[ { "image_filename": "designv11_20_0002535_ccece.2007.144-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002535_ccece.2007.144-Figure1-1.png", "caption": "Figure 1 Equivalent circuit of Induction motors", "texts": [ " Calculations of the starting voltage, starting current and starting torque are derived mathematically for both starting methods. Case studies for the cross-line and VFD starting are performed and analyzed. II. STARTING PARAMETERS OF INDUCTION MOTORS Cross-line starting applies full voltage at rated frequency to the motor terminal to start an induction motor. If there are no electrical or mechanical constraints, it can be typically used in most cases. The traditional equivalent circuit of induction motors is shown in Figure 1. 0840-7789/07/$25.00 \u00a92007 IEEE Electrical parameters of induction motors during normal operation and motor starting can be calculated using the equivalent circuit. Generally, the motor impedance Zmotor, stator current I1, rotor current I2' and the electromagnetic torque Tem can be calculated using the following equations: ( ) ( ) ( )m m motor XXj S R jXjX S R jXRZ ++ + ++= 2 1 2 2 1 2 11 (1) ( ) ( ) ( )m m motor XXj S R jXjX S R jXR V Z V I ++ + ++ == 2 1 2 2 1 2 11 11 1 (2) ( ) 1 2 1 2 2 I XXj S R jX I m m ++ =\u2032 (3) p fn 11 1 4 60 2 \u03c0\u03c0 ==\u2126 (4) 11 2 2 2 1 1 2 1 3 S RIS p P T cu em em \u2126 \u2032 = \u2126 = \u2126 = (5) Where R1 and X1 \u2013 resistance and leakage reactance of the stator, R2 and X2 \u2013 resistance and leakage reactance of the rotor referred to the stator side, Xm \u2013 magnetizing reactance, Zmotor \u2013 motor equivalent impedance, V1 \u2013 voltage per phase at the motor terminal with the power system frequency of 50Hz/60Hz, I1 \u2013 stator current, I2' \u2013 rotor current referred to stator side, Tem \u2013 electromagnetic torque, f1 \u2013 starting frequency, S1 \u2013 slip, p \u2013 pole number of induction motors" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002515_00423110701810596-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002515_00423110701810596-Figure5-1.png", "caption": "Figure 5. Control volume particular extracted from the envelope and relative meridian section.", "texts": [ " Relation (2) gives an immediate interpretation of the T variation with \u03d5: varying \u03d5 the coordinate r changes, and also the area on which the pressure pi acts. Consequently, the force that must be balanced by the tension T changes. Similarly, the Tp expression can be explained by forces of equilibrium considerations regarding the previously described control volume. After a simple passage, the second equation of Equation (1) becomes Tp\u03c1( \u2212 \u03d5) = pi ( \u2212 \u03d5)\u03c12 2 . (3) The Tp\u03c1( \u2212 \u03d5) force balances the pressure force acting on the circle sector (Figure 5) whose radius is \u03c1 and the angle is ( \u2212 \u03d5). D ow nl oa de d by [ Fl or id a St at e U ni ve rs ity ] at 0 0: 16 0 8 O ct ob er 2 01 4 Vehicle System Dynamics 19 To proceed with the control volume equilibrium considerations, it is opportune to examine the \u03d5 = case, represented in Figure 6. In this situation, Tp tension is related to the tread and not to the sidewalls and it can be regarded as the sum of the two contributions: Tp = T \u2217 p + Tp. (4) In Equation (4), the quantities T \u2217 p and Tp assume the following expressions: T \u2217 p = \u03c1pi 2 ( 1 + v w ) sin \u03d5 Tp = pir \u2032 v w , (5) Tp tension can be interpreted as the force, for length unit, balancing the pi action on the aforesaid trapezoidal surface characterised by basis w and v" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003891_978-94-007-1643-8_22-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003891_978-94-007-1643-8_22-Figure1-1.png", "caption": "Fig. 1 Description of Monimad a Robotic assistant for sit-to-stand and walking", "texts": [ " In Section 4, we present the preliminary empirical results that we obtained with the method. These results are discussed in Section 5 before we conclude. Monimad is an original assistive device, combining sit-to-stand transfer and walking aid for elderly and disabled people. It allows mobility rehabilitation and assistance, safe walking (postural stabilization) and safe sit-to-stand transfer [11]. The designed robotic system is basically a two degrees of freedom arms mechanism mounted on an active mobile platform (Fig. 1). The lower part (mobile platform) consists of two electric motors actuating the wheels. The upper part of the mechanism (articulated arms) consists of two plane parallelograms to maintain the handles horizontally, actuated by linear actuators. In addition, they are independent in order to restore lateral balance. The end effectors consists of handles equipped with a six axis forces/torques sensor that are used in the experiments presented below to evaluate the relevance of the sit-to-stand trajectory" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000769_1.1631018-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000769_1.1631018-Figure1-1.png", "caption": "Fig. 1 Configuration of a fluid-film journal bearing", "texts": [ " Moreover, in the absence of inertia effects and the usual thin-film lubrication approximation, the Reynolds equation may be written as R22 ] ]u S H3 12m ]p ]u D1 ] ]z S H3 12m ]p ]z D52 U2 2R ]H ]u 1V2 (1) where R denotes the bearing radius, H the film thickness, m the constant fluid-viscosity, p the pressure of the fluid film, U2 and V2 are the fluid velocities at the journal surface in the tangential and radial directions. Now the journal rotates counter-clockwise at the speed of v, and the journal center is situated at ~e,g! in the polar coordinate system ~Fig. 1!. Then, those velocities are explicitly given in dimensionless form U\u03042512 Cr R ~\u00abg8 cos w1\u00ab8 sin w!>1 (2) V\u030425\u00ab~g821 !sin w1\u00ab8 cos w (3) 004 by ASME JANUARY 2004, Vol. 126 \u00d5 125 s of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F where w5u2g , U25RvU\u03042 , V25CrvV\u03042 , and e5Cr\u00ab with Cr being the bearing clearance; the prime denotes the differentiation with respect to dimensionless time t5vt . For later convenience, Eq. ~3! is rewritten in the rectangular coordinate system as follows: V\u030425 x\u0303 cos u1 y\u0303 sin u1 x\u03038 sin u2 y\u03038 cos u (38) where x\u03035\u00ab sin g and y\u030352\u00ab cos g", " In particular, when the system is in a steady state, the reaction force is mathematically expressed in dimensionless form F\u030405AF\u0304x0 2 1F\u0304y0 2 (17) with F\u0304x05E E p\u03040.0 p\u03040 sin ududz\u0302 and F\u0304y052E E p\u03040.0 p\u03040 cos ududz\u0302 (18) where the subscript 0 denotes the steady state, and F\u0304x0 and F\u0304y0 are x and y components of the reaction force F\u03040 , respectively. The dimensionless force F\u03040 is related to the dimensional one F0 as F05F\u03040~24mvR4/Cr 2! (19) It is assumed that the reaction force F\u03040 is directed along the negative Y axis ~Fig. 1!. Transactions of the ASME s of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F 2.2 Optimization Problem. An optimization problem is formulated for an inner configuration of fluid-film journal bearings. The configuration is exactly given by determining the bearing radius, the distance from the center to the inner wall of the bearing, in the circumferential direction ~Fig. 1!. Moreover, the bearing clearance is given by subtracting the journal radius from the bearing radius if the journal center is placed at that of the bearing. The journal cross-section is now assumed to be a full circle, in other words, the journal radius is now constant. The optimization problem then leads to finding an optimal configuration of the bearing clearance. A Fourier series is now used to represent an arbitrary clearance configuration (0 0, such that xa E T(x) implies that (4) s(t, xo, 0) E T(z) for all t 2 to " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002332_j.mechmachtheory.2009.05.007-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002332_j.mechmachtheory.2009.05.007-Figure9-1.png", "caption": "Fig. 9. Photograph of kinematic model of a steam engine with valve control.", "texts": [ " Large collections of these copies of Redtenbacher models can be found in Italy at the Foundation for Science and Technology {FST} in Florence and at the University of Porto in Portugal. Cornell University has 18 Schr\u00f6der models including the noncircular gear pair, shown in Fig. 2 that can also be found illustrated in Redenbacher\u2019s textbooks. Thus unwittingly, Redtenbachers\u2019 influence extended beyond Karlruhe through his designs for teaching models of machines. His models largely consist of kinematic mechanisms including linkages, gear systems and straightline mechanisms. However there are cutaway models of steam engine valve systems (see Fig. 9) as well as mathematical function synthesizers shown in Fig. 10. Redtenbacher did not have a mathematics or topologically based system of categorizing kinematic mechanisms as did his student Franz Reuleaux, [19,20] enhancing [35] the topology of the French Monge school which was based on the ideas of the great physicist Andre Marie Ampere. Redtenbacher was continually interested in explaining the connection between nature and techniques on a consistent basis. Such wide ranging thoughts permanently inspired his research activities and led him to complex and variable methodology" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003320_gt2010-22058-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003320_gt2010-22058-Figure3-1.png", "caption": "FIG. 3 SCHEMATIC OF STATIC TEST RIG", "texts": [ " Leakage is measured by Rheonik mass flow meter with 0.3% uncertainty. The seal was tested on the two test rigs for the static and dynamic identification of rotordynamic coefficients at the Institute of Energy Systems, Technische Universit\u00e4t M\u00fcnchen, Germany. A detailed description of the test rigs and the experimental investigation can be found in [29]. Despite low pressure differentials and air used as a working medium, the experimental work was carried out within the project on steam turbines. Static Test Rig Figure 3 shows a schematic of the static test rig. A symmetrical rigid rotor is supported by ball bearings and driven by a variable-speed direct-current motor with the maximum rotational speed of 12000 rpm. The maximum pressure differential is 0.9 MPa. Hydraulic cylinders are used to move and adjust the testing assembly in order to preset rotor eccentricity. The value of rotor eccentricity is measured by the eight eddy current displacement sensors. The testing assembly consists of the inflow casing and two identical seals" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001000_tmag.2004.840315-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001000_tmag.2004.840315-Figure1-1.png", "caption": "Fig. 1. HDD motor distribution based on inch sizes and number of discs.", "texts": [ "ndex Terms\u2014Bubble lubrication, hydrodynamic bearing, port. I. INTRODUCTION AS SEEN in Fig. 1, the basic challenges for spindle motor manufacturers are currently high-speed performance and durability for high-grade hard disk drives (HDDs); low-cost manufacturing methods for 3.5-in standard HDDs; and compact HDD spindle designs with greater tolerance of temperature and pressure changes for mobile use. This paper reports the basic design of a hydrodynamic bearing with an oil circulation mechanism that has the state function of expelling air and eliminating bubbles inside the bearing cavity, which affects the disruption-free nature of the oil film in the bearing", " The pumping pressure, flow rate of the circulating oil during rotation, and oil-sealing capability is analyzed next. 0018-9464/$20.00 \u00a9 2005 IEEE Using the calculus of the differential equation method, Fig. 3 shows the results of analysis, the internal pressure of an existing bearing without a circulation port (thin black line), and that with a circulation port (thick gray line). Since the hydrodynamic bearing may contain a bubble inside [3], [8], its numerical analysis method may or may not assume bubbles to be present in the oil [8], [10]. The developed bearing shown in Fig. 1, however, has the capability of eliminating internal bubbles and, therefore, does not contain any bubbles inside the bearing. In this paper, therefore, numerical analysis is conducted without taking into consideration the effects of bubbles [1], [2], [9], [11]. The flow path resistance in the circulation port and adjacent narrow gap parts is substantially lower than that in the gap of the radial bearing, with the result that the pressure at the circulation port is near atmospheric pressure. The pressure indicated by \u201cP\u201d in the Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000326_1.1897410-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000326_1.1897410-Figure2-1.png", "caption": "Fig. 2 Traction-drive speed reducer with midplate in place", "texts": [ " A typical inner roller is labeled F, and G is an outer roller. Inner and outer rollers are held in proximity to one another by the roller plate, H. The outer hole, J, through which the outer roller shaft protrudes, is slightly elongated. This allows the inner and outer rollers to press firmly against one another, generating sufficient friction to transmit torque through the unit. The roller plate, along with the inner and outer rollers, is free to rotate about pin, I. The midplate, in which the three pins are mounted, is shown in Fig. 2. The slots are elongated to allow the roller plates and inner and outer rollers to seek a configuration with the inner roller pressed firmly against the input member, and the output roller pressed firmly against the output member, part K. As the input member rotates counterclockwise, the intermediate roller assemblies rotate clockwise. This wedges them between the input and output members, producing the self-actuating feature. Because of the inclined orientation angle of the intermediate roller assemblies, the clockwise torque generated on the output member by the downstream load will produce the necessary normal forces to press the rolling members together" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002478_jjap.46.4698-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002478_jjap.46.4698-Figure7-1.png", "caption": "Fig. 7. Measurement system for load characteristics.", "texts": [ " The plate end, which has a contact area of 1 0:78mm2, is connected to the surface center of the vibrator by the conductive adhesive material (Dotite FA705). The acrylic holder clamps both the plates adhering on the top and bottom surface of the vibrator, and is mounted on the linear slide base. After the vibrator is attached to the holder, resonance characteristics are measured and shown in Table II. It is indicated that Q and resonance frequency decrease by supporting the holder; therefore, we have to improve the construction for supporting the vibrator in the future. The measurement system used is shown in Fig. 7. The torque of this motor is too small to measure directly. Therefore, load characteristics are estimated from the transient responses of the revolution speed measured with an optical rotary encoder and the input electric power.12,13) In this system, the optical encoder with a code wheel (AVAGO HEDS-9100-360) can output 360 pulses per revolution, and a measurement recorder counts the pulses in 2ms periods. The rotor consists of the code wheel and a thin stainless steel shaft of 1mm diameter, and its moment of inertia J is 4:15 10 8 kg/m2" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000532_0954405041897130-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000532_0954405041897130-Figure2-1.png", "caption": "Fig. 2 View of the process of laser melting", "texts": [ " The laser beam spot was set on the upper surface of the sample, and at the focal point the beam spot size Table 2 Technical parameters of the HPDL Rofin Sinar DL 020 Parameter Value Wavelength 808 5 nm Output power (continuous wave) 2500W Range of power 100\u20132500W Area of the focus spot 1.8mm 6.8mm Focus distance 82mm Proc. Instn Mech. Engrs Vol. 218 Part B: J. Engineering Manufacture B07903 # IMechE 2004 at CAMBRIDGE UNIV LIBRARY on August 15, 2015pib.sagepub.comDownloaded from was 1.8mm 6.8mm and had a uniform energy distribution (Table 3 and Fig. 2). After the laser melting process, cross-sections of the specimens were prepared for metallographic observation by cutting samples in direction perpendicular to the laser scanning direction. The samples were mechanically polished with a series of SiC abrasive papers (Nos 400, 600, 800, 1000, 1500 and 2000), then polished to 3mm with diamond suspension and finally etched by FeCl3 reagent. The microhardness of the specimens was examined over the laser-melting depth direction of the samples; the microstructure of the samples was observed with an Olympus optical microscope" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000407_6.2005-6074-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000407_6.2005-6074-Figure4-1.png", "caption": "Fig. 4 Distributed flight control architecture.", "texts": [ " In all other respects, such as power, communication, and data acquisition, the configurations were identical. For the car-top testing the vehicle was mounted at its center of gravity to a frame which was then attached to a car, as shown in Figure 3. This provided a much more controlled environment for eval- uating controller designs. The effects on the vehicle dynamics are discussed in more detail in section describing the analytical model. The distributed control architecture is shown schematically in Figure 4. In this configuration the pilot commands received by the central micro-controller are broadcast to both the logging micro-controller and to the distributed agents through a communication and power bus running the length of the trailing edge. The power to the distributed agents is provided by centralized batteries, although the power supply could also be distributed through the airframe. The bus provides eight possible positions for the distributed agents including three positions in each outboard wing and two in the centerbody" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002255_j.reactfunctpolym.2008.06.001-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002255_j.reactfunctpolym.2008.06.001-Figure3-1.png", "caption": "Fig. 3. CVs of bare GCE (a, c) and PEBTFE (b-e) in nitrogen-saturated phosphate buffer containing 0 lM (a, b), 7.2 lM (d), and 10.8 lM (c, e) NO. Scan rate: 0.05 V s 1.", "texts": [ " Unlike polymers of some other naphthol derivatives obtained in acetonitrile [26\u2013 28], the PEBTF showed only small and broad instead of sharp-shaped redox peaks in 1 M HCl. In L form, naphthyloxy groups at the two adjacent naphthyl rings could be oxidized to naphthoquinone, thus the polymer was electro-active. While in G form, the oxidation of naphthyloxy to naphthoquinone failed, thus the polymer was electroinactive. In aqueous solutions, pathway (a) dominated the polymerization process [26]. Therefore, G form was the main component of the polymer and L form was exiguous. As a result, the redox activity of the PEBTF was very poor. Fig. 3 shows the cyclic voltammograms (CV) of bare GCE or PEBTFE in a deaerated 0.1 M PB (pH 7.4) in the absence and presence of NO. At bare GCE, only a very small and broad anodic peak was obtained (curve c). While at PEBTFE, a sharp-shaped oxidation peak was observed at 0.8 V (curve d) and the peak current increased with the increasing NO concentration (curve e). These experimental results suggest that PEBTF has an excellent electrochemical activity toward the oxidation of NO. Fig. 4 shows the cyclic voltammograms of electrochemical oxidation of NO at PEBTFE at different scan rates" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001410_0301-679x(77)90021-4-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001410_0301-679x(77)90021-4-Figure5-1.png", "caption": "Fig 5 Scanning electron micrographs of replicas taken from", "texts": [], "surrounding_texts": [ "Initial wear of gears\nS. Andersson*\nRunning-in experiments have been made in a back-to-back gear test rig with hobbed and shaved gears of quenched and tempered steels at different speeds and loads. Measurements of surface roughness and examinations of surface replicas were used to study the running-in process. The present results show that the running-in period corresponds to less than about 0.3X 10 6 revolutions. The initial wear is rather small but increases with running speed and applied load\nInitial running of gears causes final finishing of the tooth flanks: they are smoothed down by flattening and wearing of surface asperities. The topography of the tooth flanks determines the service life with respect to surface failures, yet the changes taking place during running-in are rtot well known.