of stator and rotor are constituted of electrical steel Fe-Si. The thickness is 0.5 mm and the height of plate package is 47 mm. It presents losses of 4.9 W/Kg in the conditions of 60 Hz and 1.5 T [4]. Its magnetization curve is shown in Fig. 4. III. FINITE ELEMENT METHOD The finite element method (FEM) is a numeric methods used for finding approximate solutions to partial differential equations, with boundary values given" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002439_0041-2678(70)90116-8-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002439_0041-2678(70)90116-8-Figure6-1.png", "caption": "Fig 6 Flow p a r a m e t e r aga ins t compensa t ion coeff ic ient", "texts": [ "0 The journa l d i a m e t e r is r e l a t ed to the shaf t d i m e n s i o n s and can be chosen acco rd ing to the a v e r a g e p r e s s u r e Pm\" The c l e a r a n c e ra t io ga is n o r m a l l y def ined with l im i t s (0.5 x 10 -3 to 3 x 10 -3) and depends on the r o u g h n e s s of the bea r i ng s u r f a c e s . It has to be r e m e m b e r e d that the lower the c l e a r a n c e the g r e a t e r become both c o and c2; thus , a s s u m i n g the other f a c to r s cons tan t the f r ic t ion coeff icient (see Fig 5) will i n c r e a s e while the lubr icaht de l ive ry will d e c r e a s e . Th i s conc lus ion is not evident in Fig 6 where the flow p a r a m e t e r q is plotted aga ins t the compensa t ion coeff icient c 2. However it i s apparen t that the quanti ty: tgc~ ~ = ftL Q const Q (8) = C2 P 0 d 4 = will d i m i n i s h for b igger va lues of c 2 because of the d e c r e a s e of the c l e a r a n c e g. The g e o m e t r i c a l des ign i s comple ted , when the n u m b e r m (normal ly f ro m 3 to 6) of the e l e m e n t a r y pads is fixed. As a gene ra l t r end the lubr ican t flow i n c r e a s e s with r (see Fig 6) while for spec ia l working condi t ions the load capaci ty and the s t i f f n e s s d e c r e a s e ; the v a r i a t i o n s of the load d i rec t ion however , have a s m a l l e r inf luence (see F ig 7 and Fig 8). The veloci ty coeff ic ient c o i s a des ign cons t r a in t to be dealt with, once both the g e o m e t r i c a l f o rm and the working cond i t ions a re fixed. In o rde r to comple te the be a r i ng des ign the c h a r a c t e r i s t i c s of the supply s y s t e m ought to be e s t ab l i shed , thus two feeding coef f ic ien t s m u s t be given (Equation 3)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001555_iecon.2005.1569184-Figure13-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001555_iecon.2005.1569184-Figure13-1.png", "caption": "Fig. 13. Wheel Model On Slope.", "texts": [ " 14, here we define the coordination system of COM as vertical and horizontal, then the velocity of COM can be calculated as below. { vmx = d dt (x cos \u03b6 + l sin(\u03d5\u2212 \u03b6)) vmy = d dt (x sin \u03b6 + l cos(\u03d5\u2212 \u03b6)) (18) Or { vmx = x\u0307 cos \u03b6 + l cos(\u03d5\u2212 \u03b6)\u03d5\u0307 vmy = x\u0307 sin \u03b6 \u2212 l sin(\u03d5\u2212 \u03b6)\u03d5\u0307 (19) So velocity of COM vG is v2 G = v2 mx + v2 my = x\u03072 + l2\u03d5\u03072 + 2x\u0307l\u03d5\u0307(cos \u03b6 cos(\u03d5\u2212 \u03b6) \u2212 sin \u03b6 sin(\u03d5\u2212 \u03b6)) = x\u03072 + l2\u03d5\u03072 + 2x\u0307l\u03d5\u0307 cos\u03d5 (20) Consequently, no matter how to define the coordination system of COM, the total kinetic energy T does not change. For simplicity, the coordination system is set as Fig. 13. To calculate the potential energy U , similarly we assume that the horizontal plane passing through the rear axle as the potential energy reference plane. U = mgl cos(\u03d5\u2212 \u03b6) + Mgx sin \u03b6 (21) So the Lagrange\u2019s equation can be written as \u03c4 + d\u03b8 = [(M + m)r2 + JM ]\u03b8\u0308 + mlr\u03d5\u0308 cos \u03d5 \u2212mlr\u03d5\u03072sin\u03d5 + Mgr sin \u03b6 + BM \u03b8\u0307 \u2212\u03c4 + d\u03d5 = (Jm + ml2)\u03d5\u0308 + mlr\u03b8\u0308 cos\u03d5 \u2212mgl sin(\u03d5\u2212 \u03b6) + Bm\u03d5\u0307 Similarly to the previous section, Lagrange Criteria on slope is \u03c4 < mgl sin(\u03d50 \u2212 \u03b6)\u2212mlr\u03b8\u0308 cos\u03d50 (22) Let \u03b6 = 0 in (22), it turns out to be the same as (15)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000910_iecon.2005.1569137-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000910_iecon.2005.1569137-Figure3-1.png", "caption": "Fig. 3. Rotor dimensions of two prototypes", "texts": [ " To explain this, analysis of neglecting the cage is compared with experimental values. And by using the model of considering the cage and experimental values, effect of the cage is considered. A. Prototype of single-phase LSPM Fig.2 shows the prototype of single-phase LSPM tested in this paper. The structure of this prototype is that the arc shaped PM is inserted in the rotor of multipurpose single-phase induction motor. So no optimization technique is done for this structure. Two types of rotors are constructed as shown in Fig.3. By comparing analysis of Rotor A with experimental value, the effect of squirrel cage in single-phase LSPM is examined. Furthermore, characteristics of Rotor A and Rotor B calculated by using FEM is shown to discuss the magnetic circuit. Table I shows the main design specification of the prototypes. Material of rotor cage is aluminum, and the number of conductor bar is 28. Concerning PM, arc 138deg magnet (per pole) magnetized parallel is mounted in the rotor. Material is Nd-Fe-B, and residual flux density and coersive force are Br =1", " When this motor starts, starting capacitor is connected in parallel with running capacitor. In steady state, starting capacitor is separated from the armature circuit by Positive Temperature Coefficient (PTC) resistor. 15620-7803-9252-3/05/$20.00 \u00a92005 IEEE In this chapter, the single-phase LSPM analysis that ignored squirrel cage winding is considered. By comparing analysis with experiment, the influence of disregarding the cage can be considered. Fig.4 shows the steady state of phaser diagram ignoring rotor cage. Table II shows the parameters shown in Fig.3. This phaser diagram assumes that main and auxiliary windings are winded orthogonally, so condition of constraint on induced electromotive force (EMF) is AM EE \u22a5 . (1) As voltage equations are written only by the parameters of armature circuit and induced EFM as shown in (2) and (3), effect of the rotor cage isn\u2019t considered at all. ( ) MMMMS IjXREV ++= (2) ( ){ } AACAAS IXXjREV \u2212\u2212+= (3) When load angle \u03b4 is defined as the phase difference between ME and SV , (2) and (3) are rewritten as Fig.5 and Fig", " AM UjU \u03b2= (23) From (19), \u2212 = j U U U A M \u03b2 2 . (24) In this way, MU , AU can be express by 2U . On the other hand, from (12), (13) and (18), 2U becomes \u2212\u2212 += \u03b2\u03b2 A M M A M I jI Z Vj V U 22 1 2 . (25) By substituting experimental values of AMAM VVII ,,, , the effect of the rotor cage 2U can be calculated. And substituting obtained 2U in MU , AU are given. In this way, effectiveness of rotor cage can be estimated. C. Analysis Analysis method discussed below is applied to Rotor A of Fig.3. Fig.11 and Fig.12 shows the positive/negative sequence voltage and current, respectively. To calculate MU and AU from (19), 22 ,VI are calculated. Fig.13 shows phaser locus of MU and AU when changing supply voltage SV As shown in Fig.12 (a), MU is nearly the same phase as SV . So MU acts as if it reduces supply voltage. This means that considering cage winding in analysis makes the amplitude of MI smaller. That is, in single-phase LSPM analysis, considering cage winding affects the accuracy of MI " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000868_1.1792692-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000868_1.1792692-Figure4-1.png", "caption": "Fig. 4 Schematic view of experimental apparatus", "texts": [ " To investigate the influence of the mesh size, three kinds of mesh sizes in the radial and axial directions of the porous rubber block (1035, 20310, and 40320) were used as a trial. The results showed that the difference between the squeeze film characteristics obtained by the mesh sizes of 20310 and those by 40320 is quite small. This suggests that mesh sizes of 20310 are sufficient to correctly evaluate the squeeze film characteristics. 3.1 Experimental Apparatus. The description of the experimental apparatus will be facilitated by reference to Fig. 4. The squeeze film is formed between a cylindrical porous rubber block a and a circular rigid plate b. The porous rubber block is fixed on the lower surface of a steel shaft h, which is supported by an rom: http://tribology.asmedigitalcollection.asme.org/ on 01/10/2016 Term adjusting screw via a load cell. The adjusting screw is installed in a rigid frame and it is able to change the axial position of the shaft in order to arbitrarily set an initial film thickness h0 . The displacement in the radial direction of the shaft is suppressed by an aerostatic bearing g" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002512_s12195-009-0062-x-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002512_s12195-009-0062-x-Figure1-1.png", "caption": "FIGURE 1. Schematic diagram and coordinate system for the 3D simulation of C. reinhardtii. The cell produces forces from two flagella, trans and cis (the cis flagellum is nearer the eye spot, red). Also visible are the large chloroplast (green) and the contractile vacuole (gray, center of the cell body) by which the laser trap captures swimming Chlamydomonas. The cell moves forward along its major axis (z) with velocity |v|. The cell can also rotate with velocities xx and xx about its x and z axes, respectively. The x axis is normal to a plane defined by the major axis of the cell and the minor axis through the eye spot.", "texts": [ "eywords\u2014Computational model, Spectral analysis, Laser trap, Cell motility, Flagella. Chlamydomonas reinhardtii is a unicellular green alga with two flagella that allow swimming motility (Fig. 1). These flagella are located at one end of the organism and pull C. reinhardtii through the aqueous environment. The two flagella beat in different planes relative to each other, causing the cell to rotate approximately twice per second about its long axis as it moves forward.13,19 Chlamydomonas reinhardtii also exhibits phototaxis22; it achieves this via a light-sensitive \u2018\u2018eyespot.\u2019\u2019 The flagellum adjacent to this eyespot is referred to as the cis flagellum, and the other as the trans flagellum", " A computational model of Chlamydomonas was created to quantitatively test the contributions of observed variations in force on phototaxis. The model was realized in Matlab. The computational model is a simple threedimensional representation of Chlamydomonas (Fig. 6). The cell body was modeled as a prolate ellipsoid with a major axis of b = 5.7 lm and minor axis of a = 3.7 lm.13 The cell was assumed to be neutrally buoyant and of negligible mass. This allowed for instantaneous balancing of flagellar forces and translational and rotational drag forces. Figure 1 shows the coordinate system internal to the cell. The cis and trans flagella generate force vectors FT and FC, respectively, at given moments in time. The directional component of the cell\u2019s velocity is assumed to always be parallel with the cell\u2019s major (z) axis; this means that the only translational drag force acting upon the cell is along its major axis. We allowed the cell to rotate about the X and Z axes as well, which result in rotational drag forces about these axes. The drag coefficients are those of a prolate ellipsoid, bTrZ \u00bc 4pg a ln 2b a 1 2 ; bRtZ \u00bc 16pg a2b 3 ; bRtX \u00bc 8pg a3 3 ln 2a b 1 2 where g is the viscosity of the aqueous environment (1 cP), bTrZ is the translational drag coefficient along the Z-axis, and bRtZ and bRtX are the rotational drag coefficients about the Z and X axes, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003732_icinfa.2012.6246839-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003732_icinfa.2012.6246839-Figure4-1.png", "caption": "Fig. 4. (a)indicate two consecutive laser light L1 and L2, (b) when the scan area exist a flowerpot there would be also two consecutive laser light S1 and S2", "texts": [ " While it has a coarse position of the flowerpot, the information can be used to extract the feature of the flowerpot and precise localize it. In this case, there are two works to do. Firstly, from the scan data it can create clusters of neighbor points. Secondly, the clusters are used to the feature extraction procedure which contains the features of lines and circles. A threshold is provided to define the neighborhood. In the experiment, it assumes the effective range of the laser is 10m. The threshold of the two neighbor scan light can calculate as follow (Fig. 4 (a)). \u03b8\u03bb sinL2)L2L1( \u22c5+\u2212< The two laser scanner resolution o5.0=\u03b8 , L1=10m, 0087.0sin \u2248\u03b8 . In our laser statistical experiment there is L1-L2<0.01. Therefore, our laser has the threshold 0187.0<\u03bb . When the flowerpot is in the scan area, the two consecutive points S1 and S2 (Fig. 4 (b)) would have a bigger jump compare L1 and L2 in Fig. 4 (a). After the segmentation, it need to detect the circle of the flowerpot. There is a proved effective technique called Inscribed Angle Variance (IAV) to detect circles [9]. IAV makes use of trigonometric properties of arcs: every point in an arc has congruent angle in respect to the extremes [10]. The detection of circles is achieved calculating the average and standard deviation of the inscribed angles. With the certain a circle has detected in the RFID tag existing area, if there is not ambiguous information such as multi-circle in the same existing area" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001741_icelie.2006.347213-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001741_icelie.2006.347213-Figure1-1.png", "caption": "Fig. 1. Kinematics coordinate system assignments", "texts": [], "surrounding_texts": [ "I. INTRODUCTION\nGenerally, the method of control and guidance of an AGV in the desired path is determined by the properties of the path. To be more specific, when dealing with certain routes, one should choose different control methods from a case in which there should be a plan for path selection. According to such difference, two methods of guidance can be defined: path restricted and free ranging.\nPath restricted control (external control) is done through either wire guided path navigation [1] or control by colored lines. Wire guided path navigation is the most common system for AGV guidance. Free ranging is less considered in an industrial scale. Infrared guidance [2], Inertial guidance and using ultrasonic sensors are the ways in which free ranging methods are implemented.\nThere are two driving systems for an AGV: differential steer in which the AGV uses two wheels on each side (left and right) or one or two caster wheels [3]. In steered wheel driving system a motor rotates the wheels forward and rotation around the axis perpendicular to motion plane is done by changing the angel of the steering wheel [4].\nGenerally we consider two types of path planning: global path planning and local path planning. Global path planning is more applicable in known and static environments. The major benefit of this type of path planning is that by using a complete knowledge about the environment, finding the best path becomes possible. Complete knowledge of the environment however, increases the computational complexity, causing the system to be inefficient in real-time applications [5]. Graph-based methods and genetic algorithms are examples of global path planning.\nIn local path planning a general strategy is used to reach a certain destination but the path is not predetermined. In other words, by using sensors, information about the environment and what is around the AGV is obtained. Position of the destination and position of the AGV and its direction is determined. According to all these and considering the general strategy of movement, the next step can be specified. In unknown environments, local path planning is more efficient. In addition, robot can respond quickly because\ncomputational complexity of local path planning methods is less than global path planning methods [6]. Virtual force field and fuzzy control are examples of local path planning.\nThe biggest problem with the local path planning methods is the problem which is called local minimum. That is when the robot is trapped in a place where it cannot reach its destination. To escape a local minimum, knowing the state of the AGV is essential. The robot is expected to return back to its route toward the destination as soon as it gets out of a local minimum. Also the AGV is supposed not to return to the same local minimum in which it was previously. The later the robot finds out that it is out of the trap, the later it reaches its destination; furthermore, when there is limited space for AGV movements, the AGV has to take many turns which prevents convergence toward the destination.\nOne can think of different methods to escape a local minimum; considering a virtual target, changing the navigation algorithm, changing the parameters of the algorithm, wall following and defining virtual obstacles are some of these methods.\nII. KINEMATICS\nThis paper assumes AGV consisting of one unactuated caster wheel and two conventional differential driving wheels. This type of chassis provides two degrees of freedom (DOF) locomotion by two actuated conventional non-steered wheels and one unactuated steered wheels (i.e., one castor). Robot has two degrees of freedom: y-translation and either xtranslation or z-rotation.\nThe sensed forward velocity solution is (see [8]):\n \n\n \n\n\u2212 \u2212\u2212\n\u2212 =\n \n\n \n\n2 11\n2 1 w ll ll la Ry x w aa bb \u03c9 \u03c9\n\u03b8 , (1)\nand the actuated inverse velocity solution\n641-4244-0324-3/06/$20.00 '2006 IEEE", "( ) \n\n\n \n\n \u2212\u2212 \u2212\u2212\u2212 + = \nBz\nBy\nBx\nabba\nabba\nbw\nw V V\nllll llll\nlR \u03c9\n\u03c9 \u03c9\n11 1 2\n2\n2 2\n1 , (2)\nwhere (in metric system) \u2022 x and y are translational velocities of the robot\nbody [ sm ],\n\u2022 \u03b8 is the robot z-rotational velocity [ srad ],\n\u2022 2w\u03c9 and 1w\u03c9 are wheel rotational velocities\n[ srad ],\n\u2022 R is actuated wheel radius [m ],\nal and bl are distances of wheels from robot's axis [m ].\nConstants al , bl and R should be set according to the robot proportions. Implicit values are 3=al cm, 0=bl cm and 1=R cm. One simulation step represents one second in real-time (i.e. the robot movement in one simulation step is equal to the robot movement during one-second real-time period).\nIII. ASSUMPTIONS\nThe AGV moves in a two-dimensional, static or semi static environment. It is also assumed that the environment is unknown and therefore the path planning is local. The AGV is controlled using free ranging method and a differential driving system moves the AGV which has no memory to store any information. The position of the AGV in each time step and the position of the destination are the only information given to the robot. Since the velocity of the AGV is not very high, its navigation is not substantially affected by its dynamics and therefore, equations mentioned above are used to explain its kinematics. The AGV is equipped with 3 sonar sensors by which the position of the AGV is determined.\nOur goal is that the AGV reaches the destination choosing a path with minimum probability of bumping into obstacles and then minimizing the length of the path.\nThe proposed fuzzy controller consists of two parts: moving toward the destination and obstacle avoidance.\nThe fuzzy controller which generates the virtual force toward the destination accepts the rotation angel (\u03b1 ) and side distance ( s ) as its input parameters. The left or right side distance is calculated considering the direction of rotation of the AGV and the angel between the direction in which it is moving and direction toward the destination.\nFig. 2, 3 and 4 respectively depict membership functions for difference of AGV and destination, Side distance from obstacle and rotation angel. These figures show the\nmembership functions for AGV moving toward the destination. Fig. 5 shows the FIS used for guidance of the AGV toward the destination. Table I contains the fuzzy rules for guidance.