\nAn understanding of the initial running period would be of value both from a theoretical and a practical point of view. By knowing the appearance of tooth flanks after running-in for safe running gears, it may be possible to predict service life of new gears by examining tooth flanks after running-in. This is of special interest for large gears.\nThe aim of the investigation described in this article was to study the running-in process and to test some recognised measuring methods which could be of use both on gears in the laboratory and for gears in service.\nGear test rig\nThe gear test fig used is of the usual back-to-back type (Fig 1). The load is applied by a hydraulic torque device. Each gear box had a separate lubrication system and the inlet oil temperature could be kept constant. The running conditions, applied load, driving motor torque, temperature of the oil at inlet and outlet, etc, were continuously recorded. The test gears were made of quenched and tempered steels SIS 2216 and SIS 2240 (Table 1) with a Brinell hardness of about 250. The lubricant was a mineral oil (type Nyniis TD-35X) with a viscosity 50 cSt at 50\u00b0C.\nStudying surface changes\nIn order to measure the surface roughness a special holder was made for a Pert-o-Meter profilometer (Fig 2). Since the equipment had to be inclined about 30 degrees, careful\nAtlas Copco AB, Central Materials Laboratory, Stockholm, Sweden This investigation was initiated by Stal-Laval Turbin AB and carried out at the Department of Machine Elements, The Royal Institute of Technology, Stockholm\nchecks of the surface roughness measured were made on reference surfaces. It was found that the recorded values were not influenced by the inclination. However, before every series of measurements a check was made of the equipment at the same inclination as when measuring on\nM - - - \" f ~ b . . . . . _ M\n2276\nGearbox,- ' I / / ' SIS 2240 J ~ . ~\nb ~\"1 b I \" - - Gear b\u00b0x 2 J ~ ' (\n- ~ L r \" - ~ . ~ - Reference - ~ - ' - - \" ~'-\"-~-'J~-~=~- , diameter d : 144 mm I r~l'[L'~,.J JJ~OJ'l~ Facewidth b : 40 mm \"o] ~ L X . q ~ - - ~ . . . - - - Metric module in J ~ \"1 l ~ V-'---I a normal section m n : 5 mm t . ~ to the flanks\nHelix angle /~ : 30 \u00b0 Nominal pressure angle a n : 20 \u00b0\nFig2 ProJ~lometer inclined about 30 \u00b0 in a special holder (left) used for direct measurement o f the surface roughness o f the flank in the test gear box at the approximate posilions shown (right)", "the positions marked T,C, and F are shown in Figs 6-9\nthe gears. The pick-up used recorded the surface roughness by a needle with respect to a spherical supporting surface. The measuring length, that is the distance the pick-up was moved over the flank at each measurement, was 1.5mm: the roughness of the same chosen flanks were measured before every running interval. Six flanks on each gear were measured at four positions giving 96 values of the flank roughness from which the mean roughness was determined. At each position both the R a value and the H value, a measure of the peak-to-valley roughness, were recorded.\nSurface prordes obtained by profilometers cannot give a complete view of the surface studied. As a complement,\nmicroscopic examinations were used. It is inconvenient to dismount the large test gears after each running interval: replica techniques had to be used. Using acetone soluble acetyl cellulose Ftlms, replicas could be made easily and quickly I.\nIt should be pointed out that when using replicas it is generally sufficient to examine the replica surface magni-\nTRIBOLOGY international August 1977 207", "fied with a standard optical projector: anyone familiar with the view of magnified surfaces can draw important conclusions about the surface conditions. Replicas can be studied in more detail in optical or scanning electron microscopes after depositing a thin layer on the replica surface.\nTest procedure\nA given test series started with a mesh control of the mounted test gears. The cycle of steps then followed were:\nmeasure the surface roughness of chosen teeth; make replicas of some chosen flanks; circulate the oil and heat it to about 45\u00b0C; run the gears at desired speed and apply the torque; stop the gears and oil system after running a certain time\nand let the equipment cool down; measure the surface roughness of the chosen teeth; make replicas of some chosen flanks.\nT\nC\nResults with tooth flanks as hobbed\nBefore running the test gears with as hobbed flanks, the mean roughness values were H = 4/am and R a = 0.65/am approximately. Three test series were made with a torque corresponding to a maximum Hertzian pressure of about 430 MPa (calculated according to SMS 18714) at speeds of 4 000, 2 000 and 1 000 rpm. The changes of surface roughness at 4000 and 2000 rpm are shown in Fig 3. H values of surface roughness are given since the changes of these values are greater than the corresponding changes of the more common R a values.\nThe decrease of the surface roughness during running-in was quite small, 7-16%, in the cases studied, and it was found that the changes of surface roughness are greater with increasing running speed (Fig 4). Since the theoretical film thickness, which would exist if the surfaces were\nsmooth, increases with increasing speed, the results obtained were somewhat puzzling.\nWith respect to the roughness, the running-in period seems to be less than 0.3x 10 6 revolutions. This result was confirmed by the study of the surface replicas.\nIn order to study the influence of pressure a test series was made by first running with the maximum Hertzian pressure as low as about 200 MPa and at the running speeds of 2 000 and 4 000 rpm. Surface roughness measurements and replica\n0 I 2 3\n, z\n.( , \"\n,?\n&\n\" t ; \"\n208 TRIBOLOGY international August 1977" ] }, { "image_filename": "designv11_20_0001889_ac60340a007-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001889_ac60340a007-Figure5-1.png", "caption": "Figure 5. Arrhenius plot of conductance data for the ethanolysis of acetyl chloride", "texts": [ " The same ethanol sample could be used for several runs by use of an off-set control without loss of sensitivity. Rate constants for each run were determined by plotting ln(G, - G ) us. time where G is the conductivity. The negative slopes of the straight lines resultant from these plots equal k , the rate constant. Rate constants were determined a t each of seven temperatures, and this collection of da t a was used to determine the activation energy from an Arrhenius plot. A linear least-squares routine was used to determine the activation energy. The results are shown in Table I1 while Figure 5 is the Arrhenius plot. Except for the activation energy calculation, all calculations were done to slide rule accuracy. A computer was used to calculate the activation energy. .4 comparison of our results with those obtained by other workers (5 , 6, 8) is shown in Table 111. Our value for (8) E. Euranto and R . Leimu, Acta Chem Scand , 20, 2029 (1966) ANALYTICAL CHEMISTRY, VOL. 46, NO. 4, A P R I L 1974 545 the activation energy agrees well with tha t obtained by Daum and Kelson while the results of Johnson and Enke agree more closely with those of Euranto and Leimu" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003887_19346182.2012.663534-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003887_19346182.2012.663534-Figure9-1.png", "caption": "Figure 9. The ball crosses the baseline at height H after bouncing up off the court at speed v and angle u. In a kick serve, v is about 20 m s21, u is about 208 and H is about 5 or 6 ft. The ball bounces with topspin, at about 600 rad s21 (5730 rpm). The distance from the bounce point to the baseline is about 6 to 8 m (20 to 26 ft). The perpendicular distance from the net to the service line is 21 ft (6.4 m). The perpendicular distance from the service line to the baseline is 18 ft (5.49 m).", "texts": [ " The Magnus force, FM, and its direction were determined from the magnitude of the spin and the inclination of the spin axis. If the ball is traveling in the horizontal direction then the vertical component of the Magnus force is given by FV \u00bc FMsinb where b is the inclination of the spin axis to the vertical. For example, if the spin axis is tilted by 308 then FV \u00bc 0:5FM so the vertical force on the ball due to spin is half the total force. After the ball lands in the service box, it bounces up off the court at speed v, at an angle u, with topspin v, and then crosses the baseline at height H, as shown in Figure 9. In general, the height H increases as v increases, it increases as u increases, and it decreases as v increases. Topspin causes the ball to kick up at a steep angle when the ball bounces, but after the ball bounces the effect of topspin is to reduce the bounce height. Effects of varying these parameters are shown in Figure 10. Figure 10(a) shows the effect of changing the bounce speed, assuming that the ball bounces at u \u00bc 208 with 600 rad s21 (5370 rpm) of topspin. The ball then travels a horizontal distance of either 6" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000608_155022891010015-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000608_155022891010015-Figure2-1.png", "caption": "FIG. 2. Mass-spring model of the rolling bearing.", "texts": [ " A schematic diagram of rolling element bearing is shown in Fig. 1. For investigating the structural vibration characteristics of rolling element bearing, a model of bearing assembly can be considered as a spring mass system, in which the outer race of the bearing is fixed in a rigid support and the inner race is fixed rigidly with the rotor. A constant radial force acts on the system. In the mathematical modeling, the rolling element bearing is considered as a spring mass system, and rolling elements act as nonlinear contact spring as shown in Fig. 2. Since the Hertzian forces arise only when there is contact deformation, the springs are required to act only in compression. In other words, the respective spring force comes into play when the instantaneous spring length is shorter then its unstressed length, otherwise the separation between rolling element and the races takes place and the resultant force is set to zero. A real rotorbearing system is generally very complicated and difficult to model, so for effective and simplified mathematical model the following assumptions are made: 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003271_s00170-011-3475-3-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003271_s00170-011-3475-3-Figure3-1.png", "caption": "Fig. 3 The location according to the left involute of the end mill (plane view)", "texts": [ " In these expressions, rb and ro were written as follows [23]: rb \u00bc 0; 5:m:N :Cosf \u00f04\u00de ro \u00bc 0; 5:m:N \u00fe m \u00f05\u00de b in Eq. 3 was described as a constant and 0\u2264bn\u2264bmax interval for boundary condition was given, where, bmax was described as maximum face width of the spur gear [22]. 2.3 The tool path equations for the left tooth flank Similarly, the location on the left involute curve of the end mill was taken into consideration while parametric tool path expressions for the left tooth flank of the gear were derived (see Fig. 3) [22]. The parametric expressions of the tooth path for the left tooth flank of the gear from Fig. 3 were derived as follows: Xp0 l \u00bc rn:Cos 2invamax \u00fe q invan\u00f0 \u00de \u00fe Re:Sin an 8\u00f0 \u00de \u00f06\u00de Yp0 l \u00bc rn:Sin 2invamax \u00fe q invan\u00f0 \u00de \u00fe Re:Cos an 8\u00f0 \u00de \u00f07\u00de Zp0 l \u00bc b Constant\u00f0 \u00de \u00f08\u00de where, to calculate 8 in Eqs. 6 and 7, angles corresponding to a tooth thickness and width of space on the spur gear were taken into consideration (see Fig. 4). The angles 8 and \u03b8 were derived as 8 =2inv\u03b1max\u2212 invan and \u03b8=\u03b8p\u22122(invamax\u2212inv\u03b8p) from Figs. 3 and 4. In a related paper, rotation procedures to Eqs. 1, 2, 3, 6, 7 and 8 were applied to cut the whole of teeth" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001196_1.2346690-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001196_1.2346690-Figure6-1.png", "caption": "Fig. 6 \u201ea\u2026 Labyrinth seal versus", "texts": [ " However, ince the displacement was applied only in one direction, only one f the CCS coefficients could be determined, which cannot be sed to determine the presence of destabilizing forces unless the ther coefficient is known and its sign . The knowledge of the ther coefficient requires a separate test in the orthogonal direcion, which was not conducted in their work. CS in Pocket Damper Seals Pocket damper seals 1 , patented and trademarked under the ame TAMSEAL\u2122, have been used in a number of centrifugal ompressor applications, primarily at the center seal location on ack to back machines and also on balance pistons, to attenuate otor vibration response and to solve rotordynamic stability probems. The PDS as shown in Fig. 6 is fabricated using baffle walls etween paired blades to restrict circumferential fluid flow develped by rotor rotation and impeller stage pre-swirl. In addition to estricting circumferential flow, incorporating baffle walls within aired blades gives rise to larger valued dynamic pressure oscilations in the cavities during machine operation. A notch in the ownstream blade of each pair produces a phase angle of the ressure that yields significant direct damping coefficients. Research has shown that the CCS coefficients produced by PDS n the X and Y directions possess the same sign" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000976_6.2004-4822-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000976_6.2004-4822-Figure1-1.png", "caption": "Fig. 1. Payload", "texts": [ " A test flight commences by launching the system from the ground, a 10-inch propeller powers the test system to altitudes of 250 to 400 ft where the propeller is stopped and gliding commences, lasting approximately 20 seconds for every 100 feet of altitude. Full state measurement of the parafoil required in the optimal control sequence is achieved through a sensor package that includes three single axis gyroscopes a three-axis accelerometer and a three-axis magnetometer shown in Figure 3. Inertial positions x and required in the mapping of the desired x-y path into a desired yaw angle are obtained from a Wide Area Augmentation System (WAAS) enabled Global Positioning Satellite (GPS) receiver shown in Figure 1. The sensors are supplemented with a wireless transceiver that transmits data from the parafoil and receives commands during flight. An operator controlled transmitter switches control of the parafoil to one of three modes: manual, estimation or autonomous. Manual mode allows the operator to manually fly the parafoil. Estimation mode allows estimation of linear model aerodynamic coefficients required for model predictive control. Autonomous mode controls the parafoil using the optimal control calculated from the model predictive control law" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000224_1.2114987-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000224_1.2114987-Figure3-1.png", "caption": "Fig. 3 Tooth profi", "texts": [ " The coordinates of the measured point after rotation are x3 and y3 and are calculated from x3 = cos \u2212 t \u00b7 x1 + X + y1 \u2212 Y tan \u2212 t , 5 y3 = cos \u2212 t \u00b7 y1 \u2212 Y \u2212 x1 + X tan \u2212 t . 6 The continuous monitoring of the dimension Y as the gear rotates can therefore be used, by substituting in Eqs. 1 \u2013 6 , to calculate the coordinates of the measured tooth profile. Optical Engineering 103603-3 ed From: http://opticalengineering.spiedigitallibrary.org/ on 05/15/2015 Term Measurement Setup he schematic drawing in Fig. 3 shows the main elements f the setup used to measure the tooth profile. The setup is esigned to provide specific angles of incidence and obserations that, in turn, depend on the nominal parameters of he test gear. It must also be capable of positioning the maging lens at a known distance f from the inspected gear ooth. To ensure that the setup fulfill the theoretically estimated onditions, the optical elements are fixed to a specially deigned template. The laser beam is first guided through two arrow slits, thus adjusting the angle of incidence with repect to a mirror" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002182_j.wear.2007.11.005-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002182_j.wear.2007.11.005-Figure1-1.png", "caption": "Fig. 1. Contact between a rough deformable surface and a rigid flat plane.", "texts": [ " Here t is assumed that when two rough surfaces contact with each ther under the application of a normal load, actual contact ccurs at the peak of the asperities of the rough surfaces. Due to he application of load the asperities will be deformed and the eformation may be elastic, plastic or elastic\u2013plastic depending n contact pressure, material and surface properties. In such a odel, the two rough surfaces in contact are represented by an quivalent rough surface in contact with a smooth plane as shown n Fig. 1. The model assumes that the asperities are far apart, and here is no interaction between them and no bulk deformation ccurs during contact. In this model asperities are assumed to ave spherical cap of radius R and their heights (z) to follow aussian distribution with a standard deviation of their heights, . The normalized Gaussian distribution of asperity heights is iven by \u00af (z\u0304) = 1\u221a 2\u03c0 exp [ \u2212 z\u0304 2 2 ] (1) imensionless asperity height. Following the work of Yu and Polycarpou [16], an asymmetic distribution may be described by the following normalized 5 ar 26 W \u03c6 w e t G n 2 2 e s fl d c R i o f a d a r u t d a i s a [ t ( \u03c6 a \u03c6 w 2 b b p t o d w E o s c d o 3 u s \u03b8 [ e r t e u b s a p W t t c p r e e i e t p b n p v p t m e l t u a d 56 S", " [18] of contact between a smooth sphere and a at in the presence of surface forces, the load on elastically eformed asperities may be obtained and the load on a plastially deformed asperity may be obtained from the analysis of oy Chowdhury and Pollock [21]. The contact load of an asperty in elastic\u2013plastic regime may be obtained from the analysis f Kogut and Etsion [27]. Using all these results, total loading orce (Pl) may be obtained. Considering the contact between rigid smooth surface and a rough deformable surface with a istribution \u03c6(z) of asperity heights z as shown in Fig. 1, the sperity deformation (\u03b4) is \u03b4= z \u2212 d, where d is the mean sepaation between the surfaces. If N is the number of asperities per nit area of the rough surface, then the total applied load on all he asperities in contact per unit area may be obtained in nonimensional form as P\u0304l = Pl/AnH , where An is the nominal rea (taken as unity in the present case) and H is the hardness. Pl s given by Eq. (9) in Ref. [31]. In using the above normalization cheme for loading force, a value for NR\u03c3 needs to be supplied nd the same is taken as 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000606_1.1789972-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000606_1.1789972-Figure1-1.png", "caption": "Fig. 1 Model of one link Furuta pendulum", "texts": [ " As an example of global stabilization of underactuated systems, studying the control of such systems is significant because we can get benefits of designing control system achieving light weight, low cost, and energy saving. A one link rotated-type pendulum, i.e., Furuta pendulum, has sufficient nonlinearity regardless of its mechanical simplicity. Especially the swing-up motion yields a dynamic change on the internal forces, hence, this apparatus is adequate to evaluate a proposed method that considers the internal forces. 4.1 Derivation of SDLR and Internal Forces. A model of the one-link Furuta pendulum is shown in Fig. 1. The horizontal rotational bar and the pendulum link are called a \u2018\u2018link-0\u2019\u2019 and a \u2018\u2018link-1,\u2019\u2019 respectively. Motor drives the link-0 and the link-1 rotates freely at a joint. L0 , m1 , and l1 are the total length of the link-0, the mass, and the offset to the center of gravity ~COG! of link-1, respectively. Ji and di are link-i\u2019s inertia and viscous coefficient. u0 and u1 are angles of these links, and positional coordinate values of the link-1\u2019s COG are written as (x1 ,y1 ,z1). t is an input torque to the link-0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001266_bfb0039280-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001266_bfb0039280-Figure1-1.png", "caption": "Figure 1. (a). A planar two llnk robot. (b). The two-llnk robot in a singular configuration. The", "texts": [ " The Jacobian also gives the relation between velocities in Cartesian space, ~, and joint rates, ~, = a(q),d (3) The singularities of the robot are defined as those configurations, q, for which J(q) loses rank. The purpose of this paper is to further investigate the kinematic and dynamic properties of the robot close to singularities. We wiU later be interested in control possibihties, but we wiU start with a closer look at a fa~Liliar kinematics example (Asada and Slotine, 1986; Craig, 1989; Spong and Vidyasagar, 1989). A F a m i l i a r K i n e m a t i c s E x a m p l e Consider the planar two-link, m = n = 2, manipulator in Figure 1 a. Wi thou t loss of generality one of the link lengths is normalized to 1. The kinematics is zl = cosqt + tcosCqt +q2) z2 = sinqt + lsin(ql + q2) (4) leading to the Jacobian ~(q) = [ - sin qt - t sin(qt + q2) - t sin(qt + q2) ] J \u00a2o~ ql + t cosC.q, + q2) t \u00a2o~(qt + q2) (5) The inverse kinematics, i.e. q as function of ~:, is derived using the notat ional conventions in (Craig, 1989). Star t e.g. f rom the Cosine-theorem to first yield q~ as \u00a2osq2 and sin q2 from cos q2 i 2 = ~(=1 + =] - 1 - l 2) -- c(=) and slnq2 = 4 - ~ _- s(z). The final result is q, = At=2(=~, ~:d - At~.~.(z-.(~:), 1 + t~(~)) (6) q~ = A t ~ ( , ( ~ ) , ~(~)) Consider the robot in the singular configuration where the robot is fully s tretched along the z~-axis, as seen in Figure 1 b. We will now s tudy motions inward along the zl-axis from this singular configuration. Such a mot ion is obta ined if z2 = 0 i.e. if the joint angles are related as: sinql + l,in(ql + q2) = 0 (7) A n approz i rna te analys is Before developing an explicit expressions for a mot ion along the z r ax i s , an approximate analysis based only on the leading terms in a Taylor expansion of the kinetics (4) is enlightening. The analysis clearly indicates how different t ime functions i.e. different trajectories, (z~(t), z2(t)) may describe possible motions" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001893_s00170-008-1850-5-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001893_s00170-008-1850-5-Figure5-1.png", "caption": "Fig. 5 Cutting force Fv evolution around the nominal value Fn; test case using cutting parameters ap = 5 mm, f = 0.1 mm/rev, and N = 690 rpm", "texts": [ " Consequently, the cutting force component variations and the self-excited vibrations are influenced mutually, in agreement with [17, 19, 25, 30]. Also, in agreement with research on the dynamic cutting process [28], we note that the self-excited vibration frequency is different from the workpiece rotational frequency, which is located around 220 Hz. 2.3 Forces decomposition The force resultant components detailed analysis highlights a plane in which evolves a variable cutting force Fv around a nominal value Fn (see further). This variable force is an oscillating action (Fig. 5) that generates u tool tip displacements and maintains the vibrations of elastic system block-tool BT [7] (Fig. 6). Thus, the cutting force variable (Fig. 5) and the self-excited vibrations of elastic system WTM are interactive, in agreement with research work [18, 28]. The cutting force variable part can be observed and compared. Not to weigh down this part, the cutting forces analysis is voluntarily below restricted at only two different situations: \u2013 Stable process using cutting depth ap = 2 mm (Fig. 7a), \u2013 Unstable process (with vibrations) using cutting depth ap = 5 mm (Fig. 7b). The vibration effects on the variable forces evolution are detailed in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001470_app.1979.070240608-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001470_app.1979.070240608-Figure7-1.png", "caption": "Fig. 7. Load-extension curves for two different single fibers with and without convolutions removed: (a) normal Acala; (b) Acala after treatment.", "texts": [ " It should also be noted that the fiber strength has increased, indicating that during the extension process the deconvolution effect probably creates forces which readily lead to crack propagation and hence lower breaking loads for convoluted fibers. Indeed, many workers have correlated orientation with strength. If these orientation measurements are incorrect by a convolution angle factor, then there will be a correlation between strength and convolution angle. The initial moduli have been increased considerably, as shown.in Figure 7 in the load-extension curves for Acala 1517 fibers before and after the treatment. It should be pointed out that during this water treatment it is possible to have reduced the fibrillar spiral angle, which would also lead to an increase in modulus. Meredithll has shown an increase in dynamic bending modulus as spiral angle decreases for tension-mercerized cotton fibers. It is worth noting that the total extension to break of the water-treated fibers, namely, the permanent extension after the wet stretching plus the additional extension to break, is greater by 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002891_1.4000647-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002891_1.4000647-Figure3-1.png", "caption": "Fig. 3 Pitch lines and scroll profiles", "texts": [ " 2, for a point on a continuous curve, assume that its normal angle is the positive angle between the x-axis and the normal to the curve at this point and has a property of superposition. Therefore, an arbitrary point on this curve can be expressed as a complex number by the normal component, Rn, and tangential component, Rt, of its radius-vector as follows: P = Rn exp j + Rt exp j + 2 2 Rt = dRn d According to the general profile theory, there are three types of general profile, they are Rn = c0 + c1 + c2 for type-I of scroll profile Rn = c0 + c1 cos + c2 for type-II of scroll profile 3 Rn = c0 + c1 + c2 2 + c3 3 for type-III of scroll profile As shown in Fig. 3, the inner and outer profiles of scroll wraps are formed by shifting their pitch lines by a distance of Ror /2 in the normal direction of the pitch lines. The pitch line forming outer profile of orbiting scroll and inner profile of fixed scroll is defined as A-type pitch line and similarly, the pitch line forming inner profile of orbiting scroll and outer profile of fixed scroll is defined as the B-type one. For a scroll fluid machine, A-type pitch line or B-type one may have z identical pitch lines in number" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003863_b978-0-12-417049-0.00002-x-Figure2.3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003863_b978-0-12-417049-0.00002-x-Figure2.3-1.png", "caption": "Figure 2.3 (A) Position and orientation of a solid body, (B) position and orientation of a robotic end-effector (a5 approach vector, n5 normal vector, o5 orientation or sliding vector), and (C) position vectors of a point Q with respect to the frames Oxyz and O0xbybzb.", "texts": [ " Formally, the generalized inverse Jy of a m3 n real matrix J is defined to be the unique n3m real matrix that satisfies the following four conditions: JJyJ5 J; JyJJy 5 Jy \u00f0JJy\u00deT 5 JJy; \u00f0JyJ\u00deT 5 JyJ It follows that Jy has the properties: \u00f0Jy\u00dey 5 J; \u00f0JT \u00dey 5 \u00f0Jy\u00deT ; \u00f0JJT \u00dey 5 \u00f0Jy\u00deTJy All the above relations are useful when dealing with overspecified or underspecified linear algebraic systems (encountered, e.g., in underactuated or overactuated mechanical systems). The position and orientation of a solid body (e.g., a robotic link) with respect to the fixed world coordinate frame Oxyz (Figure 2.3) are given by a 43 4 transformation matrix A, called homogeneous transformation, of the type: A= R p 0 1 \u00f02:9\u00de where p is the position vector of the center of gravity O0(or some other fixed point of the link) with respect to Oxyz, and R is a 33 3 matrix defined as: R5 n ^ o ^ a \u00f02:10\u00de In Eq. (2.10), n, o and a are the unit vectors along the axes xb, yb, zb of the local coordinate frame O0xbybzb. The matrix R represents the rotation of O0xbybzb with respect to the reference (world) frame Oxyz. The columns n, o, and a of R are pairwise orthonormal, that is, nTo5 0; oTa5 0; aTn5 0; jnj5 1; joj5 1; jaj5 1 where bT denotes the transpose (row) vector of the column vector b, and jbj denotes the Euclidean norm of b \u00f0jbj5 \u00bdb2x1b2y1b2z 1=2\u00de, with bx, by, and bz being the x; y; z components of b, respectively. Thus the rotation matrix R is orthonormal, that is: R21 5RT \u00f02:11\u00de To work with homogeneous matrices we use 4-dimensional vectors (called homogeneous vectors) of the type: XQ 5 xQ yQ zQ ? 1 2 66664 3 777755 xQ ? 1 2 4 3 5; XQ b 5 xQb yQb z Q b ? 1 2 66664 3 777755 xQb ? 1 2 4 3 5 \u00f02:12\u00de Suppose that XQ b and XQ are the homogeneous position vectors of a point Q in the coordinate frames O0xbybzb and Oxyz, respectively. Then, from Figure 2.3C we obtain the following vectorial equation: ~OQ5 ~OO0 1 ~O0A1 ~AB1 ~BQ where ~OQ5 xQ; ~OO0 5 p; ~O0A5 x Q b n; ~AB5 y Q b o; ~BQ5 z Q b a: Thus: xQ 5 p1 x Q b n1 y Q b o1 z Q b a5 p1 n o a x Q b y Q b zQb 2 64 3 755 p1RxQb \u00f02:13a\u00de or XQ = n o a p 0 0 0 1 XQ b = AXQ b \u00f02:13b\u00de where A is given by Eqs. (2.9) and (2.10). Equation (2.13b) indicates that the homogeneous matrix A contains both the position and orientation of the local coordinate frame O0xbybzb with respect to the world coordinate frame Oxyz" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002090_9780470264003-Figure13.7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002090_9780470264003-Figure13.7-1.png", "caption": "Figure 13.7 Linear transfer function in a plane.", "texts": [ " Suppose that there is another design parameter, denoted by x2, and we are interested in the overall performance of y as a function of the two design parameters. In this case, y can be given as y f x x a a x a x= = + +( , )1 2 0 1 1 2 2 linear terms (13.2) 4Civil engineers calculate the slope of a roadbed by calculating the ratio of the distance it rises or falls to the distance it runs horizontally, calling it the grade of the roadbed. TRANSFER FUNCTION MATHEMATICS 303 304 DFSS TRANSFER FUNCTION AND SCORECARDS The plot of such a transfer function is given in Figure 13.7. As you may conclude, it is a plane. Note that this transfer function may be enhanced by a cross-product term reflecting the interaction between parameters or factors x1 and x2. In Chapter 14 we call this cross-product term interaction. This may happen when there is a chemical reaction. Equation (13.2) then becomes y f x x a a x a x a x x= = + + +( , )1 2 0 1 1 2 2 12 1 2 linear terms interactio n term (13.3) We often encounter transfer functions that change at a linear rate rather than a constant rate" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.19-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.19-1.png", "caption": "Fig. 14.19 Blade kinematics of a grader", "texts": [ " Due to their advantages regarding various arrangement possibilities and considerable converting range, hydrostatic transmission has become increasing important in these machines. A decisive factor for the vehicle\u2019s drive is the nominal thrust force, which determines the machine\u2019s power capacity. The nominal thrust force depends on the engine power and the machine weight. A 2 \u00d7 2 \u00d7 2 grader delivers a ratio of engine power to machine weight of about 9\u201310 kW/t. A 1 \u00d7 2 \u00d7 3 grader is expected to deliver a ratio of about 7\u20138 kW/t [14.19]. Figure 14.19 shows a grader\u2019s blade positions. The specialized kinematics facilitate a large variety of blade positions, making the grader a universally applicable machine. Depending on the chassis type, there are two types of scrapers: the wheel scraper and the tracklaying scraper with crawler, which is mainly utilized for difficult soil conditions. The excavating bucket consists of several parts. It has movable front and back sides in order to pour the material out of the container in a particular direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003140_iros.2009.5354270-Figure11-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003140_iros.2009.5354270-Figure11-1.png", "caption": "Fig. 11. Measurement of the tip position and angle of the giant-swing robot.", "texts": [ " However, it could be confirmed that the developed giant-swing robot has the ability to perform the giant-swing motion. A. Experimental Condition and Method We attempted to avoid the stagnant state at the bottom of the horizontal bar by applying various types of rewards. In this experiment, the rewards based on the robotic physical quantities\u2014decrease in swing height, swing angle, tip angle, and mechanical energy\u2014as listed in Table II were given to the robotic actions in the Q-Learning process. The tip angle t\u03b8 was calculated by means of the kinematic information of the robot, as shown in Fig. 11. In addition to this condition, the movable ranges of enabled motors were constrained as shown in Fig. 12 in order to imitate those of human beings; we expected that the robot may acquire a human-like motion. Similar to the experiment in chapter IV, the Q-Learning with \u03b5-greedy method was applied. Then, \u03b5 was reduced with the time transition at the rate of 2.0 \u00d7 10\u22126 per learning step. First, we executed the Q-Learning for each reward by using the ODE-based dynamic simulator of the giant-swing robot in order to reduce the learning time and to avoid the fatigue breakdown", " These results imply that the kinematic information of the robot might be significant to perform the giant-swing motion. With regard to the rewards except the tip-angle-based reward, they all use the swing angle of the arm; these rewards have only the information from the horizontal bar to the robotic arm. Hence, the robot was not able to know its own posture beyond the arms in the learning process. This condition might hinder jumping out of the stagnant action loop at the bottom of the horizontal bar. On the other hand, the tip-angle-based reward includes the information of robotic posture, as shown in Fig. 11. Thus, it is implied that the posture information enabled the comprehensive exploration in the learning process and might result in achieving the giant-swing motion. As the next step, we practically implemented the learning result obtained by the application of the tip-angle-based reward in the real robot. The result demonstrated that the real robot could make a revolution around the horizontal bar. However, the robot frequently stopped the movement due to the motor spec; the employed motor has the characteristic that automatically stops the motor drive if the intensive actions are taken many times" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003355_j.jeurceramsoc.2011.03.034-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003355_j.jeurceramsoc.2011.03.034-Figure3-1.png", "caption": "Fig. 3. The microprocessing graphics on the ultrathin ceramic.", "texts": [ " he laser has high stability, with both average power and ulse to pulse energy fluctuations of less than 2%. An ultrahin ceramic plate is fixed on a vacuum table by a clamper Fig. 2). The focal plane is fixed on the surface of the samles and the laser beam spot size at focus is approximately The physical properties of the 125 m thick Al2O3 ceramics sed in the experiments are listed in Table 2. The microprocessng graphics required on the ultrathin ceramic substrates were uite delicate and complex (Fig. 3), with many narrow slots and icro holes. The minimum slot width and hole diameter were 0 m and 240 m respectively. We surmised that laser micro- F. Zhang et al. / Journal of the European Ceramic Society 31 (2011) 1631\u20131639 1633 F s 2 m c s q t T O L A B C D F c 1 e ig. 4. The effect of different factors and levels on kerf width and Ra on the kerf idewall. .3. Experimental method A 4 \u00d7 4 orthogonal design was used to determine and optiize the key parameters for laser microprocessing of Al2O3 eramics (Table 3)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000759_imece2004-60714-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000759_imece2004-60714-Figure1-1.png", "caption": "Figure 1. Single Organized Microcapsule Array.", "texts": [ " Structured arrays of engineered microcapsules strong enough to hold higher capsule pressures (10 MPa and higher) with shells that mimic cell membranes in its transport function will enable us to develop novel actuators with tailored energy density and comparable strain rates. In biological parlance, transport of ions against the concentration gradient by ion pumps is called active transport and the resulting flow of species along the established concentration gradient is called secondary active transport. Osmotic regulation occurs at the end of active transport processes to balance energy potentials. Forcing active transport by adding ATP to an Organized Microcapsule Array (OMA) as shown in Figure 1 and facilitating transport from the resulting concentration gradient will result in an expansion of the organized array. The fluid filled interstices are maintained at the required pH for the biological reactions to take place. Assembling units of organized arrays as a mosaic shown in Figure 2 and selectively expanding the arrays will produce bulk deformation in the material. Force generated in each OMA contributes to the total force developed by the actuator. Since the force in the actuator is organic (from individual units of the mosaic) and has the possibility to be controlled individually; the mosaic 2 wnloaded From: http://proceedings" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002380_j.talanta.2007.02.027-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002380_j.talanta.2007.02.027-Figure1-1.png", "caption": "Fig. 1. Scheme of flow-injection CL system. (a) Sample or blank solution; (b) carrier (1 \u00d7 10\u22123 mol l\u22121 SDS solution); (c) luminol; (d) K3Fe(CN)6; (P1, P2) peristaltic pump; (V) injection valve; (C1, C2) confluence point; (F) CL flow cell; ( (", "texts": [ " The stock solution of SDS (Xi\u2019an Chemical ndustry Reagent factory, Xi\u2019an, China) is 1.0 \u00d7 10\u22122 mol l\u22121. edical used DA ointment samples (Shenyang Minghua harmacy Ltd., China (20051101), Shenyang Yaoda Pharacy Ltd., China (061019) and Nantong Zhongbao Pharmacy td., China (061001)) are purchased from local drug tore. t L s 0 PMT) photomultiplier tube; (BPCL) ultra-weak chemiluminescence analyzer; PC) personal computer. .2. Apparatus and flow injection system The flow system used in this work is shown in Fig. 1. There are wo peristaltic pumps P1 and P2 (Shanghai Qingpu Huxi Anaytical Instrument Ltd., China). One is used to deliver the flow treams of sample and carrier, and another is used to deliver he luminol and K3Fe(CN)6. PTFE tube (0.8 mm i.d.) is used s connection material in the flow system. The flow cell is a at spiral-coiled colorless silicon rubber tube (i.d. 0.8 mm; total ength of the flow cell, 6 cm, without gaps between tubes) placed lose to the window of PMT (Hamamatsu, Tokyo, Japan, operted at \u2212400 V)", "0 \u00d7 10\u22124 mol l\u22121 K4Fe(CN)6 n 3.0 \u00d7 10\u22123 mol l\u22121 NaOH solution; (d) mixture of 1.0 \u00d7 10\u22125 mol l\u22121 4Fe(CN)6 and 4.4 g ml\u22121 DA in 3.0 \u00d7 10\u22123 mol l\u22121 NaOH solution; nsert: the UV\u2013vis absorption spectrum in 290\u2013400 nm; (e) mixture of .0 \u00d7 10\u22124 mol l\u22121 K4Fe(CN)6 and 44 g ml\u22121 DA in 3.0 \u00d7 10\u22123 mol l\u22121 aOH solution. 