\nAs mentioned before, the robot is equipped with three sonar sensors located on front, left and right sides. According" ] }, { "image_filename": "designv11_61_0001766_ichr.2004.1442673-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001766_ichr.2004.1442673-Figure2-1.png", "caption": "Fig, 2. Pelvis-thorax Rotating Angle and Yaw Moment of Stance Foot", "texts": [ " The motion capture system with twelve cameras (Vicon Motion Systems Ltd.) was used to measure three dimensional kinematic data (sampling frequency 120Hz) from reflective markers shown in Fig.l(a). Two 3-axis accelerometers were attached on both iliac crests to measure the antero-posterior and medio-lateral accelerations of the pelvis. The twisting angle of the trunk was measured using four markers shown in Fig.l(b). The thoracic and pelvic rotation around the perpendicular axis, and Bthoras in Fig.2, are measured by the markers on both clavicles and both iliac crests respect.ively. Both angles are set to 0 when the subject is exactly facing the forward direction. The yaw-axis torque exerting from the stance foot to the floor is defined as TLF and T R ~ for each foot\". When TLF increases to the positive and exceeds the maximum static friction, the body begins to rotate clockwise due to the slip occurs at the stance foot. aThe foot rotation about the ground normd is the main focus of this paper", " It is considered that the externally rotated posture of the stance feet and the COP trajectory without passing toes are for the transmission of the torque generated by the hip joints. We have reported the toes (especially the big and 2nd toes) work for balance maintaining of the upper body14. It is assumed the step width becomes wider for the balance. In addition, the translational force at the stance foot can cancel the momentum when the step width becomes wide. Next, the active compensation by the antiphase pelvic rotation is discussed. In Fig.2, TLF increases when the right leg is accelerated and swung forward in the initial part of left leg stance phase. In the normal walk where the leg and the pelvis are in-phase, the momentum due to the increase of BPelvis also increases T L F . In this case, the sum of the momentum should be compensated by trunk twisting and arm swinging. Contrary, in the trunk-twistless walk where the leg and the pelvis are antiphase, the decrease of Opelois cancels TLF. As a matter of fact, the momentum of inertia is not large compared to the legs; total momentum does not be compensated only by this active pelvic rotation" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003825_jzus.c1000224-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003825_jzus.c1000224-Figure2-1.png", "caption": "Fig. 2 Two-dimensional sketch of both inner and outer parameter hopping", "texts": [ " To avoid this situation it has to be assured that |x\u0304gi \u00b7Wgi | < \u03c1i, where \u03c1i is a design parameter determining an external limit for x\u0304gi \u00b7 Wgi . Following the same rationale as the case of weight hopping introduced above, we could now consider two forbidden hyperplanes, which are defined by the equation |x\u0304gi \u00b7Wgi | = \u03c1i. When the weight vector reaches one of the forbidden hyperplanes x\u0304gi \u00b7Wgi = \u03c1i with the direction of updating pointing toward it, a new modified hopping is introduced which pushes the weights back, inside the restricting area. A simplified 2D sketch of the procedure is given in Fig. 2. The size of hopping is given by \u2212\u03bag ( x\u0304giWgi(x\u0304gi) T ) /tr{(x\u0304gi) Tx\u0304gi} and is determined by following the vectorial proof of Theodoridis et al. (2009a), with \u03bag being a small positive number decided appropriately from the designer. By performing hopping when x\u0304gi \u00b7 Wgi reaches either the inner or outer forbidden planes, x\u0304gi \u00b7 Wgi is confined to lie in space P = {x\u0304gi \u00b7 Wgi : |x\u0304gi \u00b7Wgi | \u2264 \u03c1i , |x\u0304gi \u00b7 Wgi | > \u03b8i} lying between these hyperplanes. The weight updating law for Wgi , which embodies both hopping conditions, can now be expressed as W\u0307gi = \u2212 (x\u0304gi) T \u03beiuisi(x)dgi \u2212 2\u03c3i(x\u0304gi Wgi (x\u0304gi )T) tr{(x\u0304gi )Tx\u0304gi } \u2212 2(1\u2212\u03c3i)\u03bag(x\u0304gi Wgi (x\u0304gi )T) tr{(x\u0304gi )Tx\u0304gi } , (29) where \u03c3i = \u23a7\u23a8 \u23a9 0, if xgi \u00b7Wgi = \u00b1\u03c1i and x\u0304gi \u00b7 W\u0307gi <> 0, 1, otherwise" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002821_j.ijsolstr.2008.05.021-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002821_j.ijsolstr.2008.05.021-Figure6-1.png", "caption": "Fig. 6. Example of representation of relative rigid slips adopted in the kinematic approach.", "texts": [ " From (12), the kinematic safety coefficient, adapted to our problem is the following: mk \u00bc min g\u00f0h;b\u00de H cycle R Cc k\u00f0h; b\u00dejg\u00f0h;b\u00dejRdhdb such that G\u00f0h\u00de \u00bc H cycle g\u00f0h;b\u00dedb 2 RSH cycle R Cc qel\u00f0h;b\u00deg\u00f0h; b\u00deRdhdb \u00bc 1 ( 8><>>: \u00f031\u00de where RS is the set of relative rigid slips, i.e. such that G(h) = G for all h. Again, this convex optimization problem can be numerically solved by standard techniques of minimization. This problem is not considered in this paper. In order to obtain a raw estimate of mk, let us restrict the kinematic analysis to the subset of constant piecewise fields of the form (Fig. 6): g\u00f0h;b\u00de \u00bc CA if p=2 6 b < p=2 and p=2 6 h < p=2 CA \u00fe if p=2 < b 6 3p=2 and p=2 6 h < p=2 CB if p=2 6 b < p=2 and p=2 < h 6 3p=2 CB \u00fe if p=2 < b 6 3p=2 and p=2 < h 6 3p=2 8>><>>>: \u00f032\u00de The resulting relative slip G\u00f0h\u00de~e1 is given by G\u00f0h\u00de \u00bc I cycle g\u00f0h;b\u00dedb \u00bc p\u00f0CA \u00fe CA \u00de if p=2 6 h < p=2 p\u00f0CB \u00fe CB \u00de if p=2 6 h < 3p=2 ( \u00f033\u00de A field defined by (32) is thus admissible only if CA \u00fe CA \u00bc CB \u00fe CB \u00f034\u00de Hence, for a Tresca friction such that k(h,b) = k, the kinematic coefficient is given by ~mkT \u00bc min \u00f0CA ;C B ;C A \u00fe ;C B \u00fe\u00de p2kR\u00f0jCA j \u00fe jC B j \u00fe jC A \u00fej \u00fe jC B \u00fej\u00de such that CA \u00fe CA \u00bc CB \u00fe CB CA \u00fe CB \u00fe CA \u00fe \u00fe CB \u00fe \u00bc 1 2pQR ( 8>><>>: \u00f035\u00de Finally, the result of the minimization is ~mkT \u00bc p 2Q k \u00bc p2 1\u00fe St1 St2 Sf1 Sf2 St1 St2 " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002199_6.2007-6345-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002199_6.2007-6345-Figure3-1.png", "caption": "Fig. 3 System of axes used", "texts": [ " Governing Equations Equations governing the relative motion of the pursuer with respect to the target vehicle are along intrack, out of plane and normal axis represented by Wilshire equations1 ( )2( ) sv T sv T sv m = + = + + \u00d7 + \u00d7 + \u00d7 \u00d7 = + = \u03c1r r \u03c1 \u03c9 \u03c1 \u03c9 \u03c1 \u03c9 \u03c9 \u03c1r r Fg g +\u0393r \u03c1 = \u0393 + f(t) (1) where represent respectively the space vehicle, the target position vectors and relative position, trust, gravity vectors, the rotation vectors of the target local frame and pursuer vewhicle mass respectively , , , , ,r rsv T m\u03c9\u03c1 g . The system of axes used is shown on Fig. 3. Equation (1), is linearized, assuming that the trust F is aligned with the pursuer longitudinal axis; expressing the three components of gravity vector g function of pursuer position vector one obtains 2 3 3 2 3 2 (.) ; (.) 2 (.) ; (.) (.) ; (.) 2 x x x y y y z T z z T xFx zf f m r yFy f f m r zF rz xf f m r r \u00b5 \u03c9 \u03c9 \u03c9 \u00b5 \u00b5\u00b5 \u03c9 \u03c9 z x x z\u03c9 = + = \u2212 + + + = + = \u2212 + = + = \u2212 + \u2212 \u2212 + (2) where x, y, z are relative coordinates; \u03c9 is a rotational speed of a frame connected to the target satellite, \u00b5 .represents the gravitational constant" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000147_s0924-4247(03)00089-x-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000147_s0924-4247(03)00089-x-Figure1-1.png", "caption": "Fig. 1. Schematic diagram of a three-DOF four-wire type optical pickup.", "texts": [ " Simulations are conducted for verifying the effectiveness of the controller designed, which show that the designed controller is capable of achieving fast zero tilting with favorable tracking/focusing performances and also robust to plant uncertainty and unbalanced radial vibrations. \u00a9 2003 Elsevier Science B.V. All rights reserved. Keywords: Four-wire type optical pickup; Sliding-mode control Optical disk drives are widely used as basic data-reading platforms for various application drives such as CD-ROM, DVD, CDP, LDP, etc. One of key components for optical disk drives is the optical pickup, which performs data-reading via a well-designed optical system installed inside the pickup. Fig. 1 shows a typical commercial pickup actuator structure\u2014the four-wire type [1\u20133], which mainly consists of an objective lens, a bobbin (the lens holder), wire springs, tracking/focusing coils and permanent magnets. Due to the flexibility of these wire springs, the bobbin could be easily in motions as the electro-mechanical forces generated by the electromagnetic interaction between permanent magnets and the currents conducted in four wires and coils. A typical actuator is designed in a particular dimensional and geometric arrangements of wires, coils and magnets such that the electromagnetic interaction generates two independent actuation control forces, respectively, in the directions of focusing (vertical) and tracking (radial relative to the disk), providing control means on the fo- \u2217 Corresponding author", " Results show that the dynamic behavior of the SMC-controlled system not only acts robust to the disturbances and uncertainty, but also renders satisfactory tracking performance in three different directions simultaneously. This section starts with an establishment of the nonlinear system equations for the pickup motion and its equivalent driving circuit. Plant uncertainty caused by an uneven magnetic field, manufacturing tolerance and the various external disturbances are next formulated for control design. The four-wire type pickup, as shown in Fig. 1, exhibits motions mainly in the directions of tracking (Y-axis) and focusing (X-axis). In addition to the motions in X- and Y-directions, small tilting often occurs about Z-axis, which is due to the geometric mis-pass of the electromagnetic force acting line on the bobbin mass center. This mis-pass is often caused by manufacturing\u2013assembly tolerance and/or an uneven magnetic field. The objective of this study is to design a controller that owns three independent actuation forces/moment in X-, Y- and Z-directions in order to perform precision focusing/tracking and to simultaneously achieve zero tilting to avoid any errors in optical signals" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003987_iecon.2013.6699630-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003987_iecon.2013.6699630-Figure1-1.png", "caption": "Fig. 1. Vector diagram with position and current angle deviations. dqcoordinates are bound to the mover and -coordinates are bound to the armature", "texts": [ " In such a case both current component estimates \u00eed and \u00eeq are erroneous, and therefore, both components will create force components. As \u00eed is not actually on the d-axis it will create some force as well. With the sensorless control the electrical position angle and the current vector angle of the mover are, therefore, erroneous estimates. The actual force with the erroneously estimated \u00eeq can be given as F\u00eeq = KF\u00eeqcos( p ) where p is the error of the electric position angle of the mover, is the error of the current vector angle and \u00eeq the estimated q-axis current. The vector diagram of the definition is presented in Fig. 1. Because the phase shift between the d- and q-axis equals to /2 there will be also a lateral force component produced by the erroneously estimated \u00eed F\u00eed = KF\u00eedcos( p ( /2)) ( ) Naturally, the real d-axis current does not produce any force if it happens to be estimated correctly ( p = = 0). In linear movement a lateral force is needed for accelerating the mover and the load, and for overcoming the friction and cogging forces. At high acceleration levels the highest demand for force is formed by the acceleration of the mass of the mover and the load" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003372_978-3-642-25486-4_28-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003372_978-3-642-25486-4_28-Figure6-1.png", "caption": "Fig. 6. Unilateral Contact Model in Parallel Indexing Cam Mechanism", "texts": [], "surrounding_texts": [ "In one automated machinery indexing cam system is combined by motor, indexing cam, roller and turret and rotary table as Fig. 4. Contact between cam and rollers of turret is the key factor affecting the system performance and is dealt by Unilateral contact model. Unilateral contact model denotes a mechanical constraint which prevents penetration between two bodies. For parallel indexing cam mechanism, clearance between rollers and cam profile can be divided into three situations: Fig. 5 (a) presents that roller penetrated into cam profile, contact force is defined by complaint distance and curvature radius, roller radio, material property, oil film thick between roller and cam; (b) presents roller contract with roller justly. In his case, no compliant exist between the two ones and so no force generate. (c) Means that roller is separating with cam profile. No contact and no force exist. So we can obtain the unilateral contact model. Contact force can be calculated by stiffness and displacement based on Hertz law if compliance between two bodies exists. The relationship between contact normal force and contact displacement is shown as follows. 2 3 \u03b4kP = (6) Where 9 16 2RE k = 2RRR 111 1 += , 2EEE 2 2 1 2 1 )1()1(1 \u03bc\u03bc \u2212+\u2212= 1\u03bc , 2\u03bc are the Poisson\u2019s ratio of two cam and rollers; E1, E2 are Young\u2019s Modulus of two bodies, R1, R2 are curvature radius of contacting surface of two bodies. More generally, normal contact force between cam and roller can be written as e\u03b4kP = (7) Where stiffness value k and force exponential value e can be determined by material characteristics and geometry dimensions [13]. Normal contact force nF21 , damping force cF21 and friction force fF21 should be considered between cam1 and turret 2 during contact process. So the total contact force is '''21212121 xPycyPFFFF fcn \u03bc\u2212\u0394+=++= (8) Where c is the damping ratio, and \u03bc is the friction coefficient. According to kane equation, the differential equation of turret can be written as 0)( 6 1 =+++ =i fifididicici lFlFlFI\u03c6 (9) Where I is inertia of turret, \u03c6 is angular displacement of turret, i presents the ith roller, ciF is contact force of roller i, diF is damping force of roller i, fiF is friction force of roller i. Integrate with other dof equation we can obtained whole system dynamics equation (6)." ] }, { "image_filename": "designv11_61_0000127_s0921-4534(03)01215-2-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000127_s0921-4534(03)01215-2-Figure1-1.png", "caption": "Fig. 1. Mechanism of the rotation speed degradation of the SMB.", "texts": [ " Against the prediction of advantage (i), however, it has been found that even the SMB has the remarkable rotation speed degradation by the electromagnetic phenomena. Purpose of this research is to evaluate the rotation speed degradation of the SMB by both electromagnetic field and heat conduction. At first, we supposed the cause of the rotation speed degradation by the SMB is only the magnetic force interaction between the inhomogeneous magnetic field of the PM rotor and the eddy current induced in the cryostat (cause-A), or the shielding current induced in the HTSC (cause-B) as in Fig. 1. But the characteristics of measurement results were against the prediction derived from cause-A and B. Namely, the degradation remarkably enhanced when the levitation force of the SMB increases [2]. Such an enhancement cannot be occurred by cause-A and B, because the degradation by the inhomogeneous magnetic field of the PM rotor should be independent from the levitation force [3,4]. For explaining such anomalous characteristics of the rotation speed degradation by the SMB, we noticed the shape of the HTSC stator. It consists of the cryostat and six HTSC bulks as in Fig. 1. The magnetic field distribution made by the six bulks is periodic in the circumference direction in each 60 , because the shielding current dose not flow across the boundaries between HTSC bulks. Looking from the rotating PM rotor, the magnetic field becomes like the AC field. As a result, the eddy current is induced in the PM rotor by this magnetic field. So this is the third mechanism of the rotation speed degradation (cause-C). The rotation speed degradation of the SMB is especially the matter of the great importance to the development of the flywheel system" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003735_2011-01-2652-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003735_2011-01-2652-Figure5-1.png", "caption": "Figure 5. Traveler", "texts": [ " SwRI determined that nine degrees of freedom are required for full accessibility of the robots around a typical fighter aircraft. The first two degrees of freedom (Axes 1 and 2) are used for gross positioning of the robot's vertical column; with the remaining seven axes providing coordinated motion of the end-effector around the aircraft. This allows approximately 95% of the fighter's surface area to be stripped. Following is a more detailed description of the robot, by axis. Traveler (Axis 1 Translation) The primary structure of the robot is referred to as the Traveler (Figure 5). It is equipped with four steel wheels which ride on a track system which runs the length of the hangar. This movement, referred to as Axis 1, provides gross positioning of the robot along the long axis of the aircraft fuselage. In addition to providing support for the upper and lower beams (described below), the traveler also houses the media blast pot and the control systems for the robot. Axis 1 is driven by a servomotor/gear-driven pinion that engages a long gear rack that is welded to one of the floor tracks" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003421_48168-Figure9-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003421_48168-Figure9-1.png", "caption": "Figure 9. Scheme of the elliptically excited mathematical pendulum of length l. The pivot of the pendulum moves along the elliptic trajectory (dashed line) with semiaxes X and Y in the uniform gravitational field g.", "texts": [ "67 the transition to chaos through subcritical AH bifurcation is the most typical. In the middle of Fig. 8(d) the manifold of dark blue points reveals a typical strange attractor structure. The strange attractor inherits the basin of attraction from disappeared stationary attractor. Elliptically excited pendulum (EEP) is a mathematical pendulum in the vertical plane whose pivot oscillates not only vertically but also horizontally with \u03c0/2 phase shift, so that the pivot has elliptical trajectory, see Fig. 9. EEP is a natural generalization of pendulum with vertically vibrating pivot that is one of the most studied classical systems with parametric excitation. It is often referred to simply as parametric pendulum, see e.g. [9, 22\u201327] and references therein. Stability and dynamics of EEP have been studied analytically and numerically in [28\u201330]. Approximate oscillatory and rotational solutions for EEP are the common examples in literature [31\u201334] on asymptotic methods. Sometimes EEP is presented in a slightly more general model of unbalanced rotor [31\u201333], where the phase shift between vertical and horizontal oscillations of the pivot can differ from \u03c0/2" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001062_bfb0031456-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001062_bfb0031456-Figure1-1.png", "caption": "Figure 1: Error Function Formulation", "texts": [ " , 3 (3) where xpj is a cyclic permutation of the x coordinate of each of the three support points on the platform for the legs and pj is the distance between the leg support points on the platform. The error function was obtained by explicit differentiation, and as can be expected, the method proved to be slow. However, as mentioned above, the method proved to be highly robust to the platform passing through mechanism singularities, which we conjecture to be because of the presence of derivative information in the algorithm. The final method considered was based on an algorithm suggested by Rooney and ReesJones. Consider the platform shown in figure 1. Given a set of leg lengths then each of the points C, D, and E will lie on circles with centres F, G, and H respectively whose equations may be obtained explicitly. Now consider the points C and D when D has been specified by choosing a value for the auxiliary leg pair angle 0~. The point C may be calculated using trigonometry and the constraint that C is distance dp from D. Similarly, the position of E may be so calculated, however, the distance between C and E will not be equal in general to the leg spacing dp but will be in error by an amount e" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002122_05698196908972260-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002122_05698196908972260-Figure2-1.png", "caption": "Fig. 2-Squeeze film pressure prafiles for decreasing clearance.", "texts": [ " The flow conditions are expressed as W FP = W pB3Lh FP~ 185 [9] The known supply conditions are expressed as follows. The flow through a compensating element can be expressed by N' N' [ N' Jr; =~ Q i = ~ ~ (Xigii + Ago\\ i J = 1 J =1 1=1 N = No. of recesses fed by Fr r = 1, Nf Substituting [4] for Q i and solving for Psi gives t NR j llS j ~1 (XiQij + AQii Psi = + (X j . jj S, = PSj r = Nf + 1, NR [5] [6] [7] [8] where the dimensionless flat-plate squeeze load, WFP , is a function of L/ B. The type of pressure distribution in the fluid film associated with this squeeze load is shown, Fig. 2a , for some arbitrary section perpendicular to an edge. Also shown, Fig. 2b, is the effect of having a recess in a flat plate. For the simple case of parallel flow the pressure distribution over the small-clearance area is the same for Figs. 2a and 2b. Since the flow at any two respective points is the same by virtue of continuity, and since pressure induced viscous flow requires that the fluid velocity be proportional to the pressure gradient (i.e., slope of pressure profile), the pressure profiles across the respective land regions are identical. Since the squeeze load is essentially the integration of this pressure distri bution over the given surface, one readily sees that a flat plate with a recess has less damping capacity than the same flat plate without a recess. Now consider the effect of linking the recess to ambient pressure through a flow resistor (orifice or capillary) Fig. 2c. Since the squeeze flow now has an additional \"escape route\", the squeeze flow portion over the sills is reduced from that of the closed recess (R = Equations [5] and [8] constitute a set of NR equations in the unknown (Xi which can be solved by methods described in (1). RECTANGULAR FLAT PLATE WITH RECESS AND RESTRICTOR EFFECTS The solution for the squeeze film load capacity of a flat rectangular plate is given by Hays (2) . The results of (2) are in complete agreement with the solution of the flat plate problem using the computational scheme O", " Summarizing, it ca n be said that the da mping capac ity (squeeze load) is associated with two changes to the pad pressure distribution as observed und er static conditions, namely, 1) increased recess pressure, and 2) change to the relative shape of pressure distribution across the sills. One of the first things established by the various computed results of this study is the following relation ship for the plane rectangular recessed pad with com pensation. (0 ) thus reducing the pressure gradient across the sills and also reducing recess pressure. Figure 2c shows the pressure distribution trend as the restrictor resistance goes from infinity to zero. The addition of the restrictor is therefore shown to further reduce the squeeze load below that for a flat plate. The implication here is that the damping capacity of a flat hydrostatic pad with recess(es) is always less than that for an equally sized flat plate of the same shape operated with the same parameters. Computed results will show this to be true regardless of static pressure ratio, supply pressure or compensation type" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000540_105994905x75475-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000540_105994905x75475-Figure1-1.png", "caption": "Fig. 1 Titanium melting processes; (a) Vacuum arc remelting (VAR); (b) Plasma arc melting (PAM)", "texts": [ "00 Journal of Materials Engineering and Performance Volume 14(6) December 2005\u2014697 the result of more expensive alloy additions that make up 25% of the total composition of Beta C. One possible method to reduce the overall cost of Beta C product is to single-melt the alloy using plasma arc melting (PAM) to a smaller