5) D I n o o K g r K a n r d (2007) 1811\u20131817 More experiments to evaluate reactions (1) and (2) are caried out. When DA standard is firstly mixed with K3Fe(CN)6 ow instead of luminol in the presented flow system (Fig. 1), he obtained CL signal is much lower than that obtained withut above flow system change. By varying the tube length from A\u2013K3Fe(CN)6 confluence point to luminol confluence point C1 to C2 in Fig. 1), it is found that the longer the tube length, he weaker the CL observed. These two experiment results coinide with what we have suggested in reactions (1) and (2), and t can be concluded that the CL difference observed in above xperiments is attributed to the instability of proposed DAox. A + K3Fe(CN)6 K1\u2212\u2192DAox + K4Fe(CN)6 (1) Aox K2\u2212\u2192Pox (2) (3 (4) ( (6) Aox + I K7\u2212\u2192II (7) t is known in basic aqueous, luminol forms the dianion of lumiol (I) as shown in reaction (3). The dianion (II) can be reversibly xidized by K3Fe(CN)6 to the semidione structure shown in ne of its resonance form (II), and K3Fe(CN)6 is reduced to 4Fe(CN)6 (reaction (4))" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000839_j.euromechsol.2004.08.003-Figure12-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000839_j.euromechsol.2004.08.003-Figure12-1.png", "caption": "Fig. 12. Space-based manipulator.", "texts": [ " For \u03b2 = 2, the relation between e, c with the increase of \u03b7s , \u03b1, \u03b5 is shown in Figs. 8\u201310. It is shown that for each case, e increases with c, and that for |c| < 0.07, e is less than 0.05, which is equal to the error limit in frequency analysis. Furthermore, Fig. 11 shows that for \u03b2 = 0.5, e is less than 0.05 in case of |c| < 0.07. Thus, we make the conclusion that the linear method is suitable for the simulation of a three-dimensional clamped-free beam undergoing large overall motion for |c| < 0.07. A flexible space-based manipulator (Fig. 12) with large overall motions is simulated to verify the effectiveness of the criterion. The manipulator is composed of three flexible links: B1, B2 and B3. The time dependent influence ratio for the clamped-free beam B3 is defined as c(t) = g(t) h = 1.1931(\u03b72(t)2 + \u03b73(t)2) \u2212 1.5706\u00b51(t) min(12.3603,12.3603\u03b2) , (61) where \u03b71(t) = \u03c91(t)T , \u03b72(t) = \u03c92(t)T , \u00b51(t) = a01(t)T 2 l . (62) The motion of the three bodies are given by \u03a81(t) = \u03c0 \u2212 \u03c0 2T ( t \u2212 T 2\u03c0 sin 2\u03c0t T ) , 0 < t T, \u03c0 2 , t > T, (63) \u03a82(t) = \u03c0 \u2212 3\u03c0 4T ( t \u2212 T 2\u03c0 sin 2\u03c0t T ) , 0 < t T, \u03c0 4 , t > T, (64) \u03a83(t) = \u03c0 \u2212 \u03c0 T ( t \u2212 T 2\u03c0 sin 2\u03c0t T ) , 0 < t T, 0, t > T " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003905_ecc.2013.6669571-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003905_ecc.2013.6669571-Figure3-1.png", "caption": "Fig. 3. Kinematical structure of the drive train with gear boxes and disturbance torque.", "texts": [ " As the tilt angles of the diplacements units are limited, saturation functions satab (\u03b1\u0303i) are introduced as follows satab (\u03b1\u0303i) = \u03b1\u0303i,max \u03b1\u0303i \u2265 a \u03b1\u0303i for b < \u03b1\u0303i < a , \u03b1\u0303i,min \u03b1\u0303i \u2264 b (14) where a = \u03b1\u0303i,max and b = \u03b1\u0303i,min represent the upper and lower output limits determined by the mechanical design: {\u03b5M , 1} for the hydraulic motor and {\u22121, 1} for the hydraulic pump. In the simulation model, (13) is implemented with limited integrators for \u03b1\u0303P and \u03b1\u0303M . The longitudinal dynamics of the working machine is governed by the equation of motion. The vehicle with the drive train (vehicle mass mv , wheel radius rw, gear box transmission ratio ig , rear axle transmission ratio ia, damping coefficient dg at the drive shaft), see also Fig. 3, can be described by the following first order differential equation( JM + Jg ig 2 + Ja +mV r 2 w i2a i 2 g ) \ufe38 \ufe37\ufe37 \ufe38 JV \u03c9\u0307M + dg i2g\ufe38\ufe37\ufe37\ufe38 dV \u03c9M = (15) V\u0303M\u2206p \u03b1\u0303M\ufe38 \ufe37\ufe37 \ufe38 \u03c4M \u2212 ( \u03c4Mf tanh( \u03c9M \u03b5 ) + \u03c4gf tanh( \u03c9M ig\u03b5 ) + \u03c4L iaig ) \ufe38 \ufe37\ufe37 \ufe38 \u03c4U , where \u03c4M is the torque of the hydraulic motor. JM , Jg and Ja are the mass moments of inertia of the hydraulic motor, gear box and rear axle, respectively. The maximum values \u03c4Mf and \u03c4gf characterise the friction models of the hydraulic motor and the gear box, whereas \u03b5 << 1 represents a small number" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000788_la0604578-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000788_la0604578-Figure2-1.png", "caption": "Figure 2. Schematic illustration of (A) twist distortion of liquid crystal between two surfaces with easy axes \u03b7o and (B) angles that characterize the departure of the director of the liquid crystal from the easy axis at the top surface shown in A.", "texts": [ " In particular, we sought to determine if changes in the azimuthal anchoring energy of the nematic liquid crystal 5CB could be measured at surface coverages of biomolecules that are below the thresholds that lead to spontaneous changes in the ordering of the liquid crystals. The approach reported in this paper to measurement of the azimuthal anchoring energy of peptide-decorated interfaces is a variant of the elastic torque-balance method.27,29-31 In brief, when using the torque-balance method, liquid crystals are confined between two surfaces such that \u03b70 at the top surface is rotated approximately 90\u00b0 relative to the orientation of \u03b70 at the bottom surface (Figure 2A). This configuration induces a twist distortion across the film of liquid crystal. For films of liquid crystal that are sufficiently thin, the elastic bulk torque of the liquid crystal competes with the surface-anchoring torque such that the equilibrium position of the director \u03b7d deviates by an angle (\u00e6) from the easy axis of surfaces with finite azimuthal anchoring energies Waz. Whereas the conventional torque-balance method uses two identical surfaces, the methodology reported in this paper permits measurement of the anchoring energy on a single, patterned surface. In this modified procedure, the liquid crystal is confined by a second surface (reference plate) that strongly anchors the liquid crystal, such that deviation of the orientation of the liquid crystal from the easy axis only occurs at the surface of interest, as depicted in Figure 2B. The azimuthal anchoring energy is calculated as where K22 is the twist elastic constant for the liquid crystal, d is the thickness of the film of liquid crystal, and \u03c8 is the twist angle over which the director of the liquid crystal is rotated. The use of surfaces patterned with SAMs and peptides permits simultaneous measurement of multiple values of Waz in a single experiment, thus minimizing the volumes of solutions required for surface functionalization and analysis. Specifically, we apply this methodology to determine how the incremental addition of the peptide IYGEFKKKC (a substrate for Src kinase) to an interface leads to changes in azimuthal anchoring energy of a nematic liquid crystal", " The dark bands in Figure 4D correspond to regions of the patterned surface decorated with SAMs printed from hexadecanethiol. In these regions, the liquid crystal is not twisted (see Figure 4C), the polarization of light transmitted through the sample is not rotated, and thus light is extinguished by the analyzer.35 Using optical methods reported previously (based on rotation of the sample between the polarizers) and the sample shown in Figure 4D,27 we next measured the relative orientation of the easy axes of the two confining surfaces (see Figure 2B, angle \u03b4) in each of the six regions shown in Figure 4D to determine the uniformity of the obliquely deposited gold films used in these experiments. Our findings are summarized in the second column of the table shown in Figure 5A. These measurements reveal that there is a systematic variation in \u03b4 from the leftmost region of the substrate (Region 1) to the rightmost region of the substrate (Region 6), from 80.8\u00b0 to 85.4\u00b0, respectively. The observed variation in \u03b4 reflects small changes in the orientation of the easy axis (\u03b70) at the top (patterned) substrate, as depicted in the angle diagram in Figure 2B. The orientation of \u03b70 at the top surface is dependent on the in-plane direction of the deposition of gold used to create the top surface. We thought it likely that we would find evidence of small variations in the direction of gold deposition (as suggested by the above-described measurements) due to the geometry of the electron beam evaporator used to prepare the gold films (Figure 5B). During deposition of the gold films, the substrate is held at a distance of approximately 600 mm from the gold source", " First, we note that a reduction in sample size, or an increase in the distance between the source and substrate during deposition, will lead to smaller variations in \u03b4 within a single sample. Second, we emphasize that the above variation in the orientation of the easy axis does not prevent accurate determination of the anchoring energy of the liquid crystal. The anchoring energy is an intrinsic property of the interface and the methodology used below incorporates the measured variation of \u03b4 into the calculated values of the anchoring energy. We next determined the equilibrium orientation of the director (\u03b7d, expressed as angles \u03c8 and \u00e6, see angle diagram in Figure 2B) in each of the bright regions shown in Figure 4D. The orientation of the director was determined in each region at a fixed thickness d ) 5.5 ( 0.5 \u00b5m. Using the torque-balance expression (see eq 2) and the measurements for \u03b4, \u00e6, and \u03c8 summarized in Figure 5A, we determined Waz in each of the six tetra(ethylene glycol) regions shown in Figure 4D.36 The values of Waz also appear in Figure 5A and are graphically represented in Figure 5C. Error bars in the vertical direction reflect the uncertainties in our determination of angles \u03c8 and \u00e6 and thickness d" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003560_1.3607746-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003560_1.3607746-Figure1-1.png", "caption": "Fig. 1 Configuration of journal bearing showing the dimensions R, R0, e, hW); the angle 0; the angular velocity w; and the external loading on tha shaft Wx and V/,j", "texts": [ " Discussion of this paper should lie addressed to the Editorial Department, ASME, United Engineering Center, 345 East 47th Street, New York, N. Y. 10017, and will be accepted until October 15, 1967. Discussion received after the closing date will be returned. Manuscript received by ASME Applied Mechanics Division, April 19, 1966; final draft, December 19, 1966. Journal of Applied Mechanics 1 (duj i)M.\\ co,- = vorticity = - - \u2014 ) 11 = viscosity G = elastic shear modulus In this paper, an approximate solution is obtained for the foregoing problem using a perturbation procedure. Analysis Fig. 1 shows a cross section of the journal configuration under consideration with certain geometric relations. Neglecting end effects for this long bearing leads to the following velocity field for a cylindrical coordinate system with origin at 0: u = u(r, 6) = radial velocity v = v(r, 6) = tangential velocity' w = axial velocity = 0 The stress field may be represented as follows: 0Vr = 0 and K r c 0. It i s s t ab le a long t he z d i rect i o n b u t t h e r a d i a l i n s t a b i l i t y i s important. The inverse resu l t s a re obta ined for a face type coupling : Kz < 0 and Kr > 0.\nThe i n t e n s i t y of t h e s e s t i f f n e s s e s i s very import a n t . It i s about the same as t h e one produced by a magnetic bearing of same dimensions and constructed with t h e same magnetic material .\nA-s f o r t h e B type , it does not produce any s t i f f - ness [ 3 ] ; it means K r = KZ = 0. The two halves are in ind i f f e ren t equ i l ib r ium. P rac t i ca l ly , for small d is - placements around the centered position, the forces Fr and F, are very small .\nIn t h i s c a s e , K r and KZ may be simUltaneouSlY negative. The i d e n t i t y between t h e a1 and a2 types i s not conserved. The r a d i a l u n s t a b l e s t i f f n e s s K r becomes nore important, and I(, i s almost unaltered for t h e a2 type coupling ; b u t t h e most important effect of t h e yokes i s to inc rease the to rque r .\nFor the 6 t y p e s , t h e u t i l i z a t i o n of yokes does not conserve the indifferent equilibrium, and small in s t ab i l i t i e s appea r . I f t he equ i l ib r ium i s of f irst importance, the type must be used without soft iron\nThe a and 8' type need a c y l i n d r i c a l pa r t i t i on wa l l . -4 plane one i s necessary f o r t h e y type.\nEquations ( 1 ) and ( 2 ) are consequences of Earnshaw's theorem. It i s important t o n o t i c e t h a t they provide re la t ions about the s t i f fnesses for translational displacement only. This theorem makes no statement concerning the torque r . I11 - Study of a B'type magnetic coupling.\nTo v e r i f y t h e announced comportment of t h e 6 type couplings, a prototype has been built . The dimensions have been chosen so as to obta in very small a x i a l and r a d i a l f o r c e s i n a l a r g e a r e a around the centered pos i t ion .\nFigu - B' type coupling.\na. Experimental study. Figure 5 shows t h e 8-po1e coupling which has been realized. The magnet sec t ion dimensions a r e 9 mm x 9 mm, and the r ad ia l gap is 2,5 am wide. The exterior diameter of the inner r ing i s 118 mm. The two r ings of the outer par t are separated by 13 mm. The magnets a r e made with rubber-bonded f e r r i t e (B, = 0,22 T ) .\nFor t k i s coup l ing , t he maximum transmitted torque i s 0,26 Nm when 8 = 22,5 G. The main interest concerns i t s comportment for t ranslat ional displacements . We sha l l s tudy exper imenta l ly the ax ia l and r a d i a l f o r c e s r e s u l t i n g winen one of t h e two halves i s displaced from the centered posi t ion.\ntorqde angles." ] }, { "image_filename": "designv11_20_0001059_1.1839922-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001059_1.1839922-Figure1-1.png", "caption": "Fig. 1 Cross section view of test rig: rotor supported on three lobe gas bearings", "texts": [ " In addition, the bearing \u2018\u2018lift-off\u2019\u2019 characteristics and the identification of the threshold speed of instability and ensuing whirl frequency ratio demonstrate the stable performance of the bearing configuration tested. San Andre\u0301s 16 details the theoretical background for the analysis of hybrid gas bearing dynamic forced performance. Reference 17 is a companion paper to the work reported here and presents correlations between experimental and predicted rotordynamic behavior for the test bearings and rotor detailed below. Figure 1 shows a cross section of the test rig composed of a drive rotor, gas bearings and housings, air feed and discharge ports, and support instrumentation. The rotor weighs 0.827 kg 1.82 lb and consists of a solid steel shaft, 15 mm diameter and 190 mm long, onto which an integral brushless dc motor and two solid steel sleeves 28.48 mm 1.12 in. outer diameter are press fit. Eight 1-mm-diameter holes equally spaced at the rotor ends allow for the placement of imbalance masses.4 The integral motor maximum speed is 100,000 rpm and offers 0", " Figure 2 shows the end view of the test rig displaying two orthogonal positioned highly sensitive, eddy-current sensors that measure the displacement of the shaft at each end.5 The displacement sensors have a sensitivity of 39.4 mV/ m 1 V/mil with a linear range of 400 m. The sensor recorded voltages are conditioned to remove the large dc bias offset before connection to two separate oscilloscopes and/or the data acquisition system. The oscilloscopes display the rotor orbit at the end monitored. Force piezoelectric sensors mounted between the bearing housing and the test chamber alignment bolts measure the load transmitted through the bearings as depicted in Fig. 1 . The dynamic force sensors for the left bearing have a sensitivity of 118.8 mV/N 528.5 mV/lbf and a dynamic range of 44.48 N 10 lbf ; the sensors for the right bearing have a sensitivity of 12.0 mV/N 53.2 mV/lbf and dynamic range of 444.8 N 100 lbf . A high-speed infrared tachometer sensor is rigidly mounted inside the test chamber to indicate the shaft speed and to provide a key phasor signal for data acquisition. Thrust pins mounted rigidly to the test chamber prevent the axial movement of the rotor" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002140_1.5061038-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002140_1.5061038-Figure3-1.png", "caption": "Figure 3. Apparatus used for visualisation of the powder stream (schematic).", "texts": [ " The nozzle was positioned vertically so that the powder stream converged to a single stream in the laser focal plane, 7.5 mm below the nozzle, which is where the substrate was placed during deposition. Page 728 Laser Materials Processing ConferenceICALEO\u00ae 2007 Congress Proceedings The powder distribution beneath the nozzle due to conveyance of the Inconel 718 powder was first tested using an optical light sheet method and the results compared with the model. The experimental configuration used is shown schematically in Figure 3. A vertical light sheet from a masked halogen light was directed through the axis of the coaxial nozzle and a digital camera set to high resolution positioned at right angles to the light sheet. According to Mie theory, the average luminance of any scattering element within the stream is proportional to the concentration of powder within it [24] so the images from the camera could used directly to calculate the relative concentrations of reflective powder particles in the stream. The beam intensity distribution in the focal plane without any powder flow was next measured and the results compared with the model" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002397_2013.23130-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002397_2013.23130-Figure5-1.png", "caption": "Figure 5. Cutter bar diagram.", "texts": [ " The friction force (Fa) has ideally the same action line, while the counterbar reaction (R2) belongs to the vertical plane through this line and the X2 axis, in the middle of the blade (fig. 3). Figure 4 shows the shape of the cutting resistance at different values of the total filling percentage (p = p1\u00b7p2) and at different values of the average running speed (\u03c9). In agreement with other authors (Sitkei, 1986) and practice, it can be seen that R decreases with increasing average running speed. As shown in figure 5, analyzing the blade rotational equilibrium, it is possible to identify three couples arising from the vertical and horizontal offsets of the conrod small end from the cutting line: HFh\u2019\u2019h\u2019mh\u2019FC d\u2019\u2019mdFC dFC bTb Tb b .\u2212\u2212..\u2212.= ..\u2212.= .= 213 12 21 )( x .. 1 x .. 1 (15) where H = distance in the X1 direction between the middle of the blade and the conrod small end (m) h\u201d = distance in the X2 direction between the blade center of gravity and conrod small end (m) d = distance in the X3 direction between the cutting line and conrod small end (m) d\u201d = distance in the X3 direction between the cutting line and blade center of gravity (m) Cj = generic couple around the j axis, with j = 1, 2, 3 (N\u00b7m). The forces and couples from the blade (Fa, Fb2, C1, C2, C3), together with the external forces (R) and the forces from the crank pin (Fb1, Fb2), are supported by the counterbar carrying frame, which is fixed to the tractor chassis by means of a dap-joint. The translational equilibrium of the counterbar carrying frame (fig. 5) in the three directions is: =+ =++\u2212 =++++\u2212 0 0 0 33 2222 111 VF FVFR RFFVF i ib aib (16) 759Vol. 50(3): 755\u2212764 where Vj (N) is the generic dap-joint reaction force, and Fij (N) is the counterbar inertia force in the j direction, with j = 1, 2, 3. Using equations 7 and 8 and observing that there are no external forces in the X2 and X3 directions, these equations become: =+ =+ =+ 0 0 33 22 11 i i Ti FV FV mFV x .. 1 (17) Defining the barycentric axis as the Xj direction through the counterbar center of gravity, with j = 1, 2, 3, the rotational equilibrium around the barycentric X3 axis is: 0)( )()( )( 21221 22212 2333 =" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002567_robot.2007.364203-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002567_robot.2007.364203-Figure5-1.png", "caption": "Fig. 5. W \u2032 as seen on the surface of the sphere. The shaded region is the area need to be checked in addition to the other checking that can be done by the previous method.", "texts": [ " Also, The vectors on F2 that define the double tangents to F1 can be calculated by rotating n2 around r2 by \u03b82 and \u2212\u03b82 (see Fig. 4). C. R3-Positive Span of Four Force Cones This question is just an extension of the previous question. Let W = \u03a8(F1 \u222a F2 \u222a F3). First, let us consider the intersection of W with the unit sphere. Again, we extend each cone by the half-angle of F4, and check whether the axis of F4 lies inside W \u2032 = \u03a8(F \u2032 1 \u222a F \u2032 2 \u222a F \u2032 3) where F \u2032 i is defined in the same way as in Section IV-B. Fig. 5 illustrates W \u2032 as seen on the surface of the sphere. Obviously, a vector n4 is inside W \u2032 when either it is inside \u03a8(F1 \u222a F2), inside \u03a8(F2 \u222a F3), or inside \u03a8(F3 \u222a F1), or, finally, inside the pyramid defined by the axis of the three cones. The last area is illustrated as a shaded region in Fig. 5. The first three zones (\u03a8(Fi \u222a Fj)) can be checked directly by the method described in IV-B. The remaining problem is whether n4 lies strictly inside the pyramid of n1, . . . ,n3. Since we know that the three cones do not positively span, we know that n1, . . . ,n3 inevitably forms a pyramid. We have to calculate the inward normal vector of the three bounding facets of the pyramid. The normal vector of the three faces of the pyramid is an ordered pairwise cross product of n1, n2 and n3. However, we need to know the correct order of the axis" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002213_13506501jet575-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002213_13506501jet575-Figure2-1.png", "caption": "Fig. 2 Mapping of the flow field", "texts": [ " The constant parameters in the governing equations are given in Table 1. and ys is the normal distance of each nodal point to the nearest solid boundary that damps the turbulence kinetic energy. It should be noted that the above formulations are according to the AKN low-Re \u03ba\u2013\u03b5 turbulence model used in the present analysis [20]. The lubricant viscosity as a function of temperature is obtained from the relation \u00b5 = \u00b5i e\u2212\u03b2(T\u2212Ti) (15) In this expression, \u03b2 is the lubricant temperature\u2013 viscosity coefficient and Ti is the inlet lubricant temperature. According to Fig. 2, because of the complex flow geometry in the (x, y) plane, the governing equations are transformed into a simple computational domain, (\u03be , \u03b7), such that, the physical domain which is the region between two eccentric circles, is conformally mapped into a rectangular computational domain. The transformation between physical and computational planes can be performed in two steps. These transformations along with their transformation functions are shown in Fig. 2. For the bearing, a separate transformation function has to be used to map this region into the rectangle in the computational plane. Figure 3 shows this Proc. IMechE Vol. 223 Part J: J. Engineering Tribology JET575 \u00a9 IMechE 2009 at LAKEHEAD UNIV on March 12, 2015pij.sagepub.comDownloaded from transformation with its function. The detailed relationships of the transformation functions are given in the Appendix 2. From these transformation functions, the relations between physical and computational planes and thereby the transformed forms of the governing equations in the computational domain can be obtained", " For example, U and V , which are the \u03be- and \u03b7-velocity components in the computational domain, can be written as U = (y\u03b7u + y\u03be v) J (16) V = (\u2212y\u03be u + y\u03b7v) J (17) In these equations, J is the Jacobian of transformation, which is calculated from the relation J = y2 \u03be + y2 \u03b7 (18) such that x\u03be = y\u03b7 and x\u03b7 = \u2212y\u03be , based on the Cauchy\u2013 Reimann relations that hold for analytic functions. In order to obtain the numerical results, appropriate boundary conditions are used to solve the discretized forms of transformed equations. Referring to Fig. 2(a), the following conditions in the physical plane are considered. 1. For all dependent variables such as velocity, pressure, turbulence kinetic energy, turbulence dissipation rate, and temperature, a periodic boundary condition in the circumferential sense is imposed. 2. On the journal\u2013lubricant interface, where (x \u2212 e)2+ y2 = r2 s , based on no-slip condition, lubricant velocity components are calculated as u = \u2212 y rs (19) v = (x \u2212 e) rs (20) Experimental investigation into the thermal effects on journal bearings by Dowson et al", " Besides, because of the existence of a recirculated zone at maximum film thickness, turbulent kinetic energy has a complex variation with two maximum points as shown in Fig. 11(a). The numerical results show high values of turbulent kinetic energy in the recirculated region of lubricant flow. The following results are about an infinite length journal bearing in which the numerical procedure discussed in section 6 is employed. The journal bearing has an axial groove located on the line of centres at the section of maximum gap (Fig. 2(a)), whose parameters are given in Table 2. The main parameters in the operation of journal bearings are Re, clearance ratio, and eccentricity ratio. In the present study, the effects of these parameters on the THD characteristics of journal bearings are thoroughly explored. In Figs 12 to 14, the effects of Re, clearance ratio, and eccentricity ratio on lubricant pressure and also on bearing load are studied. Figure 12 shows the effect of Re on lubricant pressure around the shaft. It should be JET575 \u00a9 IMechE 2009 Proc", "comDownloaded from h convection heat transfer coefficient J Jacobian of transformation k lubricant thermal conductivity kB bush thermal conductivity n normal direction P pressure ri inner radius of the bush ro bush outer radius rs shaft radius Re Reynolds number T lubricant temperature Ta ambient temperature Ti inlet lubricant temperature TB bush temperature TR Taylor number (u, v) velocity components in the x- and y-directions V\u0304 linear velocity of the shaft W bearing load (x, y) coordinates in physical domain Z physical plane \u03b1 coefficient of thermal diffusivity \u03b2 temperature\u2013viscosity coefficient \u03b5 turbulence dissipation rate \u03b8 angle in direction of rotation \u03b8i half angle of groove span \u03ba turbulent kinetic energy \u00b5 viscosity \u00b5t turbulent viscosity (\u03be , \u03b7) coordinates in computational domain \u03c1 density \u03c6 dissipation \u03c9 shaft angular speed Subscripts a ambient B bush l liquid s shaft v vapor APPENDIX 2 Transformation of physical domain into computational plane is carried out by two steps. First, the transformation function \u03b6 = (ari \u2212 z)/(az \u2212 ri) maps the circle of radius ri into the unit circle |\u03b6 | = 1, while the circle of radius rs is mapped into a circle of radius Rs in the \u03b6 -plane (Fig. 2), where Rs = [ e \u2212 ari a(ea \u2212 ri) ]1/2 (35) and a = r2 i \u2212 e2 \u2212 r2 s 2eri \u2212 [( r2 i + e2 \u2212 r2 s 2eri )2 \u2212 1 ]1/2 (36) In the second step, the analytic function \u03bb = \u03be + i\u03b7 = \u2212i ln \u03b6 maps the region between two concentric circles in the \u03b6 -plane into a rectangle in the \u03bb-plane, with \u03be changing from 0 to 2\u03c0 and \u03b7 from 0 to \u2212ln(Rs). Combining the above transformations ( x ri ) fluid = e\u2212\u03b7(1 + a2) cos \u03be + a(1 + e\u22122\u03b7) 1 + a2e\u22122\u03b7 + 2ae\u2212\u03b7 cos \u03be (37) ( y ri ) fluid = e\u2212\u03b7(1 \u2212 a2) sin \u03be 1 + a2e\u22122\u03b7 + 2ae\u2212\u03b7 cos \u03be (38) from which the metric coefficients x\u03be , y\u03be , and so on in the flow field can be calculated" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001100_bf01228535-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001100_bf01228535-Figure6-1.png", "caption": "Fig. 6. First-order solution for ejection Uo=0.005 and ~bo=n/3.", "texts": [ " (36) 2 A NEW FIRST-ORDER SOLUTION FOR THE RELATIVE TRAJECTORIES OF A PROBE 85 The first-order solution in this curvilinear coordinate system is then = (N~/Np - 1)Ep - (Ns /Np + 1)ep sin Ep + ~ - ~*, (37) r / = (ap - 1) - apep cos Ep. (38) These Equations (37) and (38) can be put in the same parametric form as the equations that describe the prolate cycloid in a rectangular coordinate system (Lockwood, 1961). 8 6 T. F. BERREEN AND J. D. C. CRISP The trajectories are actually prolate cycloids in ~/[(Ns/Np + 1)/ap], r/space. It becomes: clear later that, to first-order in Xo, this space reduces to the simple ~/2, ~/space. The trajectory yielded by an ejection Uo--0.005 and ~bo-~r/3 is shown in Figure 6. As noted previously, the values of q are exact and the values of ~ are periodically exact, for Ep-kzr, where k is an integer. Points such as A are therefore exactly mapped by this approximate solution, there being no secular accumulation of error as with the solution to the first-order in displacement ratio, described previously. The maximum error in ~ is of the order 1 2 of-cep= 0.000 01 and occurs for Ep = ( 2 k - 1)zr/4 at points such as B in Figure 6. For the scale used in that figure, the approximate solution and the exact solution are indistinguishable. This new first-order solution can be reliably adopted for trajectories yielded by ever~ higher ejection speeds. Such a trajectory, namely that for Uo =0.1 and ~bo-0 where the maximum error in ~ is of the order 0.01, is illustrated in Figure 7. In this figure some exact points are also shown to emphasize that, to the scale used, this first-order solution closely approximates the exact solution, and is periodically exact at points A" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000507_2006-01-0582-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000507_2006-01-0582-Figure4-1.png", "caption": "Figure 4. Coordinate transformation from the coordinate of tri-axial force sensor to the housing coordinate", "texts": [ " In order to calculate the internal friction coefficient that is present under all CV joint conditions, one needs to find a universal equation to cover all articulation angles and rotational phase angles. This is accomplished by introducing a coordinate transformation matrix based on Euler angles representing articulation angle , and rotational phase angle . Using the defined coordinates of the tri-axial forces shown in Figure 3, one can obtain the combined rotational transformation matrix. By multiplying each individual transformation matrix, as depicted in Figure 4, in sequence, the following equation that relates the measured internal forces to the global forces in accordance with the housing coordinate is given by: Fz Fy Fx zF yF xF )90(sincos)90(cossin)90cos(]cos1[)90sin()90cos( sin)90cos(cossin)90sin( ]cos1[)90sin()90cos(sin)90sin()90(sincos)90(cos 22 22 (1) Where, the CV joint rotational phase angle is defined as shown Figure 5, from the housing end view. Thus, by using this relationship one can calculate the net friction coefficient along the housing groove by using the relationship as shown in the equation (2) below" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000844_iemdc.2005.195712-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000844_iemdc.2005.195712-Figure2-1.png", "caption": "Fig. 2 Open-circuit field distributions with 4-pole rotor at peak cogging torque position", "texts": [ "2T and a relative recoil permeability of 1.05. The cogging torque waveforms and harmonic spectra which result when the stator has only 1 slot and 2, 3, 4, 5 and 6 uniformly distributed slots are analysed, and analytically synthesised and finite element predicted results are compared with measurements. The Maxwell stress integration method is employed in the finite element calculation of the resultant cogging torque, with particular attention being paid to the discretisation so as to achieve the required accuracy. Fig. 2 shows open-circuit magnetic field distributions for the 1-slot and 6-slot motors, whilst Fig. 3 shows the cogging torque of the 1-slot motor, for which 2p=4, Ns=1, Nc=4 and C=1. Therefore, its cogging torque waveform has a periodicity of 90\u00b0 mechanical, and )4sin( ...3,2,1 1 \u03b8\u2211= \u221e = iTT sci i cog (12) where Tsci is amplitude of the ith harmonic the cogging torque. The least common multiple Nc, the goodness factor C, the harmonic order and the amplitude of the cogging torque which results in the 2-slot, 3-slot, 4-slot, 5-slot, and 6-slot, 4- pole motors are given in Table I" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000839_j.euromechsol.2004.08.003-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000839_j.euromechsol.2004.08.003-Figure1-1.png", "caption": "Fig. 1. A three-dimensional beam undergoing large overall motion.", "texts": [ " The Numerical example of a flexible spatial manipulator is used to verify the effectiveness of the criterion. In this section, the geometric nonlinear formulation of a three-dimensional beam undergoing large overall motion is established based on the following assumptions. The beam has homogeneous and isotropic material properties, the elastic and centroidal axes in the cross section of a beam coincide so that the effects due to eccentricity are not considered. The shear and torsion effects are ignored. A three-dimensional beam is shown in Fig. 1. Two coordinate systems are introduced to describe the motion of the beam: the global coordinate system O0\u2013X0Y0Z0 and the body-fixed coordinate system Ob\u2013XbYbZb . The position vector of point k on the central line of the beam r can be defined with respect to the O0\u2013X0Y0Z0 as r = r0 + \u03c10 + u \u21d2 0r = 0r0 + A(\u03c10 + u), (1) where A is the transformation matrix that defines the orientation of the body-fixed frame, 0r,0 r0 are the position vectors of k and Ob defined in O0\u2013X0Y0Z0, and \u03c10 is the position vector of k with respect to Ob\u2013XbYbZb in the undeformed state, and u is the deformation vector defined in Ob\u2013XbYbZb , which can be written as \u03c10 = xe1, e1 = [1 0 0]T, u = [u v w]T" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002964_10402000903283318-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002964_10402000903283318-Figure4-1.png", "caption": "Fig. 4\u2014Aerodynamic bearing mounted in the housing.", "texts": [ " The rotor is 350 mm long, and, following numerical estimations, the first natural frequency of its free-free beam mode is 1250 Hz. The rotor can then be considered rigid for speeds enabled by the spindle (< 60 krpm). However it is possible that the natural speeds of the hybrid Lomakin bearing system may interfere with measurements. A modal analysis of rotor\u2013hybrid Lomakin bearing system will be presented later in this article. The housing is overhung\u2013mounted at the nondrive end of the rotor in order to enable rapid dismount. The bearing shown in Fig. 4 is made of graphite contained in a stainless steel sleeve and is mounted in a housing with the aid of two conical parts. As shown in Fig. 1, the bearing is floating on the stator, the static and the dynamic forces being applied on its housing. However the bearing is not completely free because a squirrel cage made of three flexible stingers is provided between the housing and the base plate. The role of the low-stiffness squirrel cage is to reduce misalignment problems that appear during dynamic tests without interfering with measurements" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002558_acc.2009.5160468-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002558_acc.2009.5160468-Figure1-1.png", "caption": "Fig. 1. The two pairs of counter-rotating blades allow the aircraft to hover without rotating about the central point.", "texts": [ ". INTRODUCTION Quadrotor helicopters have been an increasingly popular research platform in recent years. Their simple design and relatively low cost make them attractive candidates for swarm operations, a field of ongoing research in the UAV community. Quadrotor helicopters typically consist of two pairs of counter-rotating blades mounted on a carbon fiber frame as shown in Figure 1. In designing a controller for these aircraft, there are several important considerations specific to this problem. There are numerous sources of uncertainties in the system\u2013actuator degradation, external disturbances, and potentially uncertain time delays in processing or communication. These problems are only amplified in the case of actuator failures, where the aircraft has lost some of its control effectiveness. Additionally, the dynamics of quadrotors are nonlinear and multivariate. There are several effects to which a potential controller must be robust: the aerodynamics of rotor blade (propeller and blade flapping), inertial antitorques (asymmetric angular speed of propellers), as well as gyroscopic effects (change in orientation of the quadrotor and the plane of the propeller)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002515_00423110701810596-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002515_00423110701810596-Figure3-1.png", "caption": "Figure 3. Membrane element.", "texts": [ " As regards the sidewalls, the structure assumed to describe the tyre behaviour causes a tensile stress state lying on the plan tangent to the surface in the considered point. This point can be determined (Figure 2) by coordinates (\u03b8 , r), in which r represents the parallel circumference radius and \u03b8 the anomaly of the considered meridian plan, having selected an arbitrary meridian plan as reference. A neighbourhood of this point in which the lateral faces are subjected to normal loads can only be isolated. Imposing the element equilibrium (Figure 3) along the direction normal to the surface and along the direction tangent to the meridian line, one obtains [15] T = 1 2 \u03c1pi r + r \u2032 r Tp = 1 2 \u03c1pi. (1) D ow nl oa de d by [ Fl or id a St at e U ni ve rs ity ] at 0 0: 16 0 8 O ct ob er 2 01 4 18 G. Capone et al. In Equation (1), T and Tp are the meridian and parallel tensions, respectively, while pi is the inflation pressure. The aforesaid relations can be obtained by solving the two equations system, one differential and the other one algebraical, constituting the local equilibrium conditions" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000235_bf02905937-Figure21-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000235_bf02905937-Figure21-1.png", "caption": "Figure 21. Wrinkled surface", "texts": [ " The appearance of a compression area on the surface of the membrane is considered as a physical nonlinearity combined with a significant displacement finite element. The numerical model presented in this paper and used by D\u2019Uston [17] is based on two hypotheses: 1. The membrane is not able to endure major negative stress. The appearance of a major negative stress \u03c32 p produces a pleat perpendicular to \u03c32 p. 2. An average fictitious surface replaces the distorted surface by eliminating the pleat, as shown on the Figure 21. The stress state in an element is determined by: \u2212\u2192\u03c3 = C\u2212\u2192\u03b5 +\u2212\u2192\u03c3 o in local coordinate system (48) \u2212\u2192\u03c3p = \u23a7\u23a8\u23a9 \u03c31 p \u03c32 p 0 \u23ab\u23ac\u23ad and \u2212\u2192\u03b5p = \u23a7\u23a8\u23a9 \u03b51p \u03b52p 0 \u23ab\u23ac\u23ad (49) with \u2212\u2192\u03c3p the major stress vector in the principal directions (\u2212\u2192e 1,\u2212\u2192e 2) as shown on Figure 21, where \u03c31 p \u2264 \u03c32 p and \u2212\u2192\u03b5p major strain vector in the principal directions (\u2212\u2192e 1,\u2212\u2192e 2). Two cases may occur requiring a local modification: 1. Uniaxial tensile state: \u03c32 p \u2264 0 and \u03c31 p \u2265 0. In this case, Equation 48 is reduced to that of uniaxial behaviour according to the first major direction: \u03c31 p = A\u03b51p + B, where A and B are expressed according to C and \u2212\u2192\u03c3 o p, based on the equation \u03c32 p = 0. The stiffness matrix of the element is then determined by using the new elastic stiffness matrix Cuni and stress vector \u2212\u2212\u2192\u03c3uni" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000669_s095679250400573x-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000669_s095679250400573x-Figure3-1.png", "caption": "Figure 3. (a): Case (2(i)) a = A cos(ah)/2. Solutions a are given by the intersection of the solid curves y = cos(ah) and y = 2(ah)/(Ah), and these are stable if the corresponding value of sin(ah) (the dotted curve) is positive. Thus, the first root is stable, the second unstable, the third stable, and so on, alternating between stable and unstable as a increases for as long as roots exist. (b): Case (2(ii)) a = \u2212A cos(ah)/2. Solutions a are given by the intersection of the solid curves y = \u2212 cos(ah) and y = 2(ah)/(Ah), and are stable if the corresponding value of sin(ah) (dotted) is negative. Thus, the first root is unstable, the second stable, the third unstable, and so on.", "texts": [ " For case (2(i)), a = A cos(ah)/2, if sin(ah) > 0 then both surface contributions are positive, giving a net positive 2nd variation, and a stable solution. If sin(ah) < 0 then both surface contributions to the 2nd variation (4.7) are negative, and thus the 2nd variation can be made negative for a suitable choice of \u03b7, so the solution is unstable. For case (2(ii)), a = \u2212A cos(ah)/2, the reverse holds: if sin(ah) < 0 then both surface contributions to the 2nd variation are positive, giving a net positive 2nd variation, and a stable solution, while if sin(ah) < 0 then both surface contributions are negative, so the solution is unstable. Figure 3 illustrates the situation. The solution observed in practice will be that stable solution of lowest energy, i.e. that with the lowest value of a, since ah represents the angle turned through by the director across the film. Thus, the observed solution will be the smallest stable root of a = A 2 cos ah (4.11) (case (2(i)), with b given by (4.9) (see Figure 3(a)); this solution always exists, whatever the value of A), which lies in the range 0 < a < \u03c0 2h . (4.12) The free elastic energy W is again given by W = \u03b82 z 2 = a2 2 . (4.13) We now return to the momentum equations to see how the analysis of \u00a7 3.2 is modified. Equations (3.17)\u2013(3.20) are unchanged; but the different solution for \u03b8 means that the right-hand sides of the stress conditions (3.21) and (3.22) are altered, giving the modified stress boundary conditions \u2202u \u2202z ( 2\u03b11 sin2 \u03b8 cos2 \u03b8 + (\u03b15 \u2212 \u03b12) cos2 \u03b8 + (\u03b13 + \u03b16) sin2 \u03b8 + 1 ) = \u2212Na(hax + bx) = \u2212Na 2 ( hax \u2212 ahx ) on z = h(x, y, t), (4", "26), cos(2b) = { sin(ah) case (2(i)), \u2212 sin(ah) case (2(ii)), and cos(2ah+ 2b) = { \u2212 sin(ah) case (2(i)), sin(ah) case (2(ii)), so stability depends on the sign of sin(ah). For case (2(i)), a = \u2212A cos(ah)/2, if sin(ah) > 0 then both surface contributions are negative and the solution is unstable, and if sin(ah) < 0 then both surface contributions are positive and the solution is stable. This is exactly equivalent to the case (2(ii)) considered in \u00a7 4.1 previously, and the situation is sketched in figure 3(b). (The director solution is not identical however, as the value