diameter ingot than is used in conventional double-VAR melting. Cost reductions utilizing this processing route include (a) more flexible use of raw materials and (b) significantly fewer hot working and conditioning steps to prepare the material for input to a bar-rolling mill. A comparison of the VAR and PAM melting processes is shown in Fig. 1. The PAM process permits a more flexible use of various forms of low-cost input materials and requires only one melting operation whereas, the VAR process needs at least two melting operations to assure homogeneity. Fewer hot working and conditioning steps are required during processing of the PAM material, resulting in a higher yield from ingot to bar rolling input. The processes for making bar from VAR and single-melt PAM are schematically shown in Fig. 2. The single-melt PAM process allows many processing steps to be eliminated" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000915_bfb0015079-Figure1.2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000915_bfb0015079-Figure1.2-1.png", "caption": "Fig. 1.2. Modes of oscillation of a three element biped", "texts": [ " If the ground surface is included, the third mode of behavior is ruled out since this leads to the collapse of the penduhm. When we add a second link to the simple system that we have considered and coordinate the motion of various members by applying appropriate joint moments the dynamic behavior remains similar to we have observed in the single member case [16]. In the presence of the walking surface, the system may still operate in either mode A or B, as long as all the parts of the system remain above the ground, Fig. 1.2a and Fig. 1.2b. When the system is in mode C then the swing limb will contact the ground and prompt a chain of events that may lead to stable progression. The importance of this contact event can be better understood if the motion is depicted in the phase space of the state variables. We simplify the present discussion by describing the events that lead to stable progression of a biped for a single degree of freedom system, however this approach can be generalized to higher order models. The phase plane portrait corresponding to the dynamic behavior that is described in the previous section is depicted Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001376_iros.2006.282371-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001376_iros.2006.282371-Figure3-1.png", "caption": "Fig. 3. Region where pk has to be placed on the object boundary in order to obtain a FC grasp, considering the intersection between F i, and F j, . The unitary vectors that bound the friction cone at any point between pi\u2032 and pj\u2032 spans the force space together with i and j , and sij1 lies inside the friction cone at any point between pl and pr ; the intersection of these regions determines where pk has to be placed.", "texts": [ " Proposition 1: (from [23]) Three contact points pi, pj and pk allow a FC grasp if and only if: (a) the unitary primitive vectors that bound the friction cones at these points positively span the force space, and (b) at least one intersection point between the supporting lines of two primitive forces lies inside the double-side friction cone at the other contact point. Fig. 2a shows an example of three contact points that satisfy the necessary and sufficient condition in Proposition 1, allowing a FC grasp, and Fig. 2b shows an example of three contact points that do not satisfy the condition. From Proposition 1 the following two Lemmas can also be stated (illustrated in Fig. 3). Lemma 1: Consider two contact points pi and pj , and let sijm be the intersection point between the straight lines Fi,c and Fj,c (remember that c\u0302 can be either r\u0302 or \u0302 ). In order to obtain a FC grasp, the third point pk must lie in the intersection of the following two regions on the object boundary: \u2022 The region of points where the unitary vectors that bound the friction cone together with the unitary vectors of the primitive forces that determine sijm , positively span the force space. This region is bounded by the points p i\u2032 = Op(pc i ) and pr j\u2032 = Op(pc j) (note that this region is always continuous when the object is convex, while it can be discontinuous when the object is concave)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002567_j.elecom.2008.04.019-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002567_j.elecom.2008.04.019-Figure2-1.png", "caption": "Fig. 2. (a) Schematic of the type of CFE used in this study. FESEM images of a CFE at (b) low and (c) high magnifications.", "texts": [ " the carbon sensor reference electrode during a \u2018\u2018wait\u201d period (0\u201310 s) and \u2018\u2018incubation\u201d period (10\u201330 s) and 300 mV during the \u2018\u2018detection\u201d (30\u201360 s) period (to detect Fe\u00f0CN\u00de4 6 at a diffusion-limited rate), while the CFE potential was 350 mV vs. the same reference electrode, throughout, to detect Fe\u00f0CN\u00de4 6 by oxidation to Fe\u00f0CN\u00de3 6 . An optical micrograph of the CFE positioned in a droplet on the Assure I sensor is shown in Fig. 1b. Note that for both sensors, the CFE had to be inserted through a fine mesh that acted to retain the applied droplet; no modifications were made to the sensors. Fig. 2a shows a schematic of the CFE, while Fig. 2b and c highlight the conformal polyoxyphenylene film, resulting from the fabrication procedure. Cyclic voltammograms recorded at these CFEs in a test solution of 1 mM Fe\u00f0CN\u00de4 6 and 0.2 M Sr(NO3)2 (supporting electrolyte) gave very well-defined responses with diffusion-limited currents of 1.5 nA (Fig. 3), corresponding to the one-electron oxidation of Fe\u00f0CN\u00de4 6 to Fe\u00f0CN\u00de3 6 . This current response is reasonable for the geometry and dimensions of the electrodes used herein, where the thin insulation sheath may lead to slight protrusion of the electrode and significant back diffusion from behind the electrode plane [19,20]" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003757_siitme.2013.6743683-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003757_siitme.2013.6743683-Figure2-1.png", "caption": "Fig. 2. The scheme of the electro-magnetic actuated vibrating platform", "texts": [], "surrounding_texts": [ "978-1-4799-1555-2/13/$31.00 \u00a92013 IEEE 241 24\u201327 Oct 2013, Gala\u0163i, Romania\ninduced in the body while the person sits upright on a high frequency vibrating platform. Multiple meta-analyzes have shown growing interest in this type of therapy, not only to fortify muscles, for the prevention and treatment of bone decalcification (osteoporosis), but also for patients suffering from certain chronic diseases [1]. The vibrations of the platform are specific for each patient because the frequency and amplitude of the oscillations are directly related to the person's weight. The original system, which was patented [7], was an electromechanical one with limited possibilities for further research. This paper presents a low frequency electromagnetic actuated vibrating platform and the corresponding electronic control system which automatically corrects the vibrations frequency through a feedback loop. Frequency stability in this case is \u00b1 0.5 Hz.\nIndex Terms\u2014 Oscillator, electro-magnetic actuator,\nvibrating therapy.\nI. INTRODUCTION\nVibration therapy (Whole Body Vibration [WBV] and Dynamic Motion Therapy [DMT]) has provoked an important scientific interest and gained popularity especially regarding its impact on bone density, as mechanical loading is one of the forms of inducing osteogenesis but also for fitness. The vibration platform presented in the article works on the principle of Dynamic Motion Therapy as in Table 1 [2, 6].\nThe Whole Body Vibration (WBV) is also known as Biomechanic Stimulation (BMS) and was invented by the Russian scientist Dr. Vladimir Nazarov, who researched for the Soviet gymnastics team and used the technology as physical therapy to help the cosmonauts overcome the considerable decrease of bone mass and muscle tone that they risk during extended periods of weightlessness. The effects of vibrations on the human body (prevention and treatment as well) have been researched and applied and nowadays it has worldwide a huge diversification of applications. This is the reason why the authors are so interested to continue the research and improvement of the solutions.\nLow vibrations simulate the action on the muscle cells during daily activities like standing, maintaining equilibrium and walking. These perform a sequential contraction, exerting many small combined loads on the bones and enhancing their consolidation. According to the researches, these low\nPosition Resonance frequency\nBody part Frequency value\nupright\nEye 20 Hz\nHead 18 Hz\nMuscle 7-15 Hz\nInner organs 8 Hz\nSpine 8 Hz\nWhole body 5 Hz\nThe fact that Table I presents the resonance frequencies for certain body parts and some WBV-devices work exactly with this frequency might look contradictory, but the frequency for the different technologies are applied for a short time.", "2013 IEEE 19 th International Symposium for Design and Technology in Electronic Packaging (SIITME)\n978-1-4799-1555-2/13/$31.00 \u00a92013 IEEE 242 24\u201327 Oct 2013, Gala\u0163i, Romania\nAs in [5], the frequency of 5 Hz used in case of stochastic resonance is on the lowest level of what training experts expect to be functional.\nII. THE VIBRATING PLATFORM MODEL\nAs in Fig. 1 and according to [7, 8, 9, 10], a model of a vibrating platform was developed, based on an electromagnetic actuator. The vibrating platform consists of a rigid plate 1, articulated on one side to the frame 2, an elastic element 3, an electromagnetic actuator consisting of a moving core 4 and a fixed coil 5, an electronic control system 6 and a vibration transducer 7.\nelectro-mechanic actuator\nThe weight G of the person standing on the plate is compensated by the opposed elastic force F2. The electromagnetic force F1 induced by the movement of the mobile core inside the fixed coil of the electromagnetic actuator produces the plate oscillations in a vertical plane and the vibrations, with a required frequency and amplitude, are transmitted to the bones and to the muscles of the patient. The vibrations generated by the platform are always chosen according to the training or treatment objective.\nIII. THE CONTROL SYSTEM\nThe first attempt was to use an AC electromagnetic actuator, which was extremely noisy and provided a step signal. The solution was to choose a DC electromagnetic actuator, even more adequate for biomedical devices.\nThe platform and the body parts motion during vibration are monitored by accelerometers (vibration transducer). One is placed on the platform and one is placed on the measuring point on the body.\nThe developed and built electronic control system provides a stable frequency, controlled through a loop. Fig. 3 shows the block diagram of the electronic control system. The frequency of the platform oscillation where the person stands vertically is picked up by the vibration transducer and converted into a continuous voltage. The system contains a thermally compensated reference voltage supply, which sets the initially system frequency. The DC voltage, as result of the frequency conversion, is compared with the reference voltage by a comparator system. By means of the comparison result \u2013 which is a voltage \u2013 the frequency of the oscillator is corrected, so that the error is minimal.\nTherefore, an oscillator provided with a power amplifier is used, which keeps a constant fill factor of \u00bd, regardless of", "2013 IEEE 19 th International Symposium for Design and Technology in Electronic Packaging (SIITME)\n978-1-4799-1555-2/13/$31.00 \u00a92013 IEEE 243 24\u201327 Oct 2013, Gala\u0163i, Romania\nthe frequency and the weight of the person on the platform. Fig. 3 shows the electronic scheme of the control system.\nThe novelty of this type of low frequency oscillator consists in the fact that the pilot oscillator realized with the operational amplifier U2 uses the charging and discharging of the low capacity capacitor C4 by means of two current\ngenerators. The supply of the oscillator is realized by an amplifier U1, which compares the reference voltage with the result of the frequency-voltage conversion. The presented electronic circuit is able to command high power electromagnetic actuators.\nTherefore, an oscillator provided with a power amplifier is used, which keeps a constant fill factor of \u00bd, regardless of the frequency and the weight of the person on the platform. Fig. 3 shows the electronic scheme of the control system.\nThe novelty of this type of low frequency oscillator consists in the fact that the pilot oscillator realized with the operational amplifier U2 uses the charging and discharging of the low capacity capacitor C4 by means of two current generators. The supply of the oscillator is realized by an amplifier U1, which compares the reference voltage with the result of the frequency-voltage conversion. The presented electronic circuit is able to command high power electro-magnetic actuators." ] }, { "image_filename": "designv11_61_0001745_epe.2005.219327-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001745_epe.2005.219327-Figure4-1.png", "caption": "Fig. 4: Experimental device", "texts": [], "surrounding_texts": [ "Qiw4hjaive tncEvd. ofilfccItIr.al raild ngInc[ic field for Jic.chon of ,4atorhilerA Thaoi-chdrt in inducioc mchic.U\n1:Pks,kr ,cob'S (t.as) = 2 \u00a3h c0s{[I + krNr (1 - s)]w t - (Its + ksNs + krNr)pas} hs,ks,kr\n+-ks k2kr 2 cos {[1 + krNr (I - 8)] w t - (h + p ksNs + p krNr) a - hA} (6)\nwitli ks niow- varying froni -c to +Dx. It can be noticed that cach colimpollerlt can be written with tilc following form:\nb7 -IH (t, aS) =bKH- cos (K w, t - H a') (7)\nOne can observed that the components induced by the short-circuit correspond to the second sunilniation of the relation (6) and are at [I + krNr (1 - s)] an angular frequencies. So, as these angular frequencics already exist for a hcaltlhy niachiric, the inagilitude of these ones are niodified when the fault appears.\nExternal radial magnetic field The externaf magnetic field is meastured with a coil, use(d as a flux sensor, wlhich delivers an cn.lm.f. The ancalytical cxpression of this cxtcrnal radial magnetic field can be deteriiined fromi the airga.p fluix density given by (6). Previous study [8] has shown that the radial flux (lensity componienits at the outside of the machine are linked to the airgap flux density ones by a coefficient taking into account the decreasing of the magnitude of eacli component tlhroughi the stator frame, and in the air. So, the component ba'2,H at the external of the machilne can be written as\nb\u00a7sK, = blk,H (8)\nThe coefficient AC,, (lefiie(1 from the Maxwell equations, is expressed by (9) anil depends on the pole pair iurniber H of each of the compornents which contribute to the various spectral lines at Kfw frequeicy. This coefficient is also fhinction of the relative permability Itr of the stator frame, of the interior and exterior stator radii, respectively Rsirt and Rs,,t. r represents the distance from the motor axis to the sensor put at the outside of the miachinc.\nIC= 2 ( r ) (9) PrK~b1L\\-IHI-I ( +HI]()\nKnowing the sensor characteristics, as well as the coil surface area. the numbers of turns and its frequency range, the flux linked by the sensor can be calculated. One component 'qb- at Kw angular frequency is obtained by summing up all the comnponcints oii H. Th1Cni thc c.rn.f. eJc(r), sent by the flux sensor, is given by\ne/v(r) =-- r with OK(r) =E JF7i btH (10)\nwhere TFH is a coefficicit, depending on the flux sensor charactcristics, which results from the integration of the flux density on the device area. So, it is possible to predict the frequenicy of the different spectral lines of the external magnetic field but, also, to have a good idea of the niaginitude of each of the conipoiieiit which coImlpose thesc spectral Iiies.\nISBN: 9075815-OS-3\nBRUDNYJ,II,,,,cO,\nPE 200t; - Dle,dai P.4", "Quaantitative analysis of the external radial magnetic field for detection of stator inter-ta short-circuit in induLction macitines.\nStray flux measurement and short-circuit fault The experimental mnachine is a 11kW, 380/660V, 50Hz iniduction machine with p = 2, m = 4r Ns = 24 and Nr = 16. The arrangement of the experimental machine winding is given in figure 3.\nOn this muachine, all the elementary sections are extracted as shown by the picture 4(b). Consequenitly, it is possible to performn a transformationi on the stator winidiing of the machine to simulate a short-circuit fault. A short-circuit of a whole elementary section represenits 12,55% of the comnplete winding of one phase. For the experimentations, the sensor is put against the stator frame, as shown on figure (4(a)) ill order to mailnly measure, at a fixed poinit, the radial compolnelnt of the external inagnetic field which flows in the surrondinig of the machine. The coil selds, its induced e.m.f. to a 16-bits spectral analyser, what allows to obtain spectra with a good accuracy. The supply conditions of the machine, when the fault is created, are adjusted to obtain a shortcircuit current not so important to not cause damages on the machine.\nExperimental results When a short-circuit fault exists on the machine, the wlhole spectrum of the external magnetic field is modified, and consequently, also the whole spectruin of the e.in.f., as illustrated by the\nEPE 2005 - Dresden ISBN: 90-75815-08-5 P.S\nBRtUDNY Jean Fran(;Ois\n2005 Dresden ISBN: 90-7-5815-03-5e P.5", "QLlaUtitatLve anulysis of the externa radial manetic fild for detction of sator inter-tunt short-cixruit ill iluctionL mallcllincs.\nspectra in figure 5 whre Ice = 101 (P = 3A)} The tests present a running at nio load. The spectra are given in relative value (in dB) regarding to the fundamental one which corresponds to OdD.1\nIt cani be observed that, the sensitivity of the 750 and 850Hz spectral lines, to tie shiort-circuiit fault, is imnportant. Tllese spectral lines correspond, for tlle considered miach-Iine. to tlle first slotting harmonics withi a frequencie defined by (6) for kr- \u00b11. So, the study takes an interest in the quantitative approach of the evlolution. when a slhort-circuit fault appers. of this particular spectral linles.\nMagnitude of the slotting spectral lines depending on the slhort-circuit location\nThe figture 6 presents, for a fixed position of the sensor. the evolution of the first spectral lines of the e.m.f., given l)y tlhe sensor. in rel]ative value (in (11) regaxding to the fundlamiiental., on the lhealthiy iniachiue arid whien a short-circuit is crcated witli a faulty cturreint eqiul to 10 timrecs the sttppvy cutrrent value.\nTlle abeissa represents the 24 elementaLry sections of the experimental niiac-Iine. As shown. on figuire 3. the cleicicntaxy sections 1. 2. 3. 4 and 13, 14, 15. 16 correspond to the first phiase the sections 6, 7, 8 and 179 18, 19, 20 to tlhe second one and tlhe others, to the third phase. It can lbe firstlxr seen that tlle b)oti slotting spectral lines evolve in tlhe sane way whien the location of the fauilt chaniges around the rnachinc. Ono can also observe thlat tho mnignituldCs\nISBN: 90-75815-O&-5 P.6\nBRLrDNfY Jeaa Fa~uiOis\nEPE9l0M5 - Drsdeil IS 90-758 -06-5 P.6" ] }, { "image_filename": "designv11_61_0003564_iccda.2010.5541299-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003564_iccda.2010.5541299-Figure3-1.png", "caption": "Figure 3. Grids of computation zone", "texts": [ " With the assuming of ideal gas and isothermal conditions, the equation is given by: \ufffd(ph3 ap)+\ufffd(ph3 ap) ax 12,u ax Oy 12,u Oy =\ufffd(phU )+ a(ph) ax 2 at (1) Where h is the thickness of gas film between faces. ,u is the gas viscosity. The left items in (I) denote the change of film pressure along with the coordinate x, y, and the right items in (1) present different dynamic effects of film pressure. Flow inlet is Pressure-inlet, flow outlet is Pressure-outlet, two sides interface are periodic, and the end faces set as waIl (Fig 2). Use unstructured grids, the grids of computation zone are shown in Fig 3. Because the size in Z-axis is so smaIl, the model was divided into two main zones (detailed in Fig 3) to get higher grid quality, and near the area of conjunction, finer grid is needed. In the calculation process, the X-axis direction nodes from 60 to 90, he Y-axis direction nodes from 50 to 60, and the Z-axis direction nodes from 30 to 60, ensure the V5-228 Volume 5 difference of opening force less than 1 %, meshing quantity is about 220000. Use SIMPLC discretization methods, the flow fields are described vividly, and the working principle also analyzed. III. NUMERICAL SIMULATIONS A. Pressure and path lines The numerical results in Fig 4 show the distribution of static pressure on whole mating ring face" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002885_wnwec.2009.5335810-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002885_wnwec.2009.5335810-Figure4-1.png", "caption": "Figure 4. FEA modeling when considering bearing clearance", "texts": [ " ANALYSIS OF GEAR PAIR CONSIDERING BEARING CLEARANCE There are three type bearings used mounted at the shafts in the high-speed gear pair: NCF****, NJ****, 3****. The clearance of bearing NCF**** at the two sides of wheel in hollow shaft is 225 \u03bcm and the clearance of bearing NJ**** at the left side of pinion in output shaft is 70\u03bcm. There is only axial internal clearance and the radial internal clearance of the bearing 3**** is zero. It is assumed that bearing 3**** is fixed with the gearbox here. The FEA modeling of the gear pair including the shafts and bearings is shown in Fig. 4. As shown in Fig. 5 to Fig. 9, the maximum contact stress among these five gear teeth is 933MPa, occurred at the contact line of gear tooth No.3 when considering bearing clearance. While the maximum contact stress is 1437MPa occurred at the contact line of gear tooth No.1 when not considering bearing clearance. In Fig. 5 to Fig. 9, (a) denote the contact stress when considering bearing clearance, while (b) denote the contact stress when not considering bearing clearance. By comparing the maximum contact stress at each gear tooth in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003611_amr.97-101.2119-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003611_amr.97-101.2119-Figure2-1.png", "caption": "Fig. 2 Helical drill point grinding method Fig. 3 Helical drill point grinding using the biglide parallel grinder", "texts": [ " In this paper, a biglide parallel grinder [4] (as shown in Fig. 1) is presented. During helical drill point grinding process with this biglide parallel grinder only two axes simultaneous motions are required instead of three axis simultaneous motions of conventional tool grinders . Moreover, the relationships between the helical drill point parameters and helical grinding parameters are derived. As a result, the flank surfaces of helical drill points are described. Helical drill point grinding is shown in Fig.2. The angle between drill axis and grinding wheel axis is \u03b8, semi-point angle. There are three axis simultaneous motions when helical drill points are ground Advanced Materials Research Vols 97-101 (2010) pp 2119-2122 Online: 2010-03-02 \u00a9 (2010) Trans Tech Publications, Switzerland doi:10.4028/www.scientific.net/AMR.97-101.2119 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", " Because of using parallelograms of the four links, the moving platform always moves in the plane that is vertical with the base platform whenever the biglide parallel grinder works. So, a reference coordinate system (X, Y, Z) is attached to the center O of the base platform, as shown in Fig. 3. The angle between drill axis and grinding wheel axis is \u03b8 (not shown in Fig. 3). When the glider 1 keeps stop on position of the center O and glider 2 glides along the direction from its original point to O, the drill point will move along a arc from P0 to P. The arc can be looked upon a composition of V1 and V2 in Fig.2. It means that the movement of one of two gliders will leads to two movements, V1 and V2, of mobile platform. In other words, during helical drill point grinding process with this biglide parallel grinder only two axis simultaneous motions are required instead of three axis simultaneous motions of conventional tool grinders. From Fig. 3, Fig. 4 and Fig. 5, the relationships between any point (point P) of flank surfaces of the helical drill point and helical drill grinding parameters are derived as follows: \u03c9 the angular speed of drills, t time of both glider 2 gliding and the drill rotation, r the radius of the moving platform, h the distance between drill points and the moving platform, L the original distance between glider 1 and glider 2, l the length of the parallel links, and R the radial distance between point P and the drill axis" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003364_978-94-007-5125-5_1-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003364_978-94-007-5125-5_1-Figure4-1.png", "caption": "Fig. 4 Model of wheellegged robot suspension", "texts": [ " The aim of the studies is to design a suitable control system ensuring the automatic maintenance of a constant elevation of the chassis above the ground while travelling on a bumpy surface. In the case of the computational robot model, the levelling function boils down to keeping a constant elevation of the robot chassis above the ground according to the schematic shown in Fig. 2b. This function can be effected solely by lifting the chassis by means of lift actuator qp (Fig. 3) while the other drives (protrusion, turn and rolling) remain fixed. First, simulation studies of chassis lifting were carried out. A schematic of the simulation is shown in Fig. 4. The aim of the studies was to determine the dependence between chassis elevation hk above the ground and lifting actuator extension qp. Location zA of suspension rotational couple A (Fig. 3) was assumed as chassis elevation hk. Figure 7 shows the obtained graphs of chassis elevation hk and the change in active force Fp in the lifting actuator versus lifting actuator extension qp. The range of change in actuator extension qp is determined by the basic suspension specifications shown in Table 1. For such a actuator its chassis elevation hk was found to range from 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000589_0304-8853(80)90556-9-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000589_0304-8853(80)90556-9-Figure1-1.png", "caption": "Fig. 1. Magnetization curves under tensile stress, J2(a = 2 kg/mm2), and without tensile stress J0.", "texts": [ " Explanation of the stress effects and symbols The stress effects and symbols used in the present paper are as follows: (1)Jo (or Jo) indicates the magnetization on an initial magnetization curve after the sample is demagnetized under an applied tensile stress of o (or o = 0) and then put in a field H without removing o (or o = 0), and (2) RJo (or RJo) is the reversible magnetization change observed under the following conditions. The sample is demagnetized under an applied tensile stress of o (or o = 0), and then a field H is applied while o is still applied (or o = 0). When the applied stress is cycled between 0 and o, the magnetization is reversed. Fig. 1 shows the magnetization curves o f J 2 (o = 2 kg/mm 2) and Jo for the specimen, 5 mm in width, 330 mm in length and 0.5 mm in thickness. The application of tensile stress decreases the magnetization. Jo-Jo is plotted in fig. 2 as a function of tension o at H = 1.0e. Ja-Jo decreases non-linearly with tension up to about 1.5 kg/mm 2 and then decreases linearly with tension. The magnetostriction constant for the specimen is about 7.8 X10 -6. It has long been believed that the magnetization must be increased by the application of tension (except at I = 0 or saturation magnetization) in the materials which have positive magnetostriction" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001079_iros.2004.1389572-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001079_iros.2004.1389572-Figure4-1.png", "caption": "Fig. 4. Hypothesize the sequence ofevents during the delivery.", "texts": [ " 3 we explain the sequence of events that happen during the dispense process of nanwscale volumes of viscous material. In Fig. 3.A the syringe dispenser together with needle and residue from the previous trial is at the desired height above the base of the well. In Fig. 3.8 the desired volume of the cubic phase is completely expelled out.. In Fig. 3.C, the cubic phase dispenser is back up and there is a small residue at the needle tip. This residue is due to the formation of break point as explained below in Fig. 4. In Fig. 3.D one can notice the increase in the volume of residue at the needle tip. When no force is being exened, the relaxation of the compressed viscous material results in a slow flow of the cubic phase, which accounts for the increase in the volume of residue at the needle tip. Fig. 3.E.shows the needle with complete residue before moving to the next well. The residual volume at the needle tip is due to two different phenomena: Formation of breakpoint Relaxation of compressed biomaterial A. Break Point Residue We explain the formation of break point hom Fig. 4 that hypothesizes the set of events during the delivery of viscous biomatenals. In Fig. 4.A the dispensing tool with its residue from the previous mal is at desired height above the base of the well. In Fig. 4 . 8 the desired volume of the cubic phase is completely expelled out. In Fig, 4.C, the cubic phase dispenser is moving back up and thus the shape of the cubic phase started changing. The way in which the shape changes depends on the viscosity of the biomaterial. In Fig. 4.D one can notice the formation of break point that govems the amount of the cubic phase being delivered to the well. In Fig. 4.E the needle moves back up with the residual cubic phase. B. Relaxation of Compressed Biomaterial During dispensing, syringe plunger moves with speed to expel its contents at 1 pus. Due to high viscosity of the dispensing material there is a strong resistance to the flow, As a result force builds up in the needle. Therefore material gets compressed in the syringe barrel. Atkr plunger stop moving, relaxation of the compressed material results in a slow flow of the material out of the needle. Time of relaxation and amount of material that will be expelled during relaxation depends on the viscosity and wmpressibility of the material and on the diameter and length of the dispensing needle" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001555_iecon.2005.1569184-Figure9-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001555_iecon.2005.1569184-Figure9-1.png", "caption": "Fig. 9. The construction of sensor system.", "texts": [ " Remark It has been stated before that moment criteria and ZMP criteria are essentially the same. On the contrast, Lagrange criteria are much simple, with only two variables, and are much easier to calculate. Furthermore, in the former 2 criteria, information of accelerations are used while in the Lagrange criteria, information of torque is used. Acceleration is the result of torque, and as a result Lagrange criteria are much \u201dsensitive\u201d and changes quicker. Most importantly, element \u201dTorque\u201d is included and wheelie can be prevented by regulating it directly. IV. Experiment As shown in Fig. 9, in this system there are 5 sensors for: rotating speed (\u03c8) of the wheelchair frame, vertical and horizontal accelerations (Anx, Any), angle of both wheels (\u03b8l, \u03b8r). Experiments on horizontal plane have been carried out and the limit of torque is calculated. During the experiment, if human and motor torque is bigger than this limit, motor output decreases so that the total torque is equal to the limit. Data are shown in Fig. 10-12. It is obvious that \u03d5 vibrates around \u03d50, but it never exceed 0. Similarly, Xzmp does not cross over 0, it can be said that there is no wheelie during the experiment" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001476_s1064230706030099-Figure23-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001476_s1064230706030099-Figure23-1.png", "caption": "Fig. 23.", "texts": [ " The controlling device constructs the realization \u03bd*(t), t \u2208 T, of the optimal feedback of problem w* t( ) 0.2 2t, t 0 3.2[;,[\u2208sin= w* t( ) 0, t 3.2 5,[ ]\u2208= 436 JOURNAL OF COMPUTER AND SYSTEMS SCIENCES INTERNATIONAL Vol. 45 No. 3 2006 GABASOV, KIRILLOVA (5.2), based on the known current states x*(t), u*(t), t \u2208 Th. The correspondent controlling action u*(t), t \u2208 T is presented in Fig. 22 (curve 2). The optimal value of the cost function of problem (5.2) was equal to 0.408337. The interval of saturation is the interval [2.8, 3]. In Fig. 23, the trajectories of systems (5.2) and (5.3) (curves 1 and 2) are shown. In the classical theory during the control over dynamic systems, functioning under perturbations acting all the time, three types of loops are used: (1) feedback (2) feedforward (3) combined (feedforward\u2013feedback) loop. Among them the most universal is the first one, forming the controlling actions according to the output signals of the system. However, there are situations [3] in which it is expedient to apply feedforward loops, based on the available measurements of perturbations (input signals of the system)" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002532_physreve.77.041302-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002532_physreve.77.041302-Figure7-1.png", "caption": "FIG. 7. The figures in the left column show the patterns of simple orbits for e=1. The figures in the right column show the corresponding magnitude of the eigenvalues of F vi vi . The solid curves indicate the range of for which the corresponding orbit is neutrally linearly stable. The dashed curves indicate the range of for which the corresponding orbit is neutrally unstable.", "texts": [ " IV we give a theoretical prediction about the locations of these jumps. =59\u00b0 and b =61\u00b0. The behavior in a and b is completely different even though the change in is very small. IV. PERIODIC ORBITS FOR e=1 Figures 2\u20136 showed clearly that the jumps are associated with the occurrence of repeated collision patterns. The closer e is to unity, the more pronounced the phenomenon is. This leads us to study funnels with e=1 and examine the periodic orbits in such systems. For a given angle , a number of periodic orbits are possible see the left panel of Fig. 7 for the simplest few orbits . Each periodic orbit has a different sequence of collisions with the funnel walls. For a given collision sequence with m collisions, one needs to determine if the associated orbit can exist and whether it is stable. Let xi and yi be the location of the ith collision in the sequence and ui and vi be the x and y components of the particle velocity immediately before the ith collision. Since xi and yi are constrained to be on the funnel boundaries, we can eliminate xi in favor of yi", " The maps for Frl and Frr can be similarly constructed, and the results are Frl: vi+1 = ui+1 vi+1 = ui cos 2 + vi sin 2 ui sin 2 \u2212 vi cos 2 \u2212 ti , where ti = sec ui sin 3 \u2212 vi cos 3 + sec2 ui sin 3 \u2212 vi cos 3 2 \u2212 2 ui 2 + vi 2 \u2212 2 \u2212 d tan and Frr: vi+1 = ui+1 vi+1 = ui cos 2 + vi sin 2 ui sin 2 \u2212 2 tan + vi 2 \u2212 cos 2 . Combining the effects of all m collisions we can therefore construct a two-dimensional map F defined by vi+m=F vi . Since the orbit is periodic, we have vi =F vi , where stars denote the periodic orbit. Here, we will show the calculation of the simplest periodic orbit see Fig. 7 a1 . The particle is released with zero velocity, so a simple calculation gives u1 =0, v1 = 2 1+ d 2 tan +x1 tan ; then, it jumps from left wall to the right, so v2 =Flr v1 . In order to return along the same trajectory, the velocity before the second collision must be perpendicular to the wall, so u2 =\u2212v2 tan . This allow us to compute the location of the first collision and velocities before this collision, x1 = x0 = 1 + 1 2d tan 3 \u2212 tan2 5 tan2 + 1 3 tan5 \u2212 18 tan3 \u2212 5 tan , y1 = \u2212 x1 tan , u1 = 0, v1 = \u2212 2 1 + 1 2 d tan + x1 tan , and also the location of the second collision and velocities before this collision are given by x2 = x1 + 2v1 2 cos sin 3 , y2 = x2 tan , u2 = \u2212 v1 sin 2 , v2 = 2v1 cos2 . For this orbit to exist, the locations of the collisions must be consistent with the funnel geometry which requires xi d /2\u2212a sin for i=1,2. For a 1 and d 1 this can be reduced to 6 + d 4 \u2212 a 8 3 \u2212 d 4 + 3a 8 . Other orbits shown in Fig. 7 can be constructed in the same way. Not all orbits are stable. Unstable periodic orbits cannot be observed since any small disturbance will eventually destroy the periodicity. Adding a small perturbation vi in vi yields vi+m= dF dvi vi=vi vi. Since the process is nondissipative, the map F must preserve area in the phase space and so the 2-by-2 matrix dF dvi vi=vi must have unit determinant. Therefore the characteristic polynomial of the matrix is given by 2\u2212tr dF dvi vi=vi +1=0. If tr dF dvi vi=vi 2, the orbit will be linearly neutrally stable", " This means that trajectories that start sufficiently close to the periodic orbit will neither approach nor diverge from the periodic orbit while maintaining the same collision sequence. If tr dF dvi vi=vi 2, then the orbit will be linearly unstable. In this case, trajectories that start close to the periodic orbit will diverge from the periodic orbit until the particle can no longer follow the given collision sequence. At this point the dynamics becomes complicated and the trajectories can become highly sensitive to initial conditions. Below we derive the stability limits for the simplest orbit shown in Fig. 7 a1 . Given the collision sequence for this orbit, we have v2=Flr v1 , v3=Frl v2 , and v4=Fll v3 . Using the chain rule, we have 041302-5 dF dv1 v1=v1 = dFll dv3 v3=v3 \u00b7 dFrl dv2 v2=v2 \u00b7 dFlr dv1 v1=v1 . 4 Using the maps given above, we obtain dFlr dv1 = u2 u1 u2 v1 v2 u1 v2 v1 = cos 2 \u2212 sin 2 \u2212 sin 2 \u2212 t1 u1 \u2212 cos 2 \u2212 t1 v1 , where t1 u1 = \u2212 sec sin 3 + sec2 sin 3 u1 sin 3 + v1 cos 3 \u2212 2u1 sec2 u1 sin 3 + v1 cos 3 2 \u2212 2 u1 2 + v1 2 \u2212 2 \u2212 d tan , t1 v1 = \u2212 sec cos 3 + sec2 cos 3 u1 sin 3 + v1 cos 3 \u2212 2v1 sec2 u1 sin 3 + v1 cos 3 2 \u2212 2 u1 2 + v1 2 \u2212 2 \u2212 d tan ", " Combining these results, 4 becomes a11 a12 a21 a22 , where a11 = 256 cos10 \u2212 896 cos8 + 1120 cos6 \u2212 600 cos4 + 129 cos2 \u2212 8, a12 = \u2212 tan 256 cos10 \u2212 768 cos8 + 736 cos6 \u2212 240 cos4 + 21 cos2 , a21 = \u2212 tan 256 cos10 \u2212 1024 cos8 + 1440 cos6 \u2212 848 cos4 + 197 cos2 \u2212 13 , a22 = \u2212 256 cos10 + 1152 cos8 \u2212 1888 cos6 + 1352 cos4 \u2212 393 cos2 + 34. The orbit will be linearly neutrally stable if tr dF dv1 v1=v1 2, which gives 256 cos8 \u2212 768 cos6 + 752 cos4 \u2212 264 cos2 + 26 2. We can therefore obtain the range of in which this orbit exists and is linearly neutrally stable, as 4 ,arccos 3\u2212 2 4 = 45\u00b0 ,50.97\u00b0 . Similar computations can be performed for other orbits. In the right panel of Fig. 7 we plot the magnitudes of the eigenvalues as a function of the funnel angle for the corresponding periodic orbits shown on the same row in the left panel. The ranges of for which each orbit is linearly neutrally stable are shown as solid curves while the unstable ranges of are shown as dotted curves. In particular, Fig. 7 b2 shows that the orbit shown in Fig. 7 b1 is unstable for all values of . The neutrally stable ranges are also marked in Fig. 2, and one can clearly see that these ranges correspond exactly with the ranges in which the inelastic particle stays in the funnel for an unexpectedly long time. Moreover, the sensitivity of the trajectories to initial conditions is strongly correlated with the stability of the associated periodic orbits. In order to understand the extent to which the behavior of elastic particles is dominated by periodic orbits, we have performed simulations with an elastic particle in which we record the x locations of the first 2000 collisions for a given 041302-6 input location. In Fig. 8 we plot the results as a function of input location for =61\u00b0. There are a number of periodic orbits, and the periodic orbit shown in Fig. 7 d2 can be seen for x0 approximately 0.33. The trajectories of particles that have input locations in the range 0.3, 0.37 stay close to this periodic orbit. Other orbits are also visible, for example, an orbit with eight collisions can be seen for x0 approximately 0.27 and trajectories of particles that have input locations in the range 0.24, 0.3 stay close to this orbit. However, there are other regions such as 0, 0.7 and 0.47, 0.53 in which no clear pattern exists. For =59\u00b0, the results are dramatically different", " In this case, there are no clear periodic orbits and the points denoting the collision locations fill the entire space in the figure. Therefore we do not present this figure. Now we provide an explanation of the jumps that occur in Figs. 2\u20134. For e 1 periodic orbits do not exist. Nevertheless, the particle trajectories can adopt the same collision sequence as the corresponding elastic particle e=1 until near the time when the particle exits the funnel. We will refer to these trajectories as quasiperiodic orbits. The jumps are the remnants of the neutrally stable periodic orbits shown in the left panel of Fig. 7. Periodic orbits exist for wide ranges of the funnel angle , but it is not the existence of periodic orbits that leads to the large jumps in the mean duration, impulse, and energy loss. Rather, it is the existence of neutrally stable periodic orbits that leads to the anomalous behavior. The labels on Fig. 2 allow one to easily identify each jump with the associated orbit in Fig. 7. The simpler an orbit is, the larger the jump associated with it is. This is because the locations of the collisions in simple orbits tend to be relatively far away from the exit. The quasiperiodic orbit therefore takes a relatively long time for the collision loca- tions to move down toward the exit. On the other hand, complicated orbits tend to have a collision whose location is relatively near to the exit. Therefore the particle will exit the funnel much earlier than simple trajectories. This is why the major jumps in Figs", " 11, we plot the average duration for a funnel with convex parabolic walls p=0.2 . Figures 10 and 11 clearly show that the anomalous behavior still exists for funnels with walls of parabolic shape. All large jumps that were present for the case of flat walls are still present in the case of both concave and convex parabolic walls. Furthermore, Fig. 11 shows an additional anomalous peak near =28\u00b0, which does not exist in the case of flat walls. This peak is related to the periodic orbit shown in Fig. 7 b1 , which was unstable for flat walls, but becomes stable for convex parabolic walls. To demonstrate that the anomalous phenomenon is not a singular behavior of perfectly circular particles, we consider particles with elliptic shape falling through a funnel with parabolic walls; the shapes of walls are given by y=\u2212tan x+ d 2 + p x+ d 2 x+ d 2 +H cot for \u2212 d 2 +H cot x \u2212 d 2 and y=tan x\u2212 d 2 + p x\u2212 d 2 x\u2212 d 2 +H cot for d 2 x d 2 +H cot. The opening of the funnel is located at \u2212 d 2 x d 2 and y=0" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001774_iciea.2006.257199-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001774_iciea.2006.257199-Figure1-1.png", "caption": "Figure 1: Stator phase A, rotor phase a and the mutual inductance MAa between phases.", "texts": [ " The models agree well with experimental data for the particular fault considered and are suitable for model-based fault detection, to name one perspective application. This paper is part of the work reported in more detail in [13]. Define three coordinate axes for the stator and three for the rotor. The angle between two consecutive stator or rotor axes is 2\u03c0 3 . To simplify notation, introduce the electrical angle \u03b3\u0304 = p\u03b3. The stator phases A,B,C are connected in a Y - configuration yielding \u0131A + \u0131B + \u0131C = 0. Similarly, for the rotor phases, i.e. \u0131a + \u0131b + \u0131c = 0, see [13] for details. Fig.1 illustrates the model parameters definitions with respect to stator phase A and rotor phase a. Note the interaction between stator phase A and rotor phase a through MAa. The situation is similar for the other phases. Assume now that a fault occurs in rotor phase a. Relevant examples of faults are: broken rotor bar, broken end ring and diminished air gap in front of rotor phase a. Then it is reasonable to introduce two groups of parameters according to Table 1. Thus, a rotor fault results in one or several of the following model parameter alternations; Ra = Rr; La = Lr; Mar = Mr; MSa = M " ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000161_1.1598988-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000161_1.1598988-Figure4-1.png", "caption": "Fig. 4 Scheme showing the fluid-induced forces Fr and Ft , and force moments. Mr and Mt \u201e\u00ab: eccentricity, V: shaft rotational speed, v: whirling speed, \u2018\u2018O\u2019\u2019: impeller center on the whirl orbit, \u201er,t\u2026: radial and tangential to the whirl orbit\u2026", "texts": [ " The variation of leakage flow with Seal A and B caused by the whirl motion is unsteady, however, that with Seal C is almost steady. A series of experiments was carried out to determine the effect of the seal geometry