of b differs in the two cases.) For case (2(ii)), a = A cos(ah)/2, if sin(ah) < 0 then both surface contributions are negative and the solution is unstable, and if sin(ah) > 0 then both surface contributions are positive and the solution is stable. This is exactly equivalent to the case (2(i)) considered in \u00a7 4.1 previously, and the situation is sketched in Figure 3(a) (again though, the solution for b differs). Hence the observed solution will be the smallest positive root of a = A 2 cos(ah), (4.29) with b given by (4.28). This director solution is exactly as in \u00a7 4.1 except for an additive term \u03c0/2. Thus the PDE governing the film height evolution follows almost exactly as before, with just minor changes to the definitions of the Ji: \u2202h \u2202t + \u2202 \u2202x [K1(C(\u22072h)x \u2212 Bhx \u2212 Naax) \u2212 K2Na(hax \u2212 ahx)] + \u2202 \u2202y [K3(C(\u22072h)y \u2212 Bhy \u2212 Naay) \u2212 K4Na(hay \u2212 ahy)] = 0, (4.30) where a is determined by (4" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001049_j.jelechem.2005.02.025-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001049_j.jelechem.2005.02.025-Figure1-1.png", "caption": "Fig. 1. Typical cyclic voltammograms of SQ, MeSQ and Me2SQ films prepared on AujGCE recorded in (1/15) M phosphate buffer (pH 5.0) in a nitrogen atmosphere. Temperature 20 C, scan rate 50 mV s 1: (\u2014) SQ film; (\u2014-\u2014) MeSQ film; (- - - -) Me2SQ film.", "texts": [ " Table 1 shows electropolymerization conditions of mercaptohydroquinone (H2QSH), MeH2QSH and Me2H2QSH and amounts of QU s in the resulting films. The QU values were calculated by measuring the areas of their cyclic voltammograms obtained at a scan rate of 2 mV s 1 in the phosphate buffer. The QU quantity of 10.2 nmol cm 2 corresponds to ca. 17 monolayers on the AujGCE surface, based on a close-packed monolayer of ca. 0.6 nmol cm 2 for 4-methyl-1,2-benzenediol lying flat on a perfectly flat carbon surface [31], assuming that 4-methyl-1,2-benzenediol is nearly equal to these mercaptohydroquinone derivatives in molecular size. Fig. 1 shows typical cyclic voltammograms of EESQ, EEMeSQ and EEMe2SQ prepared under the same conditions as in Table 1. Both the voltammetric responses remained qualitatively unchanged even after potential scan cycling of several hours in the phosphate buffer (pH 5.0) at 25 C. Cyclic voltammetric characteristics of these films are given in Table 2. The formal redox potentials (E s) of SQ, MeSQ and Me2SQ films which are estimated as the averages of anodic and cathodic peak potential M phosphate buffer. Other experimental conditions are the same as in Fig. 1. (\u2013n\u2013) Me2SQ film (\u2013d\u2013) MeSQ film and (\u2013s\u2013) SQ film. values of their respective cyclic voltammograms, gave linear relationship with a slope by ca. 60 mV/pH, as shown in Fig. 2. The E s of MeSQ and Me2SQ films were found to shift more negatively than that of SQ film by ca. 80 and ca. 150 mV, respectively, over the examined pH range. The amount of GOx immobilized on the electrode surface can be estimated by measuring the fluorescence intensity of flavin adenine dinucleotide (FAD) [32]. The amount of GOx immobilized on EESQ was 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001564_tie.2006.885468-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001564_tie.2006.885468-Figure3-1.png", "caption": "Fig. 3. Simulated virtual part of uniform road geometry.", "texts": [ " A desirable fabrication process is one with tolerable overfills and no underfills in each layer. Fig. 2 illustrates overfills and underfills in a typical LM application [10]. In this paper, we will focus on minimizing or reducing the defects due to deposition inaccuracy, which is caused by the variations in actual road width along a tool path. Assuming a tool path is well designed and positioning process is sufficiently accurate, a uniform-road width everywhere in a layer will guarantee desirable layer quality (see Fig. 3 for an illustration [13]). A common strategy for layer filling has been to apply contour fills for the boundaries and vector fills for the interiors. The most frequently used fill pattern for making a solid-dense part is the vector fill [see Fig. 5(b)]. Since the curvature of a vector fill is unbounded at the turn point, the reference head speed, originally set to a constant, will drop down unavoidably while turning [9]. Hence, for a narrow area of many short vector segments, the head speed will undergo frequent decelerations" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001670_s11706-007-0005-1-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001670_s11706-007-0005-1-Figure1-1.png", "caption": "Fig. 1 Schematic of the GMA welding process", "texts": [ "n the spray transfer mode [1,2], but there are few studies of weld pool dimensions in the globular transfer mode. In recent years, some researches have been carried out to study the droplets transferring from the tip of the electrode to the weld pool [3\u22125]; however, there is still a lack of quantitative analysis of weld pool geometry and dimensions in GMAW. In this study, the weld pool geometry and dimensions are numerically simulated, which provides a further understanding of the welding physics phenomena in globular transfer mode during the GMAW process. Figure 1 shows that a GMAW torch moves along the welding seam at a speed of u0. A consumable electrode wire is constantly supplied through the center of the welding torch. Electrical current, imposed on the electrode by a voltage drop between the contact tube and the metal to be welded (workpiece), generates an arc between the electrode and the workpiece. The electrode wire is melted by internal resistive power and heat transferred from the arc. The heat of the arc melts the workpiece and form a molten pool below the arc" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003906_978-3-319-00636-9-Figure2.9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003906_978-3-319-00636-9-Figure2.9-1.png", "caption": "Fig. 2.9 Illustrated explanation of the red- and blue-shift of aggregates. These cartoons illustrate how the arrangement of two planar identical molecules (e.g., porphyrins) to a dimer and thus the arrangement of their dipole moments (arrows) influences the excited state energy splitting and selection rules, which correspond to a red- or blue-shifted absorption spectrum. The thick horizontal arrows indicate a momentum view of the transition dipole moments. The thin solid vertical arrows indicate the allowed transitions from the ground states (HOMO) to the excited states, whereas broken lines correspond to forbidden transitions. The energy splitting (distance between both excited states of a dimer) is dependent to the angle between the molecules and can reach zero, as shown in Fig. 2.10. [33]", "texts": [ " Absorption spectra of isolated monomers within perfect solutions may also be shifted slightly depending on the polarity of the solvent molecules. Thereby the whole spectrum is shifted in contrast to the case of well defined aggregates where the shift is band-specific [32]. The shift of the absorption spectrum results from the splitting of the excited state (LUMO), caused by Coulomb interactions of adjacent molecules, and from the selection rule of the transition dipole moments, like shown in the illustration from Satake et al. in Fig. 2.9. The transition dipole moment (arrow) within a planar dye molecule, like a porphyrin, interacts with the transition dipole moment of an adjacent molecule due to Coulomb interactions, leading to a splitting of the excited energy state E (LUMO) of a monomer into a higher (E\u2019\u2019) and a lower (E\u2019) energetic state of the dimer. In the case of a pure face-to-face orientation of both molecules the parallel orientation of both transition dipole moments result in a higher energy and the antiparallel dipole moment orientation in a lower one" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001474_ac50010a015-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001474_ac50010a015-Figure4-1.png", "caption": "Figure 4. Schematic diagram of continuous evaporator concentrator", "texts": [], "surrounding_texts": [ "Design and Optimization of a Teflon Helix Continuous Liquid-Liquid Extraction Apparatus and Its Application for the Analysis of Organophosphate Pesticides in Water\nC. Wu\u2018 and 1. H. Suffet*\nEnvironmental Studies lnstitute, Department of Chemistry, Drexel University, Philadelphia, Pa. 19 104\nA statistical optimization study leading to the final design and operational parameters for extraction of organophosphate pesticides is described with a new \u201cTeflon Helix Contlnuous Liquid-Liquid Extractor\u201d. A coil length of 32 feet, helix winding diameter of 1 inch, aqueous flow rate of 900 ml/h, water to solvent ratlo of 1O:l without premixing were the optimal design parameters. Recoveries of I.cg to gg/i. concentrations of organophosphate pestlcides from fortified river water, sea water, and secondary sewage effluent were completed. Greater than 80 YO efficiency relative to batch extraction is demonstrated with contlnuous liquid-liquid extractlon and continuous evaporative concentration.\nThe unit operation of liquid-liquid extraction (LLE) has been used by many analysts for the isolation and concentration of organic pesticides from water with good results (I). Continuous liquid-liquid extraction (CLLE) is usually preferred over batch operation because of inherent continuous performance, labor savings, consistent sample handling, and the capability of sampling large volumes. However, many design and operational problems associated with the available continuous liquid-liquid extractors have limited the usefulness of this unit operation ( I , 2).\nA search for improvement in the design and the operation of the available continuous liquid-liquid extractors led to the design and construction of a new \u201cTeflon Helix Continuous Liquid-Liquid Extractor\u201d. The description and operational ,Jrocedures of this apparatus were reported previously (2). The mixing and extraction mechanisms of this new design are described in this paper. A statistical study leading to the final design and operational parameters for the extraction of organophosphate pesticides is also described. The analysis of pg to VgA. concentrations of organophosphate pesticides from fortified natural waters is demonstrated. Benzene is the solvent utilized as it appears to be an excellent solvent for extraction of organophosphate pesticides from aqueous solutions ( 3 , 4 ) .\nEXPERIMENTAL Apparatus and Procedure. A picture of the Teflon Helix Continuous Liquid-Liquid Extractor is shown in Figure 1. Its design is based upon a modular concept for greater analytical flexibility and is mounted on a 2 X 2.5 X 3 ft movable cart. Figures 2 to 4 show the major parts of the extractor. They are a Teflon coil mixer, a glass phase-separator, and a continuous evaporative concentrator, respectively. The design criteria as well as the operational procedures have previously been reported (2). Figure 5 shows a simplified schematic of the extractor. In short, the dual-channel micropump pumps the water and the solvent from the respective reservoirs to the Teflon helix mixer via a tee joint. The resulting water-solvent mixture from the Teflon coil enters the bottom of the phase-separator where the water and the solvent separate into two streams by gravity. The lighter solvent phase (benzene in this study) exits from the top of the sepa-\nl Present address, Gilbert/Commonwealth, Jackson, Mich. 49201.\n(1) Solvent reservoir, (2) Sample reservoir, (2\u2019) Sample reservoir and pre-filter (optional), (3) Micropump (4) Tee joint, (5) Premixer (optional), (6) Teflon helix mixer (7) Phase-separator, (8) Solvent phase exit, (9) Water phase exit, (10) Macroreticular resin bed (optional), (1 1) Solvent drying column (behind evaporator concentrator), (12) Evaporator concentrator, (13) Solvent reservoir recycle (optional), (14) Electrical power for parts 3, 5, and 12\nrator. The heavier water phase exits from the bottom of the separator. The solvent phase is passed through a Na2S04 drying column and then into the continuous evaporative concentrator. The apparatus has the capabilities of performing continuous liquid-liquid extraction, sample enrichment, and solvent recycling.\nANALYTICAL CHEMISTRY, VOL. 49, NO. 2, FEBRUARY 1977 231", "- - I\n1\nVENTING VALVE\nRAFFINATE OUTLET 1\nINTERFACE\nGLASS CONNECTOR\nWATER~SOLVENT MiXTURE INLET\nFlgure 3. Schematic diagram of glass phase separator-solvent lighter than water type\nAlthough the design of the Teflon Helix Continuous Liquid-Liquid Extractor is simple, the mixing and separation mechanisms that cause the water and the solvent to intimately mix and enable efficient extraction are very complex. This is due to interactions in the fluid acting simultaneously, Figure 5. Some components of the physical interactions believed to be responsible include pressure, inertia, gravity, viscosity, and dispersion. A quantitative description of these interactions was studied by a statistical design using an analysis of variance (anova) to analyze the experimental results.\nStatistical Optimization of Design Parameters. After the Teflon Helix Continuous Liquid-Liquid Extractor was designed, a set of preliminary statistical factorial experiments was made to optimize the design and operating parameters of the apparatus. The design parameters studied in a 25 statistical factorial experiment were (a) helix winding diameter, (b) coil length, (c) flow rate, (d) water to solvent ratio, and (e) the need for premixing. Table I lists the factors and levels studied in this experiment. The thirty-two runs of this preliminary statistical factorial experiment were made following a random order (5 , 6). Results were then used as the input data and analyzed by an IBM 370/168 Computer, using available APL statistical package ANOVA (7). Subsequently, a final statistical design (23) was completed in replicate to pinpoint relationships.\nAnalytical. Table I1 shows the group of 5 organophosphate pesticides studied. The aqueous concentrations are shown for the study of: a) the optimization of the continuous liquid-liquid extractor design (using distilled water) and b) recovery studies of natural water under different operational conditions. The gas chromatographic conditions,\n232 ANALYTICAL CHEMISTRY, VOL. 49, NO. 2, FEBRUARY 1977", "Super Centrifuge to remove the suspended solids. The water was pumped into the centrifuge by a Manostat Peristaltic pump at a flow rate of 125 ml/min while the centrifuge was operated at 40,000 rpm. The pH and the ionic strength of all water samples were adjusted to pH 4.2 and ionic strength 0.2 by adding KHzP04 except for sea water where only the pH was adjusted by adding concentrated H3P04. These conditions have been found t o minimize the effect of natural water character on the LLE process and maintain organophosphate stability (3 ) . The water samples after centrifugation, pH and ionic strength adjustment were stored in the refrigerator at approximately 3 \"C.\nAfter the optimum conditions of continuous liquid-liquid extractions were completed, water blanks, fortified distilled and natural waters were analyzed. Solvent recycling and extract enrichment were\nincluded to recover the pesticides. Final studies were completed on a total volume of 5 1. of water solution. A solvent extract of approximately 1 ml was obtained (concentration factor of 5000:l). The aqueous pesticide concentrations in this final study was from 25-66 vg/l., Table 11.\nRESULTS AND DISCUSSION Initial Factorial Experiment. Table IV is a three-dimensional presentation of t he results obtained from the 25 statistical factorial experiments. Parathion da ta are shown as a n example. The results are reported in terms of \"apparent E-value'' which is the fractional amount of a solute partitioned\nANALYTICAL CHEMISTRY, VOL. 49, NO. 2, FEBRUARY 1977 233" ] }, { "image_filename": "designv11_20_0003758_978-90-481-9262-5_31-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003758_978-90-481-9262-5_31-Figure4-1.png", "caption": "Fig. 4 Construction of a 16-PRRP-evolved radially reciprocating motion mechanism.", "texts": [ " Using a different method, Wohlhart [2] has developed the same mechanism and the geometry and kinematics of this mechanism have been studied by Wei and Dai [3]. Making use of the method presented in Section 3, two new overconstrained mechanisms are proposed below. 297 G. Wei and J.S. Dai In the type I 4-PRRP-combined chain in Fig. 2b, another prismatic joint is introduced in such a way that the axis of the prismatic joint is perpendicular to the base of the pentahedron and passes through central point O. Further introducing a square-shaped vertex P located at this new prismatic joint and rearranging the PRRP chains in a way as indicated in Fig. 4a, a further 4-PRRP-combined chain is formed. Since this 4-PRRP-combined chain contains a square-shaped vertex, it is a type II 4-PRRP-combined chain. Adding four additional vertexes V\u2032 1, V\u2032 2, V\u2032 3 and V\u2032 4 and corresponding links and removing the frame, a new mechanism is expected to be generated. However, looking at Fig. 4a, it is evident that removing the frame does not result in a mechanism because in this case links V1, V2, V3 and V4, and V\u2032 1, V\u2032 2, V\u2032 3 and V\u2032 4 are not connected. In order to form a mechanism, in addition to adding of vertexes V\u2032 1, V\u2032 2, V\u2032 3 and V\u2032 4, three square-shaped vertexes P1, P2 and P3 and corresponding links are further added to the type-II 4-PRRP-combined chain that a 16-PRRP-evolved overconstrained mechanism with mobility of one is constructed in Fig. 4b. The mechanism is a butterfly-shaped overconstrained mechanism with radially reciprocating mechanism in which when vertexes P, P1, P2 and P3 move away from central point O, vertexes V1, V2, V3, V4 move towards O and vertexes V\u2032 1, V\u2032 2, V\u2032 3 and V\u2032 4 away from O and vice versa. It should point out that in the mechanism, the angle \u03b1 between the two prismatic joints in a single PRRP chain is 54.74\u25e6 Considering the kinematic property of this mechanism, it can be used as a multi-functional grasping robot" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000936_978-1-4020-4941-5_24-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000936_978-1-4020-4941-5_24-Figure1-1.png", "caption": "Figure 1. 