on the rotordynamic fluid forces. The configurations of test facility were kept the same except the wear-ring seal, which were only changed from Seal A to C. Instrumentation and Data Acquisition System. The impeller is supported by the main shaft through a rotating force balance with a four-axis force sensor, as shown in Fig. 4. The force sensor is composed of two couples of parallel plates and four strain gauges per plate to measure the four-axis forces ~two forces and two force moments!. The output signals of the strain gauges are transferred to a data acquisition system through a slip ring, in which the four signal vector $V% are converted to four component vector of two force and two moment $FM% using a four-by-four transfer matrix @A# ~i.e., $FM %5@A#$V%). A preliminary set of dynamic calibration tests were conducted to obtain the transfer Fig", " Output signals are ensemble-averaged over 64 whirl orbits based on a triggering signal that indicates the instant when both the directions of the eccentricity and the impeller rotation come to a prescribed orientation. The fluid force and force moment are measured twice, that is, in air and in water at the same rotation and whirling speed. The former measurement provides the inertia force of the impeller itself due to the whirling motion. The fluidinduced force and force moment can be obtained by subtracting the former from the latter. Figure 4 shows the coordinate system. The r-axis is set in the direction of eccentricity \u00ab, and the t-axis is directed 90 deg from the r-axis in the direction of the impeller rotation. The fluid force F is applied to the center of the impeller \u2018\u2018O\u2019\u2019 on the whirl orbit. The fluid force F is represented in radial (Fr) and tangential (Ft) components to the whirl orbit, which are useful for the rotor vibration analysis. Measured fluid forces are normalized as ( f r , f t) 5(Fr ,Ft)/(M o\u00abV2), where M o5rpr2 2b2 is the mass of the fluid in the impeller. Uncertainty in the dimensionless fluid forces f r and f t is 60.3 ~dimensionless value!. The fluid force moment M applied on the impeller center \u2018\u2018O\u2019\u2019 is represented with radial (M r) and tangential (M t) components as shown in Fig. 4. Measured fluid force moments are normalized as (mr ,mt) 5(M r ,M t)/(M o r2\u00abV2). Uncertainty in the dimensionless fluid force moments mr and mt is 60.3 ~dimensionless value!. It should be noted here that the tangential fluid force, f t , is destabilizing for the whirl when f t3(v/V).0; i.e., f t and v/V are both positive or both negative. P0, and P1;P4 in Fig. 2 show the locations of pressure taps to measure the steady and unsteady pressure. P0 at the collector wall, and P1 at the diffuser inlet were used to measure the steady pressure performance for c, and cs with a manometer", " for the experimental support. They are also deeply grateful to Dr. Bruno Schiavello of Flowserve Pump Division for his helpful and useful comments. Also, express their gratitude for the effort of Mr. Kenta Yamamoto in support of this program as an undergraduate project at Osaka University. @A# 5 four-by-four transfer matrix b2 5 impeller axial width ~see Fig. 2! b3 5 diffuser axial width ~see Fig. 2! DCp 5 coefficient of unsteady pressure Dp ~zero-to-peak!, normalized by r(r2V)2 F 5 fluid force on impeller ~see Fig. 4! $FM% 5 four-component vector of two force and two moment Fr ,Ft 5 fluid force, radial ~r! and tangential ~t! to the whirl orbit ~see Fig. 4! f 5 dimensionless fluid force on impeller normalized by M o\u00abV25( f r 21 f t 2)1/2 f r , f t 5 dimensionless fluid force on impeller, radial ~r! and tangential ~t! to the whirl orbit, normalized by M o\u00abV2 794 \u00d5 Vol. 125, SEPTEMBER 2003 rom: http://fluidsengineering.asmedigitalcollection.asme.org/ on 01/28/20 f rp , f tp 5 dimensionless fluid force due to unsteady pressure around the impeller, radial (r), and tangential ~t! to the whirl orbit, normalized by M o\u00abV2 f RS 5 dimensionless fluid force caused by the rotating stall in vaneless diffuser, normalized by M o\u00abV2 f * 5 dimensionless frequency5frequency/(V/2p) Gap-A 5 radial clearance between impeller shroud edge and casing ~see Fig. 2! M 5 fluid force moment around the impeller center M r ,M t 5 fluid force moment, radial (r), and tangential ~t! component ~see Fig. 4! M o 5 reference value ~mass of fluid in impeller! 5rpr2 2b2 , mr ,mt 5 dimensionless fluid force moment on impeller, radial ~r! and tangential ~t! components, normalized by M or2\u00abV2 \u2018\u2018O\u2019\u2019 5 center of impeller on the whirl orbit ~see Fig. 4! p 5 pressure pt1 5 total pressure at impeller inlet Dp 5 unsteady pressure ~zero-to-peak! r 5 radius ~r,t! 5 radial and tangential axis ~see Fig. 4! r1 5 impeller inlet radius ~see Fig. 2! r2 5 impeller outlet radius ~see Fig. 2! r3 5 diffuser inlet radius ~see Fig. 2! r4 5 diffuser outlet radius ~see Fig. 2! S 5 seal radial clearance (S*5nominal radial clearance with \u00ab50) $V% 5 four-component force sensor vector \u00ab 5 radius of circular whirl orbit ~eccentricity! r 5 fluid density f 5 flow coefficient5flow rate/(2pr2 2b2V) fd 5 design flow coefficient c 5 pressure coefficient (at collector PO)5(p-pt1)/r(r2V)2 cs 5 static pressure coefficient (at diffuser inlet P1)5(p-pt1)/r(r2V)2 v 5 whirling angular velocity v8 5 angular velocity of rotating stall V 5 angular velocity of impeller v/V 5 whirl speed ratio v8/V 5 propagation speed ratio of rotating stall @1# Jery, B" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002163_j.mechmachtheory.2007.03.005-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002163_j.mechmachtheory.2007.03.005-Figure5-1.png", "caption": "Fig. 5. Geometric construction of the pitch of the LLC.", "texts": [ " The helicoidal vector field associated with homologous lines is similar to that in instantaneous kinematics, except for the difference in pitch. Furthermore, when the rotation and translation parameters of the finite displacement screw become infinitesimal, the pitch degenerates into that of an instantaneous screw. That is, p \u00bc lim d!0 /!0 d sin / \u00bc d / The pitch obtained in Eq. (14) has been used in the formation of the finite screw system for displacing a line [5] and in the calculation of the screw product of two screws [10,11]. What follows is an alternative way to determine pitch geometrically. As shown in Fig. 5, there are two pairs of homologous points: A1, A2, B1, and B2. Let A2 and B1 be coincident with point R. Connecting A1 and B1 gives the line L1, whose homologous line L2 can be obtained by connecting A2 and B2. The above geometrical construction of L1 and L2 was introduced in [3]. Next, we locate the midpoint of line segment A1B1, C1, and that of A2B2, C2. It can be seen that the vector from C1 to C2 is parallel to nb, the direction vector of the internal bisector of L1 and L2. The rise from C1 to C2 in the z direction is d, because C1 and C2 are homologous points" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003942_s13369-012-0287-1-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003942_s13369-012-0287-1-Figure3-1.png", "caption": "Fig. 3 ANSYS finite element mesh of the rotor", "texts": [ " It is a higher order 3-D element having a quadratic displacement behavior and is well suited to modeling irregular meshes (such as pump impellers). The element is defined by 10 nodes having three degrees of freedom at each node: translations in the nodal x , y, and z directions, [19]. Mesh refinement results show that a model refinement having number of elements greater than 86,446 does not have any significant improvement on the solution convergence to the values of the first five natural frequencies. In this case, the Block-Lanczos solver was used to extract the first five mode shapes. The ANSYS finite element mesh is shown in Fig. 3. The generalized eigenvalue problem was solved using the developed FEM formulation, and values of the first eight bending frequencies are obtained and listed in Table 2. Both backward and forward whirling frequencies are calculated at different rotating speeds. The results show the effect of rotor speed on the modal values. The natural frequencies of the pump rotor were also calculated using Ansys software, where the complicated impeller geometry was considered. The finite elements in such FEM structural analysis programs, in general, are not developed to account for the elemental reference rotation for non-axisymmetric components" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003612_s11044-010-9190-2-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003612_s11044-010-9190-2-Figure5-1.png", "caption": "Fig. 5 Dynamic eccentricity caused by magnetic eccentricity", "texts": [ "2 Magnetic eccentricity Here, dynamic eccentricity is caused by magnetic eccentricity of rotor mass \u00eam, which leads to non-uniform air gap, rotating with the angular rotary frequency \u03a9 . The non-uniform air gap is caused by a deviation of concentricity between the inner diameter of the rotor core and the outer diameter of the rotor core (Fig. 4). To compensate the unbalance, caused by this deviation of concentricity, a compensation unbalance is positioned, so that the centre of rotor mass U is not displaced from the rotation axis (Fig. 5). The non-uniform air gap causes a magnetic force in direction of the smallest air gap, which tries to increase the non-uniform air gap and forces the shaft centre point W to orbital movements [7]. 2.3 Bent rotor deflection Also, a bent rotor deflection \u00e2 causes a dynamic eccentricity. Bent rotor deflection is, e.g. caused by unequal thermal expansion of the rotor or due to a plastic deformation of the shaft. The rotor shaft is pre-bent. All three points W, U, and M are now displaced from the axis of rotation" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001555_iecon.2005.1569184-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001555_iecon.2005.1569184-Figure4-1.png", "caption": "Fig. 4. Standing a Broomstick Upright.", "texts": [ " Moment Criteria Assume clockwise moment positive, counterclockwise negative, moment criteria of the system can be written as m(ay + g)l sin \u03d5\u2212m(Ax \u2212 ax)l cos\u03d5 > 0 (3) Usually, when the total moment of system equals 0, system is stable. However, in this work, even when total moment is larger than 0, system is still stable because of two front wheels. B. ZMP (Zero Moment Point) Criteria The idea of ZMP was introduced by Mr. Vukobratovic at 1969 and 1972 [3]. ZMP (Zero Moment Point) is determined by the movement and gravity of the object. As shown in Fig. 4, the broomstick is influenced by gravitation and inertia force. These combined forces are called the total inertial force. Also, the point where the floor reaction force operates is called the floor reaction point. The intersection of the floor and the axis of the total inertial force have a total inertial force moment of 0, so it is called the The first two items are quite easy to calculate, while the other two are difficult. Assume the velocity of COM is vG and its horizontal component is vmx while vertical component is vmy" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000702_tro.2005.844677-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000702_tro.2005.844677-Figure3-1.png", "caption": "Fig. 3. Planar RRRR macro\u2013micro manipulator in its initial posture.", "texts": [ " Nevertheless, if, in a particular problem, taking only the twist does not produce reliable results, one can include the twist-rate in the outputs and use the linearized state\u2013output relations derived in [23]. With the linearized governing equations available, one can use the EKF relations to obtain the state estimates. These relations are derived in [30] as (40a) (40b) (40c) (40d) (40e) (40f) (40g) where and are the covariance matrices of the uncorrelated white-noise processes and , respectively. The ideas explained here regarding the estimation of the flexural states using an accelerometer array are demonstrated by simulating the dynamics of a planar manipulator, illustrated in Fig. 3, on a horizontal plane. The first two links of this manipulator\u2014constituting the macromanipulator\u2014are assumed flexible, while the other two\u2014making up the micromanipulator\u2014are assumed rigid. Moreover, we discretize the flexible links using the assumed-modes method, taking the clamped-free eigenfunctions as the shape-functions; the bending of each of the macromanipulator links is described by one flexural generalized coordinate. The specifications of the links are given in Table I. Furthermore, it is assumed that the noise covariances are constant and given by where and are the 2 2 and 4 4 identity matrices, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001604_robot.1987.1087799-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001604_robot.1987.1087799-Figure3-1.png", "caption": "Figure 3", "texts": [ " Throughout this section the word \u201cmanipulator\u201d refers to a planar 2-DOF linkage. It will also be assumed that only one obstacle is present. The consequences of having more than one obstacles will be discussed in Section 4. Let 0 be a set of obstacles in IR2, An obstacle in the workspace is defined to be a component of the set which is the intersection of the set 0 and the workspace W . It is necessary to differentiate between obstacles in IR2 and obstacles in the workspace because, as can be seen from Figure 3 , an obstacle in IR2 may become more than one obstacle in the workspace. With the above definition, obstacles in the workspace are always bounded sets because w is always bounded. From here on, the word \u201cobstacles\u201d is used to refer to obstacles in the workspace. The following definitions are also needed in the latter part of the paper. I624 Definition 1 Let X be a nonempty open set. Then X is connected if it is not the union of two disjoint nonempty open subsets of X. Otherwise, X is disconnected" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003361_978-90-481-9689-0_37-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003361_978-90-481-9689-0_37-Figure1-1.png", "caption": "Fig. 1 Basic concept of a WAMC.", "texts": [ " It is used together with actuators [1, 3] and it may contribute to a reduction in energy consumed by actuators and avoidance of loss of muscle strength and joint solidification. However, it has not been put into practical use since muscle fatigue occurs and it requires many actuators. We started to develop a walking assist machine using crutches (WAMC) in 2007. A wearable device with actuators generates motion in the lower limb to create a swing-through crutch gait while the user operates axillary crutches with both arms. The basic composition of the WAMC is presented in Figure 1. It satisfies conditions (1)\u2013(8) mentioned above. In order to make a practical WAMC, the mechanism, energy source, control system, and operating system should be determined taking into account several walking patterns in daily life, such as straight, ascending/descending steps, and turning motions, safety measures for external disturbances, etc. In our previous papers [2, 4], a basic composition for the WAMC was proposed and motion for flat floor was determined. In the present paper, motions for ascending and descending steps are discussed" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002618_detc2009-87343-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002618_detc2009-87343-Figure1-1.png", "caption": "Figure 1. Elastic-discrete model of a planetary gear and corresponding system coordinates. The distributed springs around the ring circumference are not shown.", "texts": [ " With the modal expressions of the elastic-discrete model from [3], the instability 1 Copyright c\u00a9 2009 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/22/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use boundaries are obtained as simple expressions. We show that many modes can not interact to create combination instabilities, and general instability existence rules are obtained for equally spaced planets. By adjusting the tooth numbers, contact ratios, and mesh phase one can minimize or completely suppress many potential instabilities. NOMENCLATURE MATHEMATICAL FORMULATION Figure 1 shows the elastic-discrete model of a planetary gear. The ring gear is modeled as a thin elastic body, and the sun, carrier and planets are treated as rigid bodies. The elastic-discrete model is established in detail in [3]. The same dimensionless parameters as in [3] are adopted here. The symbols with \u223c are dimensional variables, and the same symbols are used for dimensionless variables but without the \u223c. The degrees of freedom and dimensionless quantities are (also see the Appendix) v = v\u0303 R , q = q\u0303 R , q\u0303 = [x\u0303r, y\u0303r, u\u0303r\ufe38 \ufe37\ufe37 \ufe38 pr , x\u0303c, y\u0303c, u\u0303c\ufe38 \ufe37\ufe37 \ufe38 pc , x\u0303s, y\u0303s, u\u0303s\ufe38 \ufe37\ufe37 \ufe38 ps , \u03be\u03031, \u03b7\u03031, u\u03031\ufe38 \ufe37\ufe37 \ufe38 p1 , \u00b7 \u00b7 \u00b7 , \u03be\u0303N , \u03b7\u0303N , u\u0303N\ufe38 \ufe37\ufe37 \ufe38 pN ]T (1) \u03c4 = t T , T = \u221a m\u0303r k\u0303rn , ki = k\u0303i k\u0303rp , i = c,s, pn,rp,sp,bend, krbs = k\u0303rbsR k\u0303rp , krus = k\u0303rusR k\u0303rp , (2) kbend = EJ k\u0303rpR3(1\u2212\u03bd2) , m j = m\u0303 j m\u0303r , I j = I\u0303 j m\u0303rr2 j , j = r,c,s,n, Mi = M\u0303i m\u0303r , i = r,c,s,n, (3) K j = K\u0303 j k\u0303rp , j = rb,cb,sb, pp, Kn i = K\u0303n i k\u0303rp , i = r1,r2,c1,c2,s1,s2" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001747_4-431-31381-8_12-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001747_4-431-31381-8_12-Figure3-1.png", "caption": "Fig. 3. Analyical model of three degree of freedom walking mechanism", "texts": [ " In addition, the synchronized motion between the inverted pendulum motion of the supporting leg and the two-DOF pendulum motion of the swinging leg, as well as the balance of the input and the output energy, should have stable characteristics against small deviations from the synchronized motion. It is also assumed that a small viscous rotary damper with coefficient \u03b33 is applied to the knee joint of the swing leg, which produces a torque as: T3 = \u2212\u03b33(\u03b8\u03073 \u2212 \u03b8\u03072) (2) Under the assumption of a fixed bent knee angle of the supporting leg and a free knee joint of the swinging leg, the analytical model during the first phase is treated as a three-DOF link system, as shown in Fig.3. We get the equation of motion in the first phase as:\u23a1\u23a3M111 M112 M113 M122 M123 sym M133 \u23a4\u23a6\u23a1\u23a3 \u03b8\u03081 \u03b8\u03082 \u03b8\u03083 \u23a4\u23a6+ \u23a1\u23a3 0 C112 C113 \u2212C112 0 C123 \u2212C113 \u2212C123 0 \u23a4\u23a6\u23a1\u23a3 \u03b8\u03072 1 \u03b8\u03072 2 \u03b8\u03072 3 \u23a4\u23a6+ \u23a1\u23a3K11 K12 K13 \u23a4\u23a6 = \u23a1\u23a3 \u2212T2 T2 \u2212 T3 \u2212T3 \u23a4\u23a6 (3) where the elements M1ij ,C1ijand K1iof the matrices are shown in Appendix 1. T2 is the feedback input torque given by Eq.(1) while T3 is the viscous resistance torque at the knee joint, which is given by Eq.(2). When the angle between the shank and thigh of the swing leg becomes a certain value, the brake is activated and locks the knee joint" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001080_oceans.2005.1639997-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001080_oceans.2005.1639997-Figure6-1.png", "caption": "Fig. 6. Components of the MMA.", "texts": [ " The MMA was designed to mimic the size and shape of a YSI CTD sensor, which is a standard payload for the VTMAUV. Attached below the vehicle, the MMA provides passive roll stability by lowering the vehicle\u2019s CG. The VTMAUV supplies the MMA with 5 volts DC power. The actuator is controlled serially using simple position commands. The command set is the set of integers from 0 through 60, with 0 corresponding to the rearward-most mass position and 60 corresponding to a forward-most position. The layout of MMA components is shown in Fig. 6. The mass moves along an Acme 1 4 -inch lead screw with a pitch of 1 3 -inch per turn. The lead screw is driven through a customdesigned gear assembly by a high torque servomotor located at the rear of the actuator. The integrated servo potentiometer was replaced by an equivalently rated 10-turn potentiometer which is slaved to the lead screw at the front of the actuator through another gear assembly. Two brass lead screw nuts travel along the lead screw; a pair of anodized aluminum rails prevent the lead screw nuts from rotating, ensuring a smooth linear motion" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002884_vecims.2008.4592777-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002884_vecims.2008.4592777-Figure5-1.png", "caption": "Fig. 5. Needle and tissue simulation based on numerical calculation.", "texts": [ " Replacing the theoretic critical force Pe with the numerical solution, ' eP , the spring constant is derived as: 2 ' 4 2 44 eP EI l l \u03c0 \u03c0\u03bc \u2248 \u2212 (3) For a needle with a length of 50 mm, radius of 0.3 mm, and Young\u2019s modulus of 2\u00d71011 pa, the curve of tissue Young\u2019s modulus and its corresponding spring stiffness is plotted as shown in Fig. 4(b) based on Equation (3). From this curve, the critical force can be calculated and used in the haptic rendering for various tissue\u2019s Young\u2019s modulus and needle insertion depths. In the numerical solution, the homogeneous tissue is simply modeled as a block with a volume of 10\u00d710\u00d750 (mm3), and the needle has a radius of 0.31 mm as shown in Fig. 5 (a) and (b). The tissue and the needle are assembled by constraining the needle in a hole in the block volume. The contact condition between the tissue and needle is set to a bonded type. Two faces of the tissue and one end of the needle are fixed as boundary conditions. Fig. 5(c) shows the deformations of needle and tissue after an axial load is applied. It can be seen from the figure that the deformations mainly take place near the needle free end. Fig. 6(a) shows the model of needle buckling in twolayered tissues. From numerical calculation the needle buckling critical forces can be obtained in terms of different combinations of tissue elastic factors and their length proportions with regards to the overall length l of the needle. A sample set of data is given in Table 1", " From the table, it can be seen that the tissue elastic factor near the free end of the needle plays a more important role in determination of the critical force. From rows 1, 3, and 5 of the table, as we can see that the length proportions change a lot, however, the critical forces have very little change. It might be explained as that the main deformations of tissue and needle occur near the free end of the needle. The deformations of tissue and needle in the other end are relatively small as shown in Fig. 5(c). Using similar method three-layered tissues are modeled and analyzed as shown in Fig. 6(b) and Table 2. The difference between the data in row 1, 3, and 5 of Table 1 and that of row 1, and 3 in Table 2 is very small, though there is an extra layer tissue with large elastics in this model. where a is the maximum deflection, x varies from 0 to s, s is the span of the beam and \u03b1 is the slope angle at ends. \u03b1 can be numerically calculated according to [2]: 2 0 2 2 2 1 sin sin 2 e dw F P w \u03c0 \u03c0 \u03b1 = \u2212 \u222b (5) where Pe is the Euler buckling load (Equation 1), and F is the loading force applied to needle which will be explained in the next section" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001300_013-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001300_013-Figure5-1.png", "caption": "Fig. 5. Possible probe\u2013sample interactions as a probe approaches a surface with uniform properties: (a) before contact with any structure on the surface, (b) contact with the sharp surface protrusion, (c) further contact with the surface gap, and finally (d) deformation occurs over the entire contact surface. In the situation approaching (c), the probe makes locally weak contact with the surface, which is supposed to be closest to the ideal convoluted condition between the probe and the surface and where maximum roughness is also reached.", "texts": [ " The Rrms\u2013A0 relationship (results not shown) has a similar shape to the Ra\u2013A0 relationship. As the applied force A0 was increased and then reduced back to the initial value, a repeatable Ra\u2013A0 relationship was measured. As A0 increased, the measured roughness Ra first sharply increased to a maximum value Ra{p and then smoothly decreased. The variation in Ra in the adjusted range was around 0.05 nm. To achieve reproducible measurements with a high precision or where the roughness close to the atomic scale is reached,25) the force effect should be carefully taken into consideration. Figure 5 describes a possible process in which a probe approaches a surface with uniform surface properties from somewhere far from the surface in AM-AFM, under the condition that the probe\u2013sample interaction area in the gap region is larger than that in the protrusion region. When the probe is far from the surface [Fig. 5(a)], the resolution is low and the apparent height D and roughness are small. The continuous approach of the probe to the surface leads increases in resolution, apparent height D, and roughness, until the probe makes local contact with the sharper surface protrusion [Fig. 5(b)]. Further approach leads to contact with the surface gap [Fig. 5(c)]. Finally, deformation occurs over the entire contact surface [Fig. 5(d)] and both the apparent heightD and roughness decrease.21,26) The roughness rapidly increases when long-range forces dominate the imaging and slowly decreases after the probe\u2013sample contact. The roughness\u2013force relationship obtained in Fig. 4(b) might reflect this approaching process. Since the average force was attractive as determined from the phase signal, we suppose that the peak roughness Ra{p in the roughness\u2013force curve corresponds to a local weak contact close to the condition in Fig. 5(c) and is the experimental value closest to the ideal roughness as deduced from the geometrically widened image. 08KB11-3 # 2012 The Japan Society of Applied Physics On the basis of the selected imaging operating regime and the results of the analysis of the roughness\u2013force relationship, it can be concluded that a probe made of a weak adhesive material will come closer to the surface and lead to a more severe deformation when reaching Ra{p. That is, the probe made of a weak adhesive material will yield a small Ra{p if two probe radii are the same" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003412_978-3-642-37798-3_24-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003412_978-3-642-37798-3_24-Figure4-1.png", "caption": "Fig. 4. Left: Slider-Crank mec", "texts": [ " an ted t is n a od the the (3) 274 C. Esparza, R. N\u00fa\u00f1ez The X axis (North-South c shaking table are mechanic described above, as it is sh linear model for displacem with transfer functions as s Y-axis. This lineal approximation c each table axis, as it is show each axis in the range of + 6 5 for X and Y axis are pres respectively. The final tran the plant models used for th , and F. Gonz\u00e1lez els. Identification omponent) and the Y axis (East-West component) of al structures based on a slider-crank mechanism like it w own on figure 4. To simplify the model of each axi ent was obtained by the identification of a process mo hown in equation 4 for the X-axis and equation 5 for hanism used in the shaking table axis. Right: pulley mechanis .1 2. . . . . .1 1. . an be made by defining the lineal displacement region n in figure 4, which allows a maximum displacement cm to -6 cm. In table 1 the parameters of equations 4 ented. They achieve approximations of 91.8% and 91.9 sfer functions are shown in equations 6 and 7, which e Smith predictor in the branch without the delay. the as s, a del the m. (4) (5) of on and 7% are Model Reference Adaptive Position Controller with Smith Predictor 275 Equations 6 and 7 are also used to obtain the reference model transfer function in the adaptive controller, and the transfer functions used in the loops that modify the control parameters \u03b8" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003282_978-0-387-74244-1_11-Figure11.1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003282_978-0-387-74244-1_11-Figure11.1-1.png", "caption": "FIGURE 11.1. The DOF of a roll model of rigid vehicles are: x, y, \u03d5, \u03c8.", "texts": [ "1 F Vehicle Roll Dynamics In this chapter, we develop a dynamic model for a rigid vehicle having forward, lateral, yaw, and roll velocities. The model of a rollable rigid vehicle is more exact and more effective compared to the rigid vehicle planar model. Using this model, we are able to analyze the roll behavior of a vehicle as well as its maneuvering. 11.1 F Vehicle Coordinate and DOF Figure 11.1 illustrates a vehicle with a body coordinate B(Cxyz) at the mass center C. The x-axis is a longitudinal axis passing through C and directed forward. The y-axis goes laterally to the left from the driver\u2019s viewpoint. The z-axis makes the coordinate system a right-hand triad. When the car is parked on a flat horizontal road, the z-axis is perpendicular to the ground, opposite to the gravitational acceleration g. The equations of motion of the vehicle are expressed in B(Cxyz). Angular orientation and angular velocity of a vehicle are expressed by three angles: roll \u03d5, pitch \u03b8, yaw \u03c8, and their rates: roll rate p, pitch rate q, yaw rate r" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001455_36.1.77-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001455_36.1.77-Figure4-1.png", "caption": "FIGURE 4. Schematic reconstruction showing the alignment of the cuniculi (c, small arrows) and the formation of an additional tunnel (at, wide arrow) in Eopolydiexodina, by resorption and coalescence of several aligned cuniculi (according to Thompson, 1948, 1964, modified).", "texts": [ " The Fusuline Limestone sediments were deposited on a middle carbonate ramp (Vachard and Bouyx, 2002), comparable to biotopes of the modern fusiform equivalent Alveolinella (see Yordanova and Hohenegger, 2002). Approximately 650 specimens of Eopolydiexodina were examined. Only one microspheric individual (Pl. 1, Fig. 8) was identified with the following characters: larger size, more elongate, cigar-shaped test, and absence of additional tunnels. In fusulinids, the number of microspheric shells is usually very small (Hottinger, 1982), but this proportion, 1 : 650, is exceptionally low. The histogram of the megalospheric specimens (Fig. 4; n 5 430 specimens) can be interpreted as two peaks corresponding to two Gaussian curves. Among the megalospheric \u2018\u2018proloculi\u2019\u2019, nine show an internal, circular excentric globule (Pl. 3, Figs. 1\u20139); seven have the shape of a single globule in the \u2018\u2018proloculus\u2019\u2019 (Pl. 3, Figs. 1\u20137); one has the shape of two globules in a subrectangular or reniform (kidney-shaped) \u2018\u2018proloculus\u2019\u2019 (Pl. 3, Fig. 8); and one appears in the first chamber, upon the reniform proloculus (Pl. 3, Fig. 9). The early cement in both sides of the globules is an isopachous microsparite, its crystallinity differing from the other large crystals which occupy the rest of the \u2018\u2018proloculus\u2019\u2019 (Pl", " We introduce this term to designate the excentrilepidine-shaped apparatus of Eopolydiexodina. Type 2. Association of two proloculi, producing a kidneyshaped deuteroconch (\u2018\u2018reniform proloculus\u2019\u2019). Type 3. Association of two deuteroconchs (\u2018\u2018phrygische Mu\u0308tze\u2019\u2019 of Kahler, 1988, fig. 21 a\u2013b). No double or triple proloculus was observed in the studied specimens of Eopolydiexodina. Double or triple proloculi are present in other genera of fusulinids, for instance, those shown in the illustrations of Gubler (1935, pl. 3, fig. 4); White (1936, pl. 18, figs. 9\u201310, pl. 20, figs. 1\u2013 4); Wilde (1965, pl. 18, figs. 1\u20132, 4, 8); Nguyen (1979, pl. 31, fig. 8; 1980, pl. 1, fig. 12); and Kobayashi and Ujimaru (2000, pl. 2, figs. 34). Examples of polyvalence were shown by Wilde (1965, pl. 18, figs. 3, 5\u20136; pl. 19, fig. 5; pl. 20, figs. 1, 3\u20135, 7); Nguyen (1979, pl. 31, figs. 9\u201310;1980, pl. 1, Fig. 13); and Kahler (1988, figs. 22, 26). Bi-apertured proloculi were illustrated by Ozawa (1975, p. 129, pl. 26, fig. 10), and Nguyen (1979, pl", " If the \u2018\u2018proloculi\u2019\u2019 in Eopolydiexodina are embryonic apparati, the absence of the protoconch in many embryonic apparati must be discussed. Three possible explanations can be advanced: (1) The calcification of the proloculus wall is exceptional. (2) The embryonic apparati characterize a particular generation, which has few individuals, for instance, the A2 generation. (3) Both hypotheses are true. The excentrilepidinoid proloculi appear in the type material of several polydiexodinids: Eopolydiexodina douglasi (Lloyd, 1963, pl. 119, fig. 4, pl. 120, fig. 9), E. oregonensis (Bostwick and Nestell, 1965, pl. 74, fig. 1\u20132, 7) and Parafusulina deliciasensis Dunbar and Skinner (in Dunbar and others, 1936, pl. 1, fig. 7). This suggests that the excentrilepidinoid proloculus, independent of the specific subdivisions, actually corresponds to individuals of a special generation. Since megalospheric and microspheric specimens are known in many species of Eopolydiexodina (see compilation by Vachard and Bouyx, 2002, and Fig. 3 herein) and in the closely related genera Skinnerina, Polydiexodina and Parafusulina, we propose that there is a relationship between the types of embryonic apparatus and the life cycle of Eopolydiexodina darvasica" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000137_j.jchromb.2003.08.022-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000137_j.jchromb.2003.08.022-Figure2-1.png", "caption": "Fig. 2. Illustration of the microdialysis set-up for sampling from macrophages in cell culture.", "texts": [ " These tubing lines were placed into the incubator through a side port that has specially drilled holes to allow for insertion of this tubing to and from the inside of the incubator. The microdialysis probe (CMA/12 4 mm probe) was prepared by immersing it in a 70% ethanol solution coupled with perfusion through the probe with autoclaved distilled water for 30 min. Then the probe was placed in and perfused with autoclaved saline for 30 min to rinse out the ethanol. Thereafter, it was inserted through the hole drilled on the well plate cover to cell medium in the tissue culture plate as shown in Fig. 2. To confirm the formation of the azide product after the precolumn derivatization of nitrite with 2,4-DNPH, LC/MS and LC/MS/MS experiments were performed. The APCI mass spectra did not show molecular ion peaks for the azide (m/z 209). However, a fragment corresponding to a loss of 28 (N2) was detected at m/z 181. The inability of azides to form molecular cations has been reported and is consistent with their general susceptibility to release N2 [27]. MS/MS experiments were performed on mass-selected ions at m/z 181" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002532_physreve.77.041302-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002532_physreve.77.041302-Figure1-1.png", "caption": "FIG. 1. Sketch of a system in which a particle falls through a funnel with an angle .", "texts": [ " Wylie and Zhang 15 have shown that two driven inelastic particles can experience a bifurcation in which large numbers of complicated periodic orbits collapse onto a single simple orbit. II. FORMULATION In this paper, we consider a frictionless, inelastic particle of radius a falling under gravity g through a symmetric funnel with walls aligned at an angle to the horizontal and a gap of size d at the bottom of the funnel. The particle is released with zero initial velocity with its center at a height H above the bottom of the funnel and at a horizontal location x0 measured from the central axis of the funnel see Fig. 1 . When the particle collides with the walls it experiences an inelastic collision with coefficient of restitution e. Here e is defined as the ratio of the velocity normal to the wall immediately after to that immediately before the collision. We will consider the case where particles are dropped into the funnel at a random horizontal location x0. For simplicity we consider the distribution in which all locations have uniform probability of being chosen. Other distributions yield similar behavior" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003071_iccas.2008.4694528-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003071_iccas.2008.4694528-Figure4-1.png", "caption": "Fig. 4 Pilot-Scaled Helicopter Control Model.", "texts": [ " The following steepest descent method is utilized in order to modify the CMAC weights tables: Wnew P,h (t) = W old P,h(t) \u2212 g(t) \u2202J \u2202KP 1 K Wnew I,h (t) = W old I,h (t) \u2212 g(t) \u2202J \u2202KI 1 K (5) Wnew D,h (t) = W old D,h(t) \u2212 g(t) \u2202J \u2202KD 1 K where g(t) and J are the gradient to update the weights and the error criterion, and given by: g(t) = 1 c + a \u00b7 exp(\u2212b|u\u2217(t) \u2212 u(t)|) (6) J := 1 2 \u03b5(t)2 (7) \u03b5 = u\u2217(t) \u2212 u(t) (8) and a, b and c are the appropriate positive constants. Moreover, each partial differential of Eq. (5) is developed as follows: \u2202J \u2202KP = \u2202J \u2202\u03b5(t) \u2202\u03b5(t) \u2202u(t) \u2202u(t) \u2202KP = \u03b5(t){y(t) \u2212 y(t \u2212 1)} \u2202J \u2202KI = \u2202J \u2202\u03b5(t) \u2202\u03b5(t) \u2202u(t) \u2202u(t) \u2202KI (9) = \u2212\u03b5(t)e(t) \u2202J \u2202KD = \u2202J \u2202\u03b5(t) \u2202\u03b5(t) \u2202u(t) \u2202u(t) \u2202KD = \u03b5(t){y(t) \u2212 2y(t \u2212 1) + y(t \u2212 2)} In order to illustrate the effectiveness of the proposed human-skill based PID controller using CMACs, some experiments were performed by using a pilot-scaled helicopter control model. The experimental system is shown in Fig.4. This system has 2 inputs and 2 outputs. Two motors are equipped for controlling the main rotor and the tail rotor. Moreover, two rotary encoders are equipped for the purpose of measuring the pitch angle and the yaw angle. In this equipment, the human operator manipulates the operating Joystick, and regulates the position of the helicopter. For simplicity in operating, the constant input is employed for the main rotor, and the only yaw angle is regulated by manipulating the tail rotor. Therefore the control objective is to regulate the yaw angle y to any desired values by manipulating the control value u" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001525_2006-01-0460-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001525_2006-01-0460-Figure8-1.png", "caption": "Figure 8 Simulation set up for below knee impact (left) and ankle impact (right). Constraints are removed from femur and adequate activation levels are defined in the Hill type muscle bar elements to maintain an upright standing posture.", "texts": [ " The forces in the MCL are seen to be very low. This is attributed to the ligament stiffnesses being used, which are currently as originally defined in THUMS. Effect of muscle activation in a free standing posture has been studied next. In these simulations, a significant difference from the Kajzer test is that the pins on the femur were not modeled. Even though the impact locations near the ankle and knee were the same, the loading did not correspond to shear and bending. They are hence referred to as below-knee and at-ankle impacts (Figure 8). There are no earlier results for free standing impact tests to compare with. To represent cadaver tests, simulations were carried out with muscle response deactivated. In the second step, the standing posture of a pedestrian with muscle activation needed to maintain stability in a gravity field is modeled using data reported by Kuo et al., (1993). Rupture of ligaments was not modeled as ligament rupture is not common in knee injuries during pedestrian accidents (Chidester and Isenberg, 2001). The response of the standing posture modeled with active muscle forces was compared with the passive model response to determine the role of muscle loading" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002175_robot.2007.363146-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002175_robot.2007.363146-Figure3-1.png", "caption": "Fig. 3. The first case of limb singularity", "texts": [ " Determination of the former five columns, )5:1(:,limb$ , is 1 5 4 5 5 4 1 2 2 1 2 3 3 2 3 1 1 3 4 5 5 4 1 2 2 1 2 3 3 2 3 1 1 3 2 2 2 2 4 5 5 4 2 1 3 2 3 2 2 1 (:, : ) ( )[( ) ( ) ( )] ( )[( ) ( ) ( ) ( )] ( )[( )( ) ( )( )] limb l m l m x y x y x y x y x y x y l m l m x y x y x y x y x y x y x y x y l m l m x x y y x x y y = \u2212 \u2212 + \u2212 + \u2212 = \u2212 \u2212 + \u2212 + \u2212 + \u2212 = \u2212 \u2212 \u2212 \u2212 \u2212 \u2212 $ (10) Assume ],,[ zyx NNN 4545455445 =\u00d7= SSN (11) Then Eq.(10) can be rewritten as 1 5 45 2 1 3 2 3 2 2 1 (:, : ) [( )( ) ( )( )] limb zN x x y y x x y y= \u2212 \u2212 \u2212 \u2212 \u2212 $ (12) The sufficient and necessary condition for the limb singularity of RRR(RR) is Eq.(12) equals zero. According to Eq.(12), there are two special singular cases for a RRR(RR) limb listed in Table I. where ijP denote the plane determined by axes of Ri and Rj, 1) First case: As shown in Fig.3, axes of R1, R2, R3 are coplanar, where P123 is the plane determined by the axes of R1, R2, R3; P45 is the plane determined by axes of R4, R5. Assume O-YZ plane of the reference frame be the same with plane P123, Z-axis be along the axis of R3 for convenience. Thus, the kinematical screw system for a limb is $1=[0, 0, 1; y1, 0, 0]T $2=[0, 0, 1; y2, 0, 0]T $3=[0, 0, 1; 0, 0, 0]T $4=[l4, m4 , n4; (x0, y0, z0)\u00d7(l4, m4 , n4)]T $5=[l5, m5 , n5; (x0, y0, z0)\u00d7(l5, m5 , n5)]T (13) where [x0, y0, z0] is the coordinate of the rotation center" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001447_ijvd.2005.006606-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001447_ijvd.2005.006606-Figure1-1.png", "caption": "Figure 1 Vehicle model with sprung and unsprung bodies", "texts": [ " Finally, this study will present how to apply the full control concept which allows the vehicle to follow a determined motion scenario in vehicle cruise control. In the model, the steering angle, tractive or braking force, and active suspension force for each wheel are set as the individually controllable input parameters. There are 12 control inputs for the four-wheeled vehicle. Later on, after the steering angles of the same axle and the tractive forces on the same side are combined, a total of eight useful control inputs are used in the governing equations. The vehicle is divided into two parts: sprung and unsprung bodies, as shown in Figure 1(a). The unsprung (i.e. the wheels) is having only three d.o.f., where the body roll, pitch or bounce motion are not allowed. The sprung, the actual body of vehicle, has the complete six d.o.f. motion. It is assumed that the vehicle is symmetric about the xz plane and the sprung mass will revolve about the roll axis and the pitch axis of the vehicle. The roll axis A-A, as shown in Figure 1(b), is the line joining the front and rear roll centres of the vehicle. The pitch axis B-B, in Figure 1(b), is the line joining the right and left pitch centres. The longitudinal, lateral, and yaw motion of the sprung and unsprung are set to be identical. Following the procedure of deduction proposed by Yu and Liu (2001), the governing equations of the four-wheeled vehicle can be formulated in terms of longitudinal velocity U, lateral velocity V, bounce q, roll , pitch ', and yaw velocity , as shown in the following equation: M _U\u00ff V \u00ff \u00ffMs hs ' _ \u00ff \u00ff _' _q\u00ff c _' KxU 2 M _V U \u00ff Ms hs \u00ff _' \u00ff \u00ff _ _q\u00ff c _' Y FV I _ Iz _ _'\u00ff Ixz4 \u00ff _' \u00ff \u00ffMsc _ _q N M I Ix _' \u00ff Ixz1 _ _ _' \u00ff Msh _V U \u00ff _ _q \u00ff M I' ' Iy _ \u00ff Ixz2 2 Ixz3 _ 2 \u00ffMsh' _U\u00ff V _' _q\u00ff c _' \u00ffMs f c q\u00ff _' U\u00ff hs _' V _ \u00ff M' Ms q\u00ff c ' \u00ff _' U\u00ff hs _' _ V c hs _ \u00ff Fq 266666666664 377777777775 E 6 1 1 0 b 2 0 0 0 2666666664 3777777775 Fx1 Fx3 1 0 \u00ff b 2 0 0 0 2666666664 3777777775 Fx2 Fx4 0 Cf lf Cf 0 0 0 266666664 377777775 1 2 0 Cr \u00ff lrCr 0 0 0 266666664 377777775 3 4 0 0 0 \u00ff b 2 \u00ff lf f \u00ff 1 2666666664 3777777775 Fz1 0 0 0 b 2 \u00ff lf f \u00ff 1 2666666664 3777777775 Fz2 0 0 0 \u00ff b 2 lr \u00ff f 1 2666666664 3777777775 Fz3 0 0 0 b 2 lr \u00ff f 1 2666666664 3777777775 Fz4 1 where [E] is a 6 1 matrix and I Ixxs Mshsh ; I' Iyys Mshsh' Msc f c I Izzs Izzu Msc 2 Mue 2; Ix Izzs \u00ff Iyys \u00ffMshsh Iy Ixxs \u00ff Izzs Mshsh' \u00ffMsc f c Iz Iyys \u00ff Ixxs Msc 2 Ixz1 Ixzs \u00ffMsh c; Ixz2 Ixzs \u00ffMsh'c; Ixz3 Ixzs\u00ffMshs f c ; Ixz4 Ixzs\u00ffMshsc FV 2 U V Cf Cr \u00ff lf Cf \u00ff lrCr \u00ff M 2 U V lfCf \u00ff lrCr \u00ff l2f Cf l2rCr h i M b2 4 C f C r \u00ff _ K f K r \u00ff M' C f lf f \u00ff 2 C r lr \u00ff f 2 h i _' K f lf f \u00ff 2 K r lr \u00ff f 2 h i ' C r lr \u00ff f \u00ff C f lf f \u00ff _q K r lr \u00ff f \u00ff K f lf f \u00ff q Fq C f C r \u00ff _q K f K r \u00ff q lr\u00fff C r _' K r' \u00ff \u00ff lf f \u00ff C f _' K f' \u00ff C f C 1 C 2 \u00ff ; C r C 3 C 4 \u00ff ; K f K 1 K 2 \u00ff ; K r K 3 K 4 \u00ff Y \u00ffCf f1 f2 \u00ff Cr f3 f4 N \u00fflf Cf f1 f2 lrCr f3 f4 Cf C1 C2; Cr C3 C4 From Equation 1, the tractive forces on the same side, or the steering angles of the wheels of the same axle will have the same effects on the motion" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003543_978-1-84996-080-9_11-Figure11.10-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003543_978-1-84996-080-9_11-Figure11.10-1.png", "caption": "Fig. 11.10 Schematic of Blatchford\u2019s Intelligent Prosthesis (IP+)6", "texts": [ " A commercial example of an adaptive system is Blatchford\u2019s Intelligent Prosthesis (IP+) [28] which is stated to be capable of adapting to various modes of locomotion and also optimises the hip power available to the amputee. It is also stated that: The prosthesis provides stance control ranging from minimal resistance to yielding lock, capable of detecting level walking, ramp descent, stair descent, standing and instances of stumble. The stance resistance is set to preprogrammed levels for each mode which matches the user\u2019s level of control [28]. Figure 11.10 shows some details of the passive control of the adaptive prosthesis. The prosthesis has been tested on many users and the feedback has been used to refine the system. An example of a user-adaptive system is a magnetorheological knee prosthesis developed at MIT7 [33]. System design and development requirements include: \u2022 understanding of normal human locomotion; \u2022 components required for the intelligent prosthesis; \u2022 algorithms needed for the control of the prosthesis; \u2022 development, testing and verification" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003118_ettandgrs.2008.310-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003118_ettandgrs.2008.310-Figure1-1.png", "caption": "Fig. 1 3D model of the underwater manipulator", "texts": [ " It can be used in real-time control of underwater manipulator, but it has deficiency. In order to improve the tracking ability, a novel hybrid control method based on the cerebellar model articulation controller (CMAC) is proposed. The fuzzy theory is blend into the traditional CMAC. Simulation experiments results show that the hybrid method is efficiency and has strong anti-interference. The 7-function underwater manipulator is designed, and it can be used primarily in underwater robot application. The 3D model of the manipulator is shown in Fig.1, and the structure is shown in Fig.2. The underwater manipulator has six DOF and a paw, including shoulder yaw, shoulder pitch, elbow pitch, forearm roll, wrist pitch and wrist roll. The operating depth is 3000 meters, and hydraulic drive is adopted due to it is easy to seal in the underwater environment. The forearm roll joint is driven by swinging cylinder, while the wrist roll joint is driven by hydraulic motor, and the other joints are driven by hydraulic cylinder. 