3-RPR planar parallel robot with congruent equilateral base and platform.", "texts": [ " In the next section, we will briefly present the analytic expression for the singularity loci of our special 3-RPR planar parallel robot. We will identify a range of orientations for which the robot has a sufficiently large singularity-free workspace. Then, in Section 3, we will describe the several types of singular configurations by studying the degeneracies of the direct kinematics and show that they belong to self motions. Conclusions are given in the last section. The special 3-RPR planar parallel robot is shown in Fig. 1. Its mobile platform and base form congruent equilateral triangles. We denote with Oi and Bi (in this paper, i = 1, 2, 3) the intersections of the base and platform revolute joint axes, respectively, with a plane normal to these axes. Then, let Oxy and Cx\u2032y\u2032 be the base and mobile reference frames, respectively. The generalized coordinates locating the mobile platform, i.e., the mobile frame Cx\u2032y\u2032, in the base frame Oxy will be denoted by x, y, and \u03c6. We define each active-joint variable \u03b8i as the angle between the x-axis and a unit vector vi that defines the direction of the prismatic joint of leg i, measured in counter-clockwise sense" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.118-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.118-1.png", "caption": "Fig. 14.118 Plastering unit for preparing and rendering mortars made from ready-made dry plaster mixes", "texts": [ " Plastering units are equipped with a remote control system for controlling the operation of the pump. If the spraying gun\u2019s air valve is closed, the pump\u2019s drive is automatically switched off and mortar feeding stops. The opening of the air valve results in the switching on of the pump\u2019s drive. As dry plaster mixes have become increasingly popular, plastering units for traditional mortars increasingly often feature screw pumps besides piston pumps. A plastering unit for feeding and rendering mortars from ready-made dry plaster mixes is shown in Fig. 14.118. After refitting the pump and changing the mortar feeding hose, the plastering unit can also be used for self-leveling floor compounds. A characteristic feature of the plastering unit shown in Fig. 14.118 is the reduced size of the mixer, whose function is performed by a mixing chamber with an agitator in the form of helical segments. The rate of delivery of plastering units for traditional mortars is usually up to 3 m3/h. For feeding traditional plaster mortars, hoses that are 52\u201358 mm in diameter with tip elements 32\u201336 mm in diameter are usually used. For feeding and spraying special media, pressure hoses with increased strength are used. Plastering units for traditional mortars can be adapted for spraying mixes to protect steel structures against fire, self-leveling mixtures, and similar materials" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002565_j.tws.2008.08.010-Figure14-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002565_j.tws.2008.08.010-Figure14-1.png", "caption": "Fig. 14. Fillets locatio", "texts": [ " The fillet radii B (the most important factor) has been measured on the two paddles using the same metrological camera. The measurement has been taken as shown in Fig. 13, representing the profile of a fillet. The portion of the curve handled between the two tangent lines represents the fillet, which radius is measured using a camera integrated software. The measures have been taken on the external part of the disc and some fillets radii (R1, R2, R5 and R6) had to be corrected by the angle y between the camera and bending angle because they were not measured perpendicularly as shown in Fig. 14. Results of the measures are given in Table 3. The thickness C has been also measured using the same camera, by observing the disc profile side. A value of 0.662 mm was found. The second disc, whose surface was discretized using a laser probe has been post treated by importing the point cloud in n on the paddles. Table 3 Measured fillets radii First cushion disc (measured by a metrological camera) (mm) Second cushion disc (measured by a laser probe) (mm) R1 39.2 30.7 R2 64 61.2 R3 44.8 47.4 R4 31 59 R5 55" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000319_095440605x8478-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000319_095440605x8478-Figure5-1.png", "caption": "Fig. 5 Example model", "texts": [ " Using the conventional equation for a compliant mechanism with long flexure hinges causes incorrect solutions, because the length of the flexure hinge influences the off-diagonal stiffness elements. The characteristic equation of equation (34) is jlI ~M 1 ~Kj \u00bc 0 (44) where l is the characteristic root and I is the identity matrix. Natural frequencies can be obtained by solving equation (44) as fi \u00bc 1 2p ffiffiffiffi li p (45) where fi and li are the ith modal frequency and the ith characteristic root respectively. Figure 5 shows a flexure-hinge-based compliant mechanism with three rigid bodies and six flexure hinges. Let the rigid body 0 be a fixed frame, the rigid body 1 be an outer annulus plate, and the rigid body 2 be an inner circular plate. Three outer Proc. IMechE Vol. 219 Part C: J. Mechanical Engineering Science C19203 # IMechE 2005 at NATIONAL UNIV SINGAPORE on June 27, 2015pic.sagepub.comDownloaded from flexure hinges are connected between the outer annulus plate and the fixed frame, and three inner flexure hinges are connected between the inner circular plate and the outer annulus plate" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000718_1.1829068-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000718_1.1829068-Figure5-1.png", "caption": "Fig. 5 Two-cam mechanism and its acceleration-equivalent four-bar", "texts": [ " The intended benefit of this was simplification of the kinematic analysis of the original mechanism. Transactions of the ASME shx?url=/data/journals/jmdedb/27802/ on 03/07/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F In the examples presented by Hain, links connected by revolute joints were replaced by different sets of links connected by revolute joints. Hall @10# demonstrated that this concept of equivalence applies to other joint types as well, such as cam joints. He showed, as an example, that a simple pair of cams like those in Fig. 5 can be expressed alternately as one of an infinite number of equivalent four-bar linkages. One refinement of his discussion is the discernment between velocity-equivalent mechanisms ~as shown by Hain @9#! and acceleration-equivalent mechanisms, for which both velocity and acceleration correspond to those of the original mechanism. Hall showed that for the simple two-cam example, there exist an infinity of acceleration-equivalent four-bar solutions and \u2018\u2018infinity squared\u2019\u2019 velocity-equivalent solutions. Whereas Hall\u2019s example shown in Fig. 5 illustrates an equivalence relation between a cam mechanism and a four-bar mechanism, in this paper the same kind of relation will be illustrated for nutating mechanisms and epicyclic bevel-gear trains. In general, these equivalent mechanisms are only equivalent to their original mechanisms with respect to instantaneous motion. Their shared kinematic characteristics are only common for one snapshot of the overall motion of the mechanism. For gear trains ~mechanisms containing only gear and revolute joints" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002868_j.mee.2009.06.033-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002868_j.mee.2009.06.033-Figure1-1.png", "caption": "Fig. 1. Surface-tension-driven self-alignment of a part. The sum of the interfacial surface tensions is minimised.", "texts": [ " The precisely formed adhesive film pads in [4\u20137] were accomplished by hydrophobic and hydrophilic interactions on micro-structured surfaces. The principle force for surface- tension-driven self-assembly is based on this type of partial wetting. As soon as an adhesive film of an alignment structure establishes a physical contact with a complementary binding site on a part forces appear that cause the sum of the interfacial surface tensions between the adhesive and the surrounding medium to be minimised (Fig. 1). Modelling of this self-assembly approach has been performed in the software Surface Evolver. For the implemented model the pads were assumed to be completely wetted by the fluid. The following parameters were systematically varied in the presented simulations (Fig. 1/Table 1): d: one-dimensional horizontal misalignment of the sites, w: the width and length of the square-shaped sites and h: gap height between the sites on the part and on the substrate. This concept has hardly been investigated. Several complementary, conductive micro-structured patterns serve as binding sites. These are marked with P on the substrate and with P0 on the part in Fig. 2. Whereas the binding sites on the substrate form two separated electrically conductive areas, the binding sites on the part form one single continuous area" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002643_978-0-387-25842-3_6-Figure6.1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002643_978-0-387-25842-3_6-Figure6.1-1.png", "caption": "FIGURE 6.1. Schematic representation of (A) single-/multi-walled carbon nanotube formation by rolling up graphene sheet(s). (B) carbon nanotube formation based on a 2D graphene sheet of lattice vectors a1 and a2, the roll-up chiral vector Ch = na1 + ma2, and the chiral angle \u03b8 between Ch and a1. When the graphene sheet is rolled up to form the cylindrical part of the nanotube, the chiral vector forms the circumference of nanotube\u2019s circular cross-section with its ends meeting each other. The chiral vector (n, m) defines the tube helicity. (C) Schematic representation of SWCNTs. (a) (5,5) armchair nanotube; (b) (9,0) Zigzag nanotube; (c) (10,5) chiral nanotube.", "texts": [ " In this chapter, we present an overview of the recent progress in carbon nanotube functionalization and electrode fabrication for biosensing applications by discussing some selected examples from the research and development carried out by many research groups, including our own work. As can be seen in Figures 6.1A&B, carbon nanotubes may be viewed as a graphite sheet that is rolled up into a nanoscale tube form (single-walled carbon nanotubes, SWCNTs) or with additional graphene tubes around the core of a SWCNT (multi-walled carbon nanotubes, MWCNTs) [1, 29, 30]. These elongated nanotubes consist of carbon hexagons arranged in a concentric manner, with both ends of the tubes normally capped by fullerenelike structures containing pentagons (Figure 6.1C). They usually have a diameter ranging from a few \u25e6 Angstroms to tens of nanometers and a length of up to centimeters. Because the graphene sheet can be rolled up with varying degrees of twist along its length, carbon nanotubes have a variety of chiral structures [1, 30]. Depending on their diameter and helicity of the arrangement of graphitic rings in the walls, carbon nanotubes can exhibit interesting electronic properties attractive for potential applications in areas as diverse as sensors, actuators, molecular transistors, electron emitters, and conductive fillers for polymers [1, 2]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000583_polb.20298-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000583_polb.20298-Figure1-1.png", "caption": "Figure 1. Schematic illustrations of (a) interdigitated line array on Al-Si-SiO2 substrate and (b) POMA\u2013FET structure. The (c) axes coordinate system used in the theoretical model.", "texts": [ " Aluminum was deposited by a conventional sputtering process, whereas the SiO2 layer was obtained by thermal oxidation of a 100 p-type Si wafer, and the gold line array (25 pairs) by the lift-off lithographic technique. In such a device, the line array acts as a source or drain contact, the underlying SiO2 as an insulating layer, and Al-Si as the gate. The length (l) and thickness (a) of the arrays were 800 and 0.1 m, respectively. Three different devices were fabricated with interdigitated array distance (W) equal to 10, 20, or 30 m. The thickness of POMA and Al films was equal to 0.1 m. Figure 1(a) shows a schematic interdigitated line array of the device without the final POMA deposition; Figure 1(b), the final deposition. The devices were dipped in different aqueous HCl solutions (10 3 and 10 2 M) to investigate the effect of the chemical doping level in the electrical behavior of the transistor. Undoped POMA was considered as weakly doped material because the polymer is not completely dedoped during the deprotonation step after the chemical synthesis.3 Source\u2013drain current (ISD) versus source\u2013drain voltage (VSD) curves were measured using a HP4145B Data Parameter Analyzer, and the electrical drain, source, and gate contacts were made using a Wentwoth Model 900 micromanipulator", "10 On the other hand, the unusual Is versus VG exponential dependence may be explained by the fact that the structure of POMA is more disordered at the polymer\u2013air interface in which the modifications of the chemical structure of the polymer are induced upon interaction with oxygen and/or water.11,12 This interfacial structure plays an important role in decreasing the carrier mobility. To explain the experimental curves, mainly those presented in Figure 3(a), we developed a model that is summarized by the following steps: 1. the coordinates x, y, and z, displayed in Figure 1(c), represent the source\u2013drain distance, the distance from the SiO2/ POMA interface, and the length of the electrodes (source and drain), respectively. dISD vldy (1) where is the carrier density, v is the carrier drift velocity, and ldy is the differential area in the polymer bulk (related to the zy plane). 2. As mentioned above, it is accepted that the POMA film has a less disordered structure near the SiO2/POMA interface (y 0). This may induce a gradient in the carrier mobility, in the y direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000849_sensor.1995.717235-FigureI-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000849_sensor.1995.717235-FigureI-1.png", "caption": "Fig. I ISFET glucose sensor", "texts": [], "surrounding_texts": [ "A non-invasive blood glucose monitoring system has been developed by integrating an Ion Sensitive FET (ISFET) based biosnsor technology and a suction effusion fluid (SEF) colllection technology. The system consists of an SEF measurement part, where the ISFET glucose sensor is mounted, and a main part which consists of a small vacuum pump, a liquid pump and a controller. A quasicontinuous measurement of blood glucose Concentration was achieved by the system." ] }, { "image_filename": "designv11_20_0003304_bio.1272-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003304_bio.1272-Figure1-1.png", "caption": "Figure 1. Schematic depiction of the proposed SIA\u2013MSFIA-CL set-up. MS, multisyringe module; S1\u2013S4, syringes (S1 = 10 mL; S2 = S3 = 5 mL; S4 = 1 mL); V1\u2013V4, three-way solenoid valves; T1 and T2, confluence points; SV, multiposition selection valve; HC, holding coil (2.5 m); RC, reaction coil (100 cm); L1 and L2, connecting tubings (20 and 4 cm); PSM, photosensor module; C, carrier (100 mM Tris\u2013HCl buffer, pH 8.5; R1, 1.0 \u00a5 10-4 mol/L luminol solution; R2, 10 U/mL peroxidase enzyme solution; S, sample (phenol derivative); R3, 2 \u00a5 10-3 mol/L hydrogen peroxide solution; W, waste. The exchange options of the commutation valves were classified in On/Off lines. The Off line was assigned to the solution flasks, and the On line was reserved for the flow network. The arrows indicate the flow direction.", "texts": [ "0 \u00a5 10-4 mol/L luminol working solution was daily obtained from the stock solution by dilution with the buffer. The hydrogen peroxide solution (2 \u00a5 10-3 mol/L) was daily prepared by diluting 30% v/v H2O2 (Panreac) in buffer solution. The standard phenol derivatives were daily prepared in buffer solution or in a 25% dimethyl sulphoxide aqueous solution for those compounds with low solubility in water. All the solutions were protected from light exposure throughout use. The proposed SIA\u2013MSFIA manifold (Fig. 1) comprised a multisyringe piston pump module (MS) from Crison (Barcelona, Spain), used as a liquid driver for performing the MSFIA\u2013CL operations. It was equipped with four syringes (Hamilton, Bonaduz, Switzerland), labelled S1\u2013S4. Each syringe was connected to a three-way commutation valve (V1\u2013V4; N-Research, Caldwell, NJ, USA) that allowed access to two different channels: solutions flask (Off), allowing its recycling, or flowing system (On). An eight-port multiposition selection valve (SV; Crison) was used as a sampling device to load the sample/standard and H2O2 into the holding coil (HC) in order to avoid carry-over into S2", " The main purpose was establishing conditions that enabled the maximization of the analytical signal corresponding to the blank (absence of phenol derivative) to a certain value, as well as the CL intensity difference between the blank and the phenol derivative, since the proposed procedure could involve CL enhancement or inhibition. Physical parameters: sample volume, flow rate and stopped-flow period in the reaction coil The manifold based on SIA\u2013MSFIA concepts was devised to accommodate the reaction between the phenolic compound, hydrogen peroxide and peroxidase prior to the addition of the chemiluminogenic reagent and subsequent direction toward the CL detector (Fig. 1). The CL measurements for the peroxidasecatalysed luminol oxidation scheme needed to be carried out at extremely high flow rates to enable proper light collection by the photosensing module (PSM), which was placed as close as possible to the confluence point T2, as shown in Fig. 2A. Total flow rates applied during the detection step (Table 1, step 15) from 3.75 to 45 mL/min were investigated and a total dispensing flow rate of 45 mL/min (the maximum rate available with the set of syringes assembled by activating simultaneously syringe/valve S1/V1 and S3/V3) was then selected for further experiments" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001544_iros.2007.4399373-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001544_iros.2007.4399373-Figure4-1.png", "caption": "Fig. 4. In our experiments, both the center of mass of the last link ln and the load lo are constrained to lie on the y axis of the last link\u2019s reference frame, so that the center of mass of their union l\u0301n also lies on the y axis.", "texts": [ " However, for control purposes, the inverse model of the robot will also have to be learned and another regressor matrix will have to be estimated. The six models learned in the previous section were used to obtain the regressor matrix and for performing load estimation and control under changing loads. Varying loads were randomly chosen with the constraint that out of the ten inertial parameters of the last link / load, only five were not zero. This was achieved by constraining both the center of mass of the link and the load to lie on the y axis of the link\u2019s reference frame (see Fig. 4) such that mnlnx and mnlnz are zero. Furthermore, the off-diagonal elements of the inertia tensor are zero. Out of the five non-zero inertial parameters, three were identifiable and inferred: mn (mass), mnlny (product of mass and the y-position of center of mass) and Inxx (moment of inertia around the x axis). We ran the simulation for 50 different trials (with random loads) and tried to estimate the inertial parameters of the compound last link / load. At the same time, the estimates were used for applying control as in eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002374_12.817405-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002374_12.817405-Figure8-1.png", "caption": "Fig. 8 \u2013 Distribution of Von Mises stress, the upper limit of the scale is 790 MPa, which is the yield stress of Ti-6Al-4V at room temperature. The deformations are multiplied by a factor of 10.", "texts": [ " The final part presents a mid-section where the hardness and the Young\u2019s modulus are lower than in the upper and lower regions, due to the presence of retained \u03b2, which is softer and less stiff than \u03b1\u2019. Proc. of SPIE Vol. 7131 713120-6 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 10/23/2015 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx Proc. of SPIE Vol. 7131 713120-7 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 10/23/2015 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx ,7 (MP) .......... .......... .......... .......... .......... The residual stress distribution in the final part is shown in Fig. 8. The residual stress in the part does not exceed 200 MPa, much lower than the yield stress of Ti-6Al-4V at room temperature, 790MPa. The calculations show that deformation occurs predominantly at high temperature, and the plastic behavior of the material at these temperatures generates low residual stresses in the part. Heating up of the workpiece during buildup is one of the main reasons for the non-uniform microstructure and properties distribution in the part. Heating of the substrate, in turn, is caused by the idle time between the deposition of consecutive layers being too short to allow sufficient cooling of the workpiece before the deposition of a new layer" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002239_j.euromechsol.2008.06.008-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002239_j.euromechsol.2008.06.008-Figure9-1.png", "caption": "Fig. 9. Planar 2-link WMM.", "texts": [ " 7(d) gives an example showing a trial path and the final path obtained after convergence of the process. Although initially a trial path might present undesirable distortions, they will be attenuated progressively as the process converges. We now give some numerical results obtained on a 1.6 GHz Intel Centrino Duo computer. The first example compares the pro- posed RPA to another method based on PMP, using a 1-link WMM (see Table 1 and Fig. 8). Most of the other examples will concern a planar 2-link WMM (see Table 2 and Fig. 9). The last example will deal with a spatial 3-link WMM. Unless otherwise specified, we have made use of a clamped cubic-spline model with Nm = 3 free nodes for the motion function \u03bb(\u03be) and a fourth-order B-spline model, respectively, with N p = 7 free nodes for the platform path P p(\u03bb) and Na = 5 free nodes for the manipulator path P a(\u03bb). Parameters of the simulatedannealing algorithm have been set as given in Haddad et al. (2007a). rw is radius of the wheel. L1 is the length of the link. Lw is the distance between the middle of the wheel axis and the center of the wheel", " It has required a runtime of 22 minutes. We note that the arm manipulator is generally oriented towards the center turn of the path; this is in order to compensate, by the weight of the arm manipulator, the influence of the centrifugal force on the dynamic stability of the robot. We consider the minimum-time trajectory problem with constraints on maximum torque and on stability for the spatial 3-link WMM (Fig. 21) whose characteristics are listed in Table 3. The plat- form is the same as that of the planar WMM shown in Fig. 9. The only difference concerns the maximum torques (which now are fixed as follows: \u03c4max p1 = \u03c4max p2 = 20 N m). All inertia moments are given with respect to the center of mass of the corresponding bodies. The moving frames attached to the rigid bodies of the manipulator are defined using the Khalil\u2013Kleinfinger method (Khalil