978-0-7695-3563-0/08 $25.00 \u00a9 2008 IEEE DOI 10" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000498_j.mechmachtheory.2005.09.001-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000498_j.mechmachtheory.2005.09.001-Figure2-1.png", "caption": "Fig. 2. Global variables of a four-bar mechanism.", "texts": [ " / Mechanism and Machine Theory 41 (2006) 505\u2013524 where [Me], [Ce] and [Ke] are mass, equivalent damping and equivalent stiffness matrices of a element, respectively; {Fe} is a load vector of an element and /e is a vector of element variables. Transforming the governing equations of the elements such that they are expressed in terms of global variables, and then assembling the element equations leads to the global governing equations of a four-bar mechanism. Many previous researchers defined global variables in a manner that kept a constant orientation with fixed axes. Fig. 2 [10\u201312] illustrates the set of global variables used in this paper when considering a flexible four-bar mechanism with each link modeled with one element of the type derived in Section 2.1. Based on this set of global variables, the base of the input link of the four-bar mechanism is assumed to be rigid. Also, curvatures at all pins are set to zero, except at the base of the input link where the mechanism is driven. Increasing the number of elements or using higher-degree shape functions, the number of global variables generated from the nodes of elements increases" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000795_tac.2003.821418-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000795_tac.2003.821418-Figure1-1.png", "caption": "Fig. 1. Schematic of a cellular network.", "texts": [ " In a sense, this class of algorithms is robust by allowing uncertainty in the system structure (see [30] for the issue of uncertainty allowed by feedback). This is a desirable property for many wireless network and Internet control problems, where it is not cost effective or simply infeasible to perform detailed system parameter estimates. The performance-target-tracking problem is modeled after the Quality-of-Service (QoS) tracking problem in wireless communication [8]. Consider a cellular network consisting of M mobile units distributed over L cells as depicted in Fig. 1. Each mobile unit intends to communicate with the base station controlling the cell to which it belongs. The communication channels are usually duplex, that is, bidirectional. For simplicity, we will focus on the uplink channel, that is, the communication channel from mobile units to the base stations. Due to propagation characteristics of the electromagnetic wave, the 0018-9286/04$20.00 \u00a9 2004 IEEE signal received for each communication channel is corrupted by interference noises coming from other mobile units" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003153_2009-26-0051-Figure7-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003153_2009-26-0051-Figure7-1.png", "caption": "Figure 7 : Contact Point Analysis of a Door Cover", "texts": [ " As we can see easily approximately half of the problems are squeak and half of the problems are rattle problems. SOLUTIONS IN PREVENTION FOR SQUEAK: Solutions are different for squeaks and rattles, since the root causes are different. There are certain parts in the car that are moving up to 3 mm and for air ducts to the IP or cross car beam 6 mm relative displacement were measured. As a rule of thumb, everything less than 3 mm is a contact and for special cases like air ducts even larger distances have to be considered as a contact. Fig. 7 shows a detailed contact point analysis of a door cover. For squeak the main root cause are material incompatibilities. Therefore solutions are in material selections, surface structures, antifriction coatings, additives and similar possibilities to eliminate squeak. For rattle we have to consider fixing systems, fixing strategies, tolerance chains, mechanical properties of materials, stiffness and resonance frequencies of parts. It is far beyond the scope of this paper to explain all the possible solutions" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001772_s11172-006-0499-1-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001772_s11172-006-0499-1-Figure2-1.png", "caption": "Fig. 2. Effect of pH on the activities of free invertase (1) and invertase immobilized on the PHEMA microbeads activated by ECH (2) and CC (3). 4 6 8 pH", "texts": [ " Parameters affecting the enzyme activity. We studied effects of the pH value, temperature, and substrate con centration on the enzymatic reaction rate, as well as the enzyme stability and repeated use capability. The activi ties of free and immobilized invertase were calculated by measuring the absorbance of the solution at 640 nm as described above. Effect of pH. The pH dependence of the activity of the free and immobilized enzymes was determined in the pH range of 3.0\u20148.0 at 30 \u00b0C. The results are shown in Fig. 2. The pH value for the maximum substrate conver sion was found to be 4.5 for free invertase and 5.0 and 5.5 for invertase immobilized on the PHEMA microbeads activated by ECH and CC, respectively. The optimum pH value of immobilized invertase was shifted by unity to the alkaline region. The change in the optimum pH value for the supported enzyme can be attributed to diffusional limitations and secondary interactions, because enzyme immobilization on the support affects its three dimen sional structure and arrangement of functional groups" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002748_ls.82-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002748_ls.82-Figure3-1.png", "caption": "Figure 3. Schematic diagram of high-pressure diamond anvil cell.", "texts": [ " The two TR Gel-Lubes used in this study were P-A-10 (Base oil, poly-\u03b1-olefi n, gelling agent A and Weight Percent of gelling agent) and P-B-10 (Base oil, poly-\u03b1-olefi n, gelling agent B and Weight Percent of gelling agent). In P-A-10, 10% mono-amido is added as gelling agent to the base fl uid PN-0; in P-B-10, 10% bis-amido is added as gelling agent to the base fl uid P-N-0. Properties of TR Gel-lube are shown in Table II. Copyright \u00a9 2009 John Wiley & Sons, Ltd. Lubrication Science 2009; 21: 183\u2013192 DOI: 10.1002/ls DAC was employed for generating high pressure, as shown in Figure 3. Figure 4 shows a schematic diagram of the optical system. The pressure is generated by compressing a cylindrical pressure chamber drilled in an austenitic stainless steel SUS303A gasket between opposed diamond anvils each weighing 0.2 carat. Mechanical properties of SUS303A gasket have Vickers hardness (HV) of 322 HV, a proof stress of 331 MPa, a tensile strength of 671 MPa, Young\u2019s modulus 171 GPa and elongation of 73.3%. The sample lubricant occupies the space within the pressure chamber with an inner diameter of 0", " The pressure is usually determined by the ruby scale method. However, pressure cannot be measured continuously in the ruby method. Here, the change of gasket thickness is used for pressure determination. The method of estimating the pressure in chamber is established by measuring the thickness of austenitic stainless steel gasket as follows. The phase diagram of n-dodecane CH3(CH2)10CH3,12 n-nonane CH3(CH2)7CH3 13 and water H2O14 has previously presented by the researchers. The apparatus shown in Figure 3 is also applicable to direct observation of crystallisation through the diamond under high pressure. Figure 5 shows the relation of solidifi cation pressure under each condition and the gasket thickness change \u2206h. Repeatability of the experiment is good. The thickness of gasket was measured per 2 \u00b5m for pressure determination. The picture of CCD (charge-coupled device) camera was recorded on DVD (digital versatile disc) recorder. Copyright \u00a9 2009 John Wiley & Sons, Ltd. Lubrication Science 2009; 21: 183\u2013192 DOI: 10" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000290_2004-01-1231-Figure8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000290_2004-01-1231-Figure8-1.png", "caption": "Figure 8. Simulation model for clunk analysis", "texts": [ " \u2022 The leaf spring is connected to chassis by a 6 dimensional GForce with nonlinear characteristic in driving direction (other components with linear spring characteristics) and to the rear leaf spring shackle by a bushing. The leaf spring characteristic is represented by a bushing between leaf spring part and rear axle. \u2022 The leaf spring shackles are connected to chassis and leaf spring by bushings. \u2022 Rear axle is supported on leaf springs by bushings and connected to chassis by two dampers. \u2022 Dynamometer representing vehicle inertia is supported on ground by a cylindrical joint. \u2022 The clunk model installed in ADAMS is described in Figure 8. During the start of the simulation the model is driven from both ends of the driveline. Engine torque (150Nm) is applied to the crankshaft and the overall acceleration of the system is controlled by a (smooth) step motion at the dynamometer. After 2.5 seconds the motion is deactivated. Now the whole driveline accelerates/ decelerates freely depending on the engine torque input. After 3 seconds of simulation (the model has reached a quite stable condition) the gear-shifting event creating the clunk is simulated" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000278_1-84628-179-2_7-Figure7.8-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000278_1-84628-179-2_7-Figure7.8-1.png", "caption": "Fig. 7.8. Body and inertial axis system.", "texts": [ " \u2022 A last reason to also avoid stalled flight conditions or tumble manoeuvres is because they worry potential users. With these considerations in mind some work has been done in performing numerical optimizations of the transition manoeuvres to try and achieve the above goals. The starting point for the work done on transition manoeuvre optimization as well as vehicle control design is a full nonlinear 6-DOF model of the T-Wing vehicle that has been developed by [160]. Basically this model consists of the normal nonlinear, rigid-body, 6-DOF equations of motion that apply to any aircraft (Figure 7.8); a simple mass model of the vehicle; and a large database of basic forces (or coefficients) and aerodynamic derivatives covering a large number of flight conditions. The next subsection shows the derivation of the nonlinear model of an airplane, after which the equations are particularized to the case of the TWing. The rigid body equations of motion are obtained from Newton\u2019s second law [21]: \u2211 F = d dt (mv) (7.1)\u2211 M = d dt H (7.2) \u2022 The summation of all external forces acting on a body is equal to the time rate of change of the momentum of the body" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001112_bfb0041186-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001112_bfb0041186-Figure1-1.png", "caption": "Figure 1. The system considered in the example.", "texts": [], "surrounding_texts": [ "S(rl) ffi (1~11 +e)-~ll. (5.9) (5.10) 6. AN ILLUSTRATIVE EXAMPLE Consider a system consisting of two rotors B i and 8 2 connected by a massless shaft B 3; see Figure I. Relative to inertial reference frame e, the system is constrained to rotate about a line L parallel to 61; It, Iz > 0 are the moments of inertia of B i. B2 respectively, about L. Rotor Bz is subject to a control moment u(t~t. Rotor B 1 is subject to an unknown disturbance torque \u00a2o(t)~s; the only information assumed available on e0 is an upper bound on I \u00a20(0 I, i.e., , ~ t ) t < 1 3 V t G s (6.]) where [3 is known. The system configtwation can be described by ~b and 0, the angular displacements of B 1 relative to e and B2 relative to B l, respectively; thus q =a [~ of. (6.2) We shall treat B3 as a neglected component; when it is rigid, 0 = O." ] }, { "image_filename": "designv11_61_0001459_00423110600872416-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001459_00423110600872416-Figure1-1.png", "caption": "Figure 1. Running gear with link suspension; (a) link suspension freight bogie; (b) link suspension single-axle running gear for two-axle freight wagons.", "texts": [ " This paper presents non-linear multibody simulations investigating these matters. As long as the characteristics cannot be controlled within closer limits than found in this study, there is a strong need for the sensitivity analysis to be made, both in predictive multibody simulations of vehicle dynamics as well as for verification and acceptance tests. Keywords: Freight wagon; Link suspension; Dynamics; Ride quality; Simulation; Test Freight wagons with link suspensions have existed for more than 100 years. Today, the link suspension, shown in figure 1, is the most common suspension type for freight wagons in Europe.Already in 1890, the principle of link suspension was defined as a standard on two-axle freight wagons, and the first bogie with link suspension was introduced in mid-1920s. The present designs originate from the early 1950s. In an ongoing research project at KTH, the dynamic behaviour of freight wagons is studied. Background to the project includes requirements to improve ride qualities and increase axle load, loading gauge and sometimes also the speed of freight trains in order to make rail freight traffic more competitive", " In Sweden, a whole network for freight traffic with an enlarged loading gauge and 25 ton axle load (for some lines 30 ton) is built. The largest rail freight company, Green Cargo AB, is running overnight mail service with specially equipped two-axle freight wagons running at 160 km/h with a maximum axle load of 20 ton. This paper emphasizes on the variation in link suspension characteristics and how these variations may influence the dynamics of a two-axle freight wagon. In two-axle wagons, the vehicle body is connected by single or double links to the parabolic or trapezoidal leaf spring that rests on the axle box (cf. figure 1(b)). This arrangement allows the axle box to move in both the longitudinal and the lateral directions relative to the carbody. The suspension is quite simple and robust and also occupies a modest amount of space laterally and vertically. Stiffness and damping are both provided by one system and is intended to be proportional to the vertical load Fbox on the axle box. This type of running gear is also quite light, thus allowing a maximum of payload in the wagon. Stationary test results of link suspension force\u2013displacement characteristics on freight wagons are known from a number of sources, for example, references [1\u20133], of which a sample is shown in figure 2", " A low amount of hysteresis means a low amount of damping of vehicle motions, when excited by the vehicle\u2013track interaction; see, for example, ref. [1]. In various on-track tests, it has been observed that two-axle wagons with new links and bearings can exhibit very poor ride qualities, with large lateral dynamic oscillations, under certain conditions; see, for example, ref. [5]. A parametric study of freight wagon lateral dynamics is performed by means of the multibody simulation software GENSYS [9]. The vehicle is a two-axle freight wagon with UIC double-link suspension (cf. figure 1(b)). Axle spacing is 9 m and the carbody length is 13 m. Total mass of the wagon is 45 ton, equally distributed on the two axles. The vehicle masses are modelled as rigid bodies. The non-linear wheel\u2013rail geometry is precalculated within GENSYS for a number of varying conditions; for further details, see section 5. The non-linear creep forces are interpolated from a four-dimensional table, generated by the FASTSIM algorithm of Kalker [10]. A simplified model of the link suspension is a linear spring in series with a friction element in parallel with another linear spring" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002973_robio.2009.5420758-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002973_robio.2009.5420758-Figure4-1.png", "caption": "Fig. 4 Actuator layoput", "texts": [ " In order to realize flexible neck motion like a human cervical spine, we adopted a coil spring for the head motion mechanism. Furthermore, the center of rotations for pitch rotation (\u201cPitch1\u201d) and yaw rotation were set in the base of the head. Movable positions and movable ranges were defined as shown in Fig. 3 by referring to anatomical knowledge. The forward and backward motions are realized by combination of the head rotation (\u201cPitch1\u201d) and the neck bending (\u201cPitch2\u201d). The roll-rotation, both pitch-rotations (\u201cPitch1\u201d and \u201cPitch2\u201d) and the yaw-rotation are also actuated by McKibben artificial muscle as shown in Fig. 4. We use electro-pneumatic regulators for controlling contraction of McKibben artificial muscle. Although various kinds of autonomous robots have been studied and developed a lot, intelligence of robots is lacked still to interact with human and act in daily lives automatically. Thus, it is difficult that robots conduct the classes autonomously. On the other hand, it is thought that tele-operated robots by hidden operator are practical rather than autonomous robots. Actually, tele-operated robots seem to behave automatically from the view point of human even though robots are moved by tele-operation" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002375_s11012-008-9143-5-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002375_s11012-008-9143-5-Figure3-1.png", "caption": "Fig. 3 Plot of stresses which arise in the disk rim", "texts": [ " The un-homogeneity of the heterophase material structure may be lessened by repeated cycles of deformational impact in turns with stabilization annealing. For the rotating disk of constant strength, whose straining occurs under two-axial state of stress, it follows from the equilibrium equation r d\u03c3r dr + r h dh dr \u03c3r + \u03c3r \u2212 \u03c3\u03d5 + \u03b3\u03c92r2 = 0, r \u2208 [0,RC] (1) that the condition \u03c3r(r) = \u03c3\u03d5(r) \u2261 \u03c3C = const is satisfied if the disk thickness h(r) changes according to the following: h = h0 exp [ \u2212\u03b3\u03c92r2 2\u03c3C ] , r \u2208 [0,RC], (2) where RC is the inner radius of the rim of disk (Fig. 3). The interaction between the central part of the disk and the rim occurs on the surface L with height hC (Fig. 3). The magnitude \u03c3C shall be determined from the equation for the rim equilibrium and the compatible deformation of the central part of the disk and its rim. The rim of the disk is considered as the ring with an external surface free of loading and on the surface L the radial stress \u03c3C acts (Fig. 3). As is accepted in the theory of shells and rings, it is taken that stress \u03c3\u03d5 is constant along the radius of the rim (the distribution of radial stress is not considered here). On the basis of the plot of the acting forces (Fig. 3), the equilibrium equation for the rim takes the form: \u222b RC+\u03b4 RC \u222b \u03c0/2 \u2212\u03c0/2 H\u03b3\u03c92r cos\u03b7r d\u03b7 dr \u2212 2\u03c3\u03d5H\u03b4 \u2212\u03c3C \u222b \u03c0/2 \u2212\u03c0/2 cos\u03b7 d\u03b7 = 0. After integration, neglecting \u03b43/3, we obtain H\u03b3\u03c92RC\u03b4(R + \u03b4) \u2212 \u03c3\u03d5H\u03b4 \u2212 \u03c3ChCRC = 0. (3) Because in the central part of the disk \u03c3r(r) = \u03c3\u03d5(r) = \u03c3C = const, then from the condition of continuity of radial displacements Ur = \u03b5\u03d5r , it follows that the stress \u03c3\u03d5 in the rim equals \u03c3C . From formula (3), provided that \u03c3\u03d5 = \u03c3C , it follows: \u03c3C = \u2212\u03b3\u03c92RC(RC + \u03b4) 1 + hCRC H\u03b4 . (4) If the stress \u03c3C1 and the profile of the disk h(r,\u03c91) are determined at a given frequency \u03c91 by formulae (4) and (2), then the stress \u03c3C2 , which satisfies the equilibrium equation of the disk with profile h(r,\u03c91) at another frequency \u03c92 = \u03c91 + \u03c9, shall be determined as follows" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000746_12.619143-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000746_12.619143-Figure5-1.png", "caption": "Fig. 5 Schematic configuration of a compliant continuous ring flange that axially preloads a lens in its cell (From Yoder6).", "texts": [ " The surface deformations, Ax, occur equally at each side of the lens if the surface radii are numerically the same. Otherwise, we calculate Ex individually for each surface. The deformations each act as \"springs\", with compliances of: CD=AX/P. (15) To consider this effect in our estimation ofK3, we add (CD)! and (CD)2 to the denominator ofEq. (4). 3.2 Effect of retainer deflection To include the effect of retainer flexure, we note a basic similarity between a threaded retainer and a continuous ring flange as used to provide preload in some lens mounts. See Fig. 5. The threaded joint is assumed to be rigid. We apply Eqs. (16) through (18) (adapted from Roark8) to determine the axial deflection of the retainer of axial length tR under a given preload P. Proc. of SPIE 587705-3 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 06/28/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx Ax = (KA_KB)(P I tR3 ) (16) KA 3(m2 \u2014 1)[a4 \u2014 b4 \u2014 4a2b2 ln(a Ib)] I (4itm2EMa2 ) (17) 3(m2 1)(m + 1)[2 ln(a Ib) + (b2/ a2) \u2014 1J[b4 + 2a2b2 ln(a I b) \u2014 a2b2] KB= (18) 4irm2 EM [b2 (m + 1) + a2 (m \u2014 1)] The compliance ofthis retainer is: CR=L\\X/PA", " The magnitudes of the Ax terms would then be smaller than predicted by Eqs. (20) through (22). This would modify the resultant K3 factor. Further study is needed in order to quantify this frictional influence. Two examples are given in this section to illustrate the estimation of K3 with all six of the above discussed temperaturerelated effects considered. The first example involves a large single element eqi-bi-convex 5K15 lens with diameter DG 6.5350 in. (165.9890 mm) clamped axially by a threaded retainer in a simple 606 1-T6 aluminum cell as shown in Fig.5. The design parameters of Table 2 apply. In the second example, the glass is changed to BK7 and the metal parts are changed to 6A14V titanium so the CTEs are more nearly equal than in Example 1. Proc. of SPIE 587705-4 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 06/28/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx The lens weight also changes from 4.7383 lb (2.1493 kg) to 3.2666 lb (1.4817 kg) because of the material change. All dimensions and temperatures remain as indicated in Table 2", "org/ on 06/28/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx AXIAL COMPLIANCE Because we do not always know K3 precisely during the design of optical assemblies or the value of K3 of a given design is too large, it is common practice for designers to provide a controlled amount of axial compliance in some component of the assembly so it will deflect slightly when the temperature changes, but not significantly affect the applied preload. This can be done by using a circular flange as illustrated in Fig. 5 or by shaping a threaded retainer so it acts as a ringshaped flexure. The deflection of the flange or retainer is then related directly to the preload as discussed in Section 3.2. When the temperature changes, the differential axial expansions of the metal and glass parts will modify the deflection, but, if carefully designed, the change in deflection will be a small fraction of that deflection and cause a proportionately small change in preload. This will, in most cases, allow the contact stress in the lenses to be reduced to quite tolerable values" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000572_1.1637648-Figure5-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000572_1.1637648-Figure5-1.png", "caption": "Fig. 5 The canonical system and the basis screws associated with a line displacement", "texts": [ " can be easily rewritten to represent a general screw associated with the finite line displacement. To do so, one only needs to further extract s12 from the parentheses to obtain the following result: S135lxSx1lxpSxp1ly Sy1lypSyp (25) where Sx5~1,0,0;0,0,0! (26) Sxp5~0,0,0;1,0,0! (27) Sy5S 0,sin u23 2 ,0;0,s23 cos u23 2 ,0D (28) Syp5S 0,0,0;0,sin u23 2 ,0D (29) and lx5sin u12 2 , lxp5 s12 2 cos u12 2 , (30) ly5cos u12 2 , lyp52 s12 2 sin u12 2 In Eq. ~25!, S13 is described as a screw of a 4-system whose basis screws can be Sx , Sy , Sxp and Syp ~Fig. 5!. Sx and Sy have finite quatches 0 and (s23/2)cot(u23/2), and are coaxial with the x- and JANUARY 2004, Vol. 126 \u00d5 59 shx?url=/data/journals/jmdedb/27774/ on 05/12/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F y-axes respectively. The other two screws, Sxp and Syp , with infinite quatches, are pure translation screws along the x- and y-directions respectively. The four coefficients shown in Eq. ~30! satisfy a constraint equation below: lxlxp1lylyp50 (31) Therefore, under Parkin\u2019s definition of quatch, the set of screws associated with a finite line displacement can be represented by a subset of the 4-system" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002223_j.1749-6632.1969.tb54291.x-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002223_j.1749-6632.1969.tb54291.x-Figure2-1.png", "caption": "FIGURE 2. Distribution of red cell acid phosphatase activities in the English population (dotted line) and the common phenotypes (continuous lines). Diagram based on data of Hopkinson el al. (1964).", "texts": [ " There are marked differences in the relative activities of the two isozymes present in the different homozygous types. In type A, they are roughly equal in activity. In type B, the front (i.e., more anodal) band is more active than the back, and in type C, the back band is much more active than the front one. There are also striking differences in total activity between the types. (Hopkinson et al., 1964). On the average, the acid phosphatase activity of type B red cells is about 50% greater than that of type A cells, and type C cells show an even greater total activity. (FIGURE 2). The type A isozymes differ from types B and C in their electrophoretic mobilities, but some curious effects may be produced by certain additions to the buffer systems (Harris et al., 1968). Thus, in a phosphate buffer system at around pH 6, they migrate more slowly toward the anode than the corresponding B or C isozymes. If, however, a tricarboxylic acid, such as citric acid, is added, the mobilities of both the A isozymes are increased relative to the corresponding B and C isozymes and migrate more rapidly toward the anode" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000297_s0076-6879(04)81045-0-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000297_s0076-6879(04)81045-0-Figure4-1.png", "caption": "Fig. 4. Diagram of the fully assembled oxygen sensor positioned into a stopcock.", "texts": [ " Gently pull the electrodes from the back of the sensor body to ensure that the working electrode is not pushed against the SR plug in the distal tip of the sensor sleeve. Insert a 90-mm steel support wire through the sensor body and into the epoxy plug used to join the sensor sleeve and sensor body. Wipe any excess epoxy from the outside of the sensor and allow the epoxy to completely cure. 7. Fix the sensor into a four-way stopcock to allow attachment to the catheter used to introduce the sensor into the vessel. Figure 4 shows positioning of the sensor within the stopcock. Prior to placing the sensor in the stopcock, the valve should be turned to the completely open position because the valve cannot be turned once the sensor transects the stopcock. The end that the electrode leads protrude from is then sealed with 5-min epoxy, and the end of the sensor body and the steel wire are also fixed into place with the epoxy. After the epoxy has dried completely, trim the steel support wire to the sensor body. Chemical Microsensor potentiostat (Diamond Electro-Tech, Ann Arbor, MI) DATAQ-700 USB data acquisition card (DATAQ Instruments, Akron, OH) 8" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0003040_icems.2009.5382795-Figure14-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0003040_icems.2009.5382795-Figure14-1.png", "caption": "Fig. 14 Basic experimental apparatus.", "texts": [ " In the case of the distributed winding motors, the magnet eddy-current loss is mainly produced by the carrier harmonics of the PWM inverter [3]. As a consequence, the magnet segmentation effect of the IPM often became smaller than that of the SPM. On the other hand, in the case of the concentrated winding motor, the slot harmonic losses are considerably larger than the carrier harmonic losses, as shown in Fig. 6. Thus, this effect can be considered as negligible because of the lower time-harmonic orders. To support the above-mentioned results and considerations, the basic experiment by the magnet specimens is carried out. Fig. 14 shows the experimental apparatus. The magnets with a thermal sensor are placed at the center of an exciting coil. After surrounding the magnet by heat insulation material, the coil is excited by AC current whose frequency is 1 to 30 kHz. Then, the eddy-current loss is estimated by the following expression [7] with the increase in the temperature during 5 minutes. dt dTmcWmag = (11) where m is the mass, c is the specific heat, and T is the temperature. Fig. 15 shows the experimental and calculated results" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000016_s0167-8922(03)80067-6-Figure1-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000016_s0167-8922(03)80067-6-Figure1-1.png", "caption": "Figure 1. Experimental set-up", "texts": [ " The present study investigates the film behaviour of two different lubricants during sinusoidal variation of entrainment speed. The lubricants are chosen to have almost identical dynamic viscosity at atmospheric pressure under the test temperatures used but a quite large difference in pressure-viscosity coefficient. 2. EXPERIMENTAL SET-UP 2.1. The apparatus The experimental method for measuring film thickness used in this study was ultra-thin film optical interferometry. A schematic of the apparatus is shown in Figure 1. The EHD contact is formed between a sapphire disc, 100 mm diameter and 5 mm thickness and a 19 mm diameter steel ball. The disc is coated with a partially reflective chromium layer and a silica layer. Both contacting surfaces have roughness of about 10 nm Ra, much less than the smallest film thickness studied. The disc and the ball are driven at controlled speeds, which are logged continuously during each test. The microscope has a double role, to direct the light from a xenon arc lamp light source to the contact and also from the contact to the entrance slit of a spectrometer" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001255_iciis.2006.365733-Figure2-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001255_iciis.2006.365733-Figure2-1.png", "caption": "Fig. 2. Body frame attached to the ECAV", "texts": [ " Kinematic Model The configuration space of an ECAV is given by the special Euclidian group SE(3), where [p(t), R(t)] C SE(3) denotes the position and orientation of the ECAV with respect to the inertial frame at time t. The motion of an ECAV is then described by a smooth curve in SE(3). The rotation R is parameterized by the Euler angles \"yaw, pitch, roll\", ( b, , ) away from the singularity of a and the translation p is parameterized by p(t) = (X, y, z). Consider an orthonormal coordinate frame (T, B, N) fixed to the body of the ECAV at its mass center as shown in Fig.2. We call this the ECAV body frame. The Euler angles of the ECAV with respect to the inertial frame are shown in Fig.3. Two nonholonomic equations apply on the ECAV which arises from the fact that the ECAV cannot move in a direction normal to its heading direction: From a pure kinematic point of view, the equivalent control system of the ECAV can then be written as: \u00b1 = coS&cos 0 is the torsion of a(s) at s. We can then easily derive: ds N(s) = -T(s) B(s) k(s)B(s). These relations give the inherent properties of the 1-4244-0322- 7/06/$20. 00 C2006 IEEE B p=0 N p=0 (1) 256 parameterized curve, a(t), and are the well known SerretFrenet formulas, written compactly as follows: [T(s)1 [N(s) LB (s) J L 0 k(s) -k(s) 0 O T(S) 01 [T(s)1 -T(s) N(s) 0 J LB(s)] T, N, B makes up an orthonormal basis and is what was used as the body frame of the ECAV, shown in Fig.2. It is clear that the two intrinsic quantities of curvature (k) and torsion (T) can be used as the controls of the space curve a(t). The fundamental theorem of space curves [12], given below, assures us that the resulting a(t) is unique. Theorem 1: Given two single-valued continuousfunctions k(s) (curvature) and T(s) (torsion) for s > 0, there exists exactly one space curve a(s), up to a rigid body motion. We assume the speed V to be constant, leading to the following simple relation between the parameters of arc length and time: s = V t" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0000390_05698190500313478-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0000390_05698190500313478-Figure3-1.png", "caption": "Fig. 3\u2014Experimental apparatus.", "texts": [ " Test bearings A, B, and C have a clearance between the polymer lubricants and the raceways. Therefore, there is no direct contact between the polymer lubricants and the raceways in all the test bearings. Even if the polymer lubricants do not directly contact the balls and the raceways, the polymer lubricant supplies the lubricating oil to the contact points between the raceways and balls. In addition, the oil discharge rate from the polymer lubricant increases as the temperature rises (Takajo (13)). The experimental apparatus is shown in Fig. 3. Each test bearing was attached to the end of the shaft in the experimental apparatus and was axially loaded by a weight. While the inner ring was rotated at a certain speed, the running torque of the test bearing was detected by a load cell (NEC San-ei, T1-550), which was located 120 mm from the test bearing\u2019s center axis. The temperature of the polymer lubricants was detected by a radiation thermometer (Impac, IN3000). In addition, the rotational vibration of the outer ring in operation was dampened by oil dampers, and the contact vibration between the arm of the housing and the load cell was dampened by silicone gel" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002100_jmes_jour_1969_011_071_02-Figure6-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002100_jmes_jour_1969_011_071_02-Figure6-1.png", "caption": "Fig. 6. Overcut angle 6, and conjugation", "texts": [ " It makes J O U R N A L M E C H A N I C A L E N G I N E E R I N G SCIENCE a better insight possible, especially into the manufacturing process. Suppose an internal gear pair not satisfying rule (12: has a centre distance a, at which the overcut occurs, a, (rh2 -rb l ) . The overcut spans an arc 8, = - T/z , on the tip circle of the internal gear, where &,, &, and uU,, are defined by equations (3), (4) and ( 5 ) for a = a,. Let the centre distance a, increase to a centre distance uII, until the decrease in tooth thickness of the internal gear corresponds with the overcut arc B,, Fig. 6. z2-z1 (inv a,,,I-inv C C ~ ~ , ~ ) = 8, . (14) 2 2 where aWI1 is defined by equation (5) for a = air. Then a,, is called the conjugate centre distance to a,. z,(4,,+inv q,l-inv Elimination of 6, results in -z2(+,,+inv c:,,,-inv (r,\u201dII) = 01 (15) r b 2 - r b l cos a,,, = a, rb2- r b l a,, = - cos %,I 1 (19) , At the new centre distance a,, new overcut may be expected. The conjugation of centre distances seems similar to the story of Achilles a, pursuing the tortoise a,,, but in our case the tortoise a,, goes to a well-defined point of maximum overcut and then creeps back until he meets Achilles a, in the point described by existing theory" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002168_isie.2007.4374753-Figure4-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002168_isie.2007.4374753-Figure4-1.png", "caption": "Fig. 4. Section of a \u201cthree-phase asynchronous torque-motor\u201d", "texts": [ " In figure 2, that phenomenon is illustrated, showing that the deeper currents are weaker, and have also some delay in the direction of the displacement. Forming a rotor with spiral shape sheets, distributed in a radial disposition around the shaft, it is possible to generate magnetic fields stay more in the rotor's periphery, inducing peripheral e.m.f, and currents along the same sheets, that are only active in their periphery. The peripheral currents of this rotor have more section to circulate, compared with a normal cage rotor's current The figure 4 shows a developed plain representation of the disposition of the sheets. Instead of being shaped in an angular shape, in order to make a difference between the both zones, one where active currents go through, and the other which is used to receive the possible returning currents (A returning currents proposal). In spite of this, the returning currents can be established in two manners: 1) Option A: Through short-circuit rings. With that kind of construction the only rotor resistance that must be considered, is the one corresponding to the outside of the iron sheets that form the rotor" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002542_00029890.2009.11920919-Figure18-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002542_00029890.2009.11920919-Figure18-1.png", "caption": "Figure 18. Special cases of Example 2 with negative k, and \u03bb = 0.3 in (a)\u2013(c).", "texts": [ " The trammel and the trace have the same slope at their point of intersection when (19) is satisfied. If n = mk then dn/dm = kmk\u22121 = kn/m, and (19) implies n m ( \u03bck \u03bb + 1 ) = 0. (26) Relation (26) is satisfied if n/m = 0 (when the trammel is horizontal) but (26) also holds for a nonhorizontal trammel when the point of subdivision (\u03bbm, \u03bcn) on the trammel satisfies \u03bb/\u03bc = \u2212k, or \u03bb = k/(k \u2212 1), k = 1. For this choice of \u03bb, the single trace \u03c4(\u03bb) coincides with the envelope of the family of traces, and a line through the trammel is tangent to both in the fourth quadrant as shown in Figure 17d. Figure 18 shows special cases of Example 2 when the exponent k is negative. In Figure 18a, b, c we have k = \u22121, k = \u22122, k = \u22123, respectively. The exceptional case 130 c\u00a9 THE MATHEMATICAL ASSOCIATION OF AMERICA [Monthly 116 (26) is shown in Figure 18d. For the special choice of ratio \u03bb/\u03bc = \u2212k, the trace \u03c4(\u03bb) coincides with the envelope and the trammel is tangent to both in the first quadrant as indicated in Figure 18d. The governing curve, each trace, and the envelope in Figure 18a are rectangular hyperbolas with the axes as asymptotes. The next example treats kth power governing hyperbolas written in standard form. Example 3. G(x, y) = ( x A )k \u2212 ( y B )k \u2212 1 (kth power hyperbola, k = 0). By Theorem 5 the trace is also a generalized kth power hyperbola given by ( x \u03bbA )k \u2212 ( y \u03bcB )k = 1. If k = \u22121, the envelope has Cartesian equation resembling that in (25) with a difference instead of a sum: ( x A )k/(k+1) \u2212 ( y B )k/(k+1) = 1. 12. APPLICATION: GRAPHIC CONSTRUCTION OF ENVELOPES AND GOVERNORS AS A CLASSROOM ACTIVITY" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0002645_gt2009-60243-Figure12-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0002645_gt2009-60243-Figure12-1.png", "caption": "Figure 12: Unit B Shaft Seal Mating Seal Ring", "texts": [ " As shown in Figure 7, vibration sensors are located at five planes along the engine assembly. Velocity pickups on the generator are shown in Figure 8 and are labeled 3HV and 3VV (aft end) and 4HV and 4VV (forward end). Figure 9 shows the turbine output shaft end displacement sensor (eddy current proximity probe) used to generate vibration plots. nloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 10/04/2017 Ter Details that bring clarity on the cause and effect of the shaft heating hypothesis can be found in Figure 10 through Figure 12. Figure 10 shows an engine cross-section with the locations of potential close-clearance rub sites that could generate heat, and adjacent shaft elements, which could be affected by localized heating. The shaft seal at the lower left of the figure has four rotating steel teeth in close proximity to a bronze seal ring. The bearing-bronze alloy is listed as being wear-resistant, moderately machinable and suitable for gears and bearings. 5 Copyright \u00a9 2009 by ASME ms of Use: http://www.asme.org/about-asme/terms-of-use Down The rotating seal ring is adjacent to the aft bearing inner ring, which carries the bending load from the compressor to the shaft", " Localized, unsymmetrical heating at this location would cause unequal axial expansion of the inner ring and put a temporary, local \u201ckink\u201d in the shaft. Both compressor seals have steel teeth that run against an alloy designed to be abraded-away under contact. The inner seal is very close to the Hirth coupling, which joins compressor and turbine wheels. Localized, unsymmetrical heating of the compressor wheel in this area also would cause unequal axial expansion across the joint and increase the turbine runout, but to a lesser extent than heating at the shaft seal. The shaft seal teeth and seal ring are shown in Figure 11 and Figure 12. The seal labyrinth teeth have small deposits of bronze visible, and the deposits are localized to a relatively small angular extent. The seal mating ring has evidence of rubbing on the inside diameter, again with a restricted angular extent. The bronze material was not abraded, but was pushed into raised ridges alongside the rub grooves. This implies that a localized seal rub has occurred, with continuous heat generation as the material was not simply cut away. This heat would inevitably find its way into the shaft and result in non-uniform heating and thermal growth of the shaft contributing to a bowing of the rotor assembly" ], "surrounding_texts": [] }, { "image_filename": "designv11_61_0001255_iciis.2006.365733-Figure3-1.png", "original_path": "designv11-61/openalex_figure/designv11_61_0001255_iciis.2006.365733-Figure3-1.png", "caption": "Fig. 3. Euler angles of the ECAV", "texts": [ " Kinematic Model The configuration space of an ECAV is given by the special Euclidian group SE(3), where [p(t), R(t)] C SE(3) denotes the position and orientation of the ECAV with respect to the inertial frame at time t. The motion of an ECAV is then described by a smooth curve in SE(3). The rotation R is parameterized by the Euler angles \"yaw, pitch, roll\", ( b, , ) away from the singularity of a and the translation p is parameterized by p(t) = (X, y, z). Consider an orthonormal coordinate frame (T, B, N) fixed to the body of the ECAV at its mass center as shown in Fig.2. We call this the ECAV body frame. The Euler angles of the ECAV with respect to the inertial frame are shown in Fig.3. Two nonholonomic equations apply on the ECAV which arises from the fact that the ECAV cannot move in a direction normal to its heading direction: From a pure kinematic point of view, the equivalent control system of the ECAV can then be written as: \u00b1 = coS&cos