and Dombre, 2002). The workspace is a (6 m \u00d7 6 m) flat floor with one obstacle (Fig. 22(a)). Boundary conditions are defined as follows: \u03a9START = (1,0.5,90\u25e6,90\u25e6,\u221245\u25e6,45\u25e6) and \u03a9GOAL = (3" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002984_iros.2010.5650918-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002984_iros.2010.5650918-Figure2-1.png", "caption": "Fig. 2. Geometric scheme of SUSM", "texts": [ " depicts the principle of SUSM and the classification of the torque control strategy of SUSM. Section III describes three kinds of position control methods based on the torque control strategy. Section IV demonstrates the experimental system and the experimental results. Finally, Section V provides the conclusions and future works. The SUSM used in this study consists of one spherical rotor and three ring-shaped stators. Figure 1 shows an overview of the SUSM. The geometric schemes are illustrated in Fig. 2. The stator includes a metallic elastic body and piezoelectric elements. When an AC voltage is applied to the piezoelectric vibrator, a standing wave is generated on the elastic body. By applying two AC voltages with a phase difference to the positive and negative sections of the piezoelectric elements, a traveling wave is generated due to combination 978-1-4244-6676-4/10/$25.00 \u00a92010 IEEE 3061 of the two standing waves[7]. The stators and the rotor are in pressure contact with each other, and the rotor is driven by the tangential force of the elliptical motion of the traveling wave. A single stator is shown in Fig. 3. Another piezoelectric element on the stator is used as a sensor detecting the resonance, and the signal is called the feedback signal. There are two inputs (AC voltage A and B), one output (Feedback) and FG (Frame Ground) terminals. The stators, namely vibrators, are located as shown in Fig. 2. Geometric parameters (stators\u2019 alignment) are \u03b81, \u03b82, \u03b83 and \u03c6. Using the parameters, the moment vector of each stator, mi, can be expressed as follows: mi = \u2212 cos \u03b8i cos\u03c6 \u2212 sin \u03b8i cos\u03c6 sin\u03c6 \u03c4i, (i = 1, 2, 3) (1) Here, \u03c4i is the generated torque of each stator. As a result, the output moment vector of the rotor, mrotor, can be described as the summation of the vectors mi in Eq. (1). mrotor = mx my mz = m1 + m2 + m3 = \u2212c\u03b81c\u03c6 \u2212c\u03b82c\u03c6 \u2212c\u03b83c\u03c6 \u2212s\u03b81c\u03c6 \u2212s\u03b82c\u03c6 \u2212s\u03b83c\u03c6 s\u03c6 s\u03c6 s\u03c6 \u03c41 \u03c42 \u03c43 = D\u03c4 ", " In the posture control strategy, we rotate the SUSM to the target position along the moment md. From the Eq. (2), the required stator torque \u03c4 can be obtained as follows: \u03c4 = D\u22121md . (3) Since the geometric parameter \u03c6 is nearly zero in the used SUSM, we premeditate such the 2DOF motion so that mz can be neglected, and we obtain the following relationship. \u03c4 = 1 d c\u03c6 s\u03b82 \u2212 s\u03b83 \u2212c\u03b82 + c\u03b83 s\u03b83 \u2212 s\u03b81 \u2212c\u03b83 + c\u03b81 s\u03b81 \u2212 s\u03b82 \u2212c\u03b81 + c\u03b82 [ mx my ] (4) where, d = sin(\u03b81 \u2212 \u03b82) + sin(\u03b82 \u2212 \u03b83) + sin(\u03b83 \u2212 \u03b81). As shown in Fig. 2(a), geometry of the each stator becomes \u03b81 = \u03c0/2, \u03b82 = \u2212\u03c0/6 and \u03b83 = \u22125\u03c0/6, and the Eq. (4) can be expressed as follows: \u03c4 = 2 3 \u221a 3 c\u03c6 0 \u2212 \u221a 3 \u22123/2 \u221a 3/2 3/2 \u221a 3/2 [ mx my ] (5) = \u2212 2 3 c\u03c6 cos ( \u03c0 2 ) sin ( \u03c0 2 ) cos ( \u2212\u03c0 6 ) sin ( \u2212\u03c0 6 ) cos ( \u22125\u03c0 6 ) sin ( \u22125\u03c0 6 ) [ mx my ] (6) = 2 \u2016md\u2016 3 c\u03c6 sin (\u03c8\u2032 \u2212 \u03b81) sin (\u03c8\u2032 \u2212 \u03b82) sin (\u03c8\u2032 \u2212 \u03b83) , (7) where, \u03c8 = atan2(my, mx) is the direction of the target moment md on the X-Y plane, and \u03c8\u2032 = \u03c8 \u2212 \u03c0/2 is the direction of the motion of lever as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001378_ijmmm.2007.015474-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001378_ijmmm.2007.015474-Figure6-1.png", "caption": "Figure 6 Simplified force diagram", "texts": [ "25 cos sin Pe cos sin cos sin n n n n n n FD B b B B \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 + = \u23a1 \u23a4+ \u2212 \u2212\u23a3 \u23a6 (17) \u2022 The maximum contact temperature at the tool-workpiece interface, TN, is supposed to be reached at the middle of tool-workpiece contact (see point N in Figure 5) and is given by the following equation: 0.25 eq 1.25 0.36sin 0.5 0.53 Pe n N A NT T B E \u03b1 \u03c8 \u239b \u239e = + +\u239c \u239f \u239d \u23a0 (18) where ( )( ) \u03b1 \u03c8 \u03b1 + = 1.25 eq 1 0.25 eq 0.6 1 Pe cos sin erf Pe / 4 n N n p B E B (19) and 0.3 0.1 eq 1 0.2 0.1 0.24 sin Pe nFD p E B \u03b1 = (20) Figure 6 shows simplified diagram of the forces generated in orthogonal cutting, as proposed by Astakhov. In this diagram R is the resultant cutting force, F\u03c4 the force along the plane of the maximum combined stress, FC and Ft are the power and trust components of the cutting force, respectively and Fn\u03b3eFt\u03b3 are the normal and the friction forces on the tool rake face. Because a tri-axial state of stress is the case in the deformation zone, even in simple orthogonal cutting, a good estimation for the power component of the cutting force can be given by the following equation (Isakov, 1997a,b): c c uF A \u03c3= (21) where Ac is the uncut chip cross-sectional area (m2) (calculated as described in the Appendix of Astakhov (1998)) and \u03c3u is the ultimate compression strength of the work material (Pa), function of temperature and strain rate. As follows from Figure 6, the resultant cutting force R then calculates as ( )cos ( / 4) cos(( / 4) arctan( )) c cF F R B\u03c0 \u03c6 \u03c0 = = \u2212 \u2212 (22) where B is a similarity number, calculated using Equation (13). Thus, the trust components of the cutting force, Ft, can be calculated as tan arctan( ) 4t cF F B \u03c0\u23a1 \u23a4= \u2212\u23a2 \u23a5\u23a3 \u23a6 (23) The experiments were carried out on a numerically controlled lathe using a bar turning process. Round bars of 140 mm diameter made of AISI 316L and AISI 1045 steels were selected for this study. The tests were conducted using both uncoated and coated tungsten carbide inserts" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003468_s12283-011-0074-3-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003468_s12283-011-0074-3-Figure1-1.png", "caption": "Fig. 1 Free-body diagram for a football moving through air. The figure shows a ball kicked with backspin, leading to a lift force with a component pointing up. If the ball\u2019s rotation axis were not pointing perpendicular to the page, a sideways component of the Magnus force would point perpendicular to the page", "texts": [ " This study used a 32-panel football manufactured by Umbro, model X III 250. With the ball\u2019s mass m = 0.424 kg, the gravitational pull on the ball from the Earth, i.e. the ball\u2019s weight, is mg = 4.155 N, where g = 9.80 m/s2, the magnitude of gravitational acceleration. With a constant gravitational force on the ball, the more interesting force on the ball is the one from the air. The air exerts a single force on the ball, though convention dictates that force is separated into three components because of the ball moving through, in general, three spatial dimensions. Figure 1 shows a free-body diagram of a football moving through air. One component of the air force is called \u2018\u2018drag,\u2019\u2019 and it points opposite the ball\u2019s velocity. The magnitude of the drag force [12] is FD = q A CD v2, where q is the air density (1.2 kg/m3), A is the cross-sectional area of the football (0.0375 m2), CD is the dimensionless drag coefficient, and v is the ball\u2019s centre-ofmass speed with respect to the still air far from the ball. The drag coefficient depends on the ball\u2019s speed and its rate of spin", "58 to 19.63 m/s, a range of speeds that Fig. 4 A football moves at an average speed of approximately 19 m/s through a dust cloud. The distance between the centre-of-mass points for the first circle and the last circle is 47.6 cm. Adjacent circles are separated in time by 0.0025 s Fig. 5 Schematic of how the boundary-layer separation angle u is defined. Image by Tracy Chase contains the drag crisis (see Fig. 2). To help put the dust tests into context, also shown in Fig. 8 is the fitted drag curve from Fig. 1. A total of 16 tests are labelled in Table 1. There are six clusters of tests to be discussed: 1\u20135, 6 and 7, 8 and 9, 10 and 11, 12\u201314, and 15 and 16. The low-speed cluster of tests (1\u20135) has separation angles in the range 179 B u B 206 . A close inspection of the \u2018\u2018Top\u2019\u2019 and \u2018\u2018Bottom\u2019\u2019 images in Table 1 reveals that the boundary layer separated on smooth patches on the top and bottom of the ball in test 1. For tests 2\u20134, the boundary layer separates on a smooth patch on one side of the ball and near seams on the other side" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002945_amr.97-101.3366-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002945_amr.97-101.3366-Figure3-1.png", "caption": "Fig. 3 Temperature distribution of (a) inner ring, (b) outer ring", "texts": [ " (4) where f1 is the friction coefficient, p (N) is the load, and v (m/s) represents the mean sliding velocity, dk is the sphere diameter of the inner ring, \u03b2 (\u00b0) represents half angle of oscillation degrees, and f2 (r/min) is the frequency of oscillation. Structural model. A 3D structural stress analysis model was used with exactly the same geometry as the thermal model (Fig. 2). The temperature distribution from the thermal analysis was applied as body loads. The mesh methods, element size, time step size were all the same as the thermal model in order to keep good compatibility. Fig. 3 shows the steady-state temperature distribution of the inner ring and outer ring. The maximum temperature, 78.1\u00b0C, occurs on the outer surface of the inner ring and the inner surface of the outer ring. The maximum temperature of the side surface on the outer ring is 66.8\u00b0C. Through comparison with the experimental results [10] under different room temperature, good agreement is found for the maximum temperature of the side surface on the outer ring (Table 1). The experimental results are larger than the FEM results, because the material properties and environmental conditions may change with temperature rise" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002821_j.sna.2010.07.002-Figure12-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002821_j.sna.2010.07.002-Figure12-1.png", "caption": "Fig. 12. Cross-sections of the RIUSM, th", "texts": [ " 10 shows the olarization pattern of piezoelectric ceramic. The vibration mode f the stator is B (3, 1). The stator is made of yellow brass. The onstruction of the stator is shown in Fig. 11. A PCB circuit is bonded on the steel backing and piezoelectric ing is bonded on the PCB circuit, by epoxy. Electrode patterns n PCB conduct the electricity to the piezoelectric ring segments. ibrating brass ring (stator) is bonded on the piezoelectric ring by ltrasonic epoxy (UHU plus epoxy) which is strong and high fatigue ife cycle epoxy. Fig. 12 shows the cross-section of the RIUSM. The mechanical characteristics of the motor with various freuencies and amplitude of the applied voltage have been calculated sing the proposed analytical model. The unknown boundaries of he contact region are founded as follows: = 0.065 mm, L\u2032 = 1.099 mm Using parameters shown in Table 1, mechanical characteristic of he motor is plotted by MATLAB software. The experimental value f speed is measured by optical tachometer. Since the torque of IUSM is very small, the experimental value of torque is measured sing equivalent circuit of ultrasonic motor [14]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000718_1.1829068-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000718_1.1829068-Figure10-1.png", "caption": "Fig. 10 Sun-ring-planet spur-gear train and its kinematically similar bevel-gear train \u201eS1 \u2026", "texts": [ " Thus the correct declaration is that these two mechanisms are kinematically similar. The degree of similarity is S1 since every link of the sun-ring-planet train has a corresponding link in the Humpage train, but the velocities of link 2 in each are not parallel. Alternatives to the similarity relation of Fig. 8 are given in Figs. 9 and 10. If one is concerned with the motion of link 5, the planar train can be modified as shown in Fig. 9 such that each mechanism has five links. Otherwise, link 5 can be deleted as in Fig. 10, leaving each mechanism with four links. While the similarity relations of Figs. 8, 9, and 10 are each of degree S1, those in Figs. 9 and 10 may be considered better because they compare mechanisms with the same number of links. This distinction will become clear in the discussion of efficiency analysis later. Since the speed ratios in Tables 1 and 2 match, as indicated earlier, the declaration of similarity between the sun-ring-planet train and the Humpage train may seem an obvious example. The intent of this comparison, however, is to illustrate the flexibility with which the principle of similarity and equivalence can be applied, or in other words, the variety of forms that a similar or equivalent geared mechanism can take", "/Nr512Np /Nr (2) where vo and v i are the output and input velocities and Nr and Np represent the numbers of teeth on the outer ring and the flexspline, respectively. Using notation similar to that used for the simple planetary set earlier, this becomes v4 /v35~N42N2!/N4512N2 /N4 (3) and the mechanism is labeled as shown in Fig. 11. This is a planet-fixed case, and the speed-ratio results match exactly with those for the planet-fixed case of the simple planetary set. Thus the harmonic drive is kinematically similar to both gear Journal of Mechanical Design rom: http://mechanicaldesign.asmedigitalcollection.asme.org/pdfaccess.a trains shown in Fig. 10. It also has a ~more intuitive! similarity to the planar gear train of Fig. 11. Link 2 of the harmonic drive is actually the planet, and instead of moving on a carrier around the central axis of the mechanism, it flexes such that its \u2018\u2018axis of rotation\u2019\u2019 ~not well defined for a compliant member! remains at that central axis. In this way, the geared members of the harmonic drive can be assigned tooth numbers which give a dramatic reduction ratio while avoiding the problems of dynamic balancing and manufacturing a tiny carrier link" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002578_robot.2007.363787-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002578_robot.2007.363787-Figure6-1.png", "caption": "Fig. 6. Slope model of a screw", "texts": [ " Rotating gear 2 to the right results in ungrasping. The following assumption is also necessary so that the output jaw maintains the grasp and gear 2 pushes the lever in Fig. 4 (2) through (4). Assumption 2: Screws 1 and 2 are non-backdrivable, that is, they do not turn when a force is applied to the output jaw and gear 2, respectively. The condition for non-backdrivability is shown next. We review the mechanics of a screw[6]. A male screw can be regarded as a slope by unrolling it as shown in Fig. 6. A female screw can be regarded as an object on the slope. Let \u03b8 be the angle of the slope (lead angle), and \u03c1 be the friction angle between the male and female screws. Let also \u03c4\u2295 and \u03c4\u2296 be the torques to tighten and loosen the female screw, respectively. Their relationship is given by: \u03c4\u2296 = tan(\u03c1 \u2212 \u03b8) tan(\u03c1 + \u03b8) \u03c4\u2295. (2) Therefore, as \u03b8 increase, the torque \u03c4\u2296 to loosen the female screw decreases even if the tightening torque \u03c4\u2295 is the same. If \u03c1 < \u03b8, \u03c4\u2296 is negative, which means that the pair of male and female screws are backdrivable" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001770_1.27289-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001770_1.27289-Figure1-1.png", "caption": "Fig. 1 Schematic view of the trajectory resulting from the constant control torque method.", "texts": [ " The second step is a trajectory sliding on the manifold, by means of a single constant torque, until the transient goal is reached. The final step is a trajectory from the transient goal to the origin. The control timings, durations, and the sign of control torques can then be determined by calculating the intersection points between the manifold and polhode or the boosted/damped trajectory, between the trajectory sliding on the manifold and the transient goal, and between the single spin motion and the origin. A schematic representation of these trajectories is shown in Fig. 1. The idea of the constant torque manifold for stabilization of the rotational motion of an asymmetric rigid body is inspired by [12], which describes a method to obtain a trajectory of the angular velocities of an asymmetric rigid body when a single constant torque is employed along either themaximum,minimum, or middle principal moment of inertia axis. Contrary to the robust feedback schemes [9\u201311], a demerit of the proposed method is that it is not robust to themodeling errors, and external disturbances" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000688_s00170-003-1853-1-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000688_s00170-003-1853-1-Figure2-1.png", "caption": "Fig. 2. Calculating the edge contact length on the 7-parameter APT tool [6]", "texts": [ " In this manner, the simulation could effectively \u201czoom in\u201d on the cutting teeth, allowing the simulation to more accurately determine the in process chip geometry (Fig. 1c). Notice the depth buffer sized to the tool diameter and the depth buffers zoomed to the tool teeth are approximately the same size, yet the latter is a more efficient usage. Furthermore, the tool shape used in [4] was a simple flat end mill. This is but one type of tool used in milling. The theoretical 7-parameter (d, e, f , h, r , \u03b1, \u03b2) APT tool is shown in Fig. 2. Through a judicious choice of parameters, this tool may mathematically model flat end, ball nose, bull nose and even more complicated tool shapes. This exact shape tool will rarely, if ever, be used to machine a part. However, if the simulation is able to handle its complicated geometry, then it is capable of handling the wide assortment of tool shapes used in manufacturing. 2.2 Improvements in the determination of the edge contact length, b The simulation described herein is ideally suited for determining the edge contact length, b", " The depth buffer measures the depth of each element from the current viewing datum, and is denoted by the variable p. As the simulation is concerned with two states of the depth buffer, p\u2032 refers to the state of the depth buffer before the rendering of the current tool position and p the state after. For a simple flat end mill, the maximum depth of cut can be calculated as \u2206p = pmax \u2212 p\u2032 min, and the edge contact length is simply b = \u2206p/ cos \u03c8, where \u03c8 is the cutter helix angle. However, for the APT tool of Fig. 2, the edge contact length may only be calculated by projecting the segment \u2206p onto the tool shape. Referring to Fig. 2, the tool shape consists of three distinct piecewise sections \u2013 a linear section OA, described by the angle \u03b1, a circular arc AB described by centre point (e, f ) and radius r , and a linear section BC., described by the angle \u03b2. The linear section projections are determined by similar triangles and the circular projection are determined by the subtended angle of the depth intersection point. If f(z) is a function that computes the edge contact length of a height z in the shown tool coordinate system, then the overall edge contact length, b, can be calculated from: b = f(p\u2032 min \u2212 p0)\u2212 f(pmax \u2212 p0)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003231_xst-2010-0242-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003231_xst-2010-0242-Figure7-1.png", "caption": "Fig. 7. Force computation by summating four force components on every contacted bur surface element (e.g. A) that is considered as burring the tissue recorded in the closest endpoint (p). Solid circles: endpoints contacting tissue. Hallow circles: endpoints contacting null. (A) Lateral view. (B) Top view.", "texts": [ " l indicates the vector from C to c along the bur axis, i.e. the \u2212Z-axis. C and c are the centers of the bur and a cross-section that the element is located on. r is the radius of this cross-section. X, Y and Z are the vectors along the X-axis, Y -axis and Z-axis, respectively. Rmeans a rotation matrix about the Z-axis. \u03b1 indicates the angular position of the element relative to the X-axis. The element area, dA is calculated as r d\u03b1 dl. d\u03b1 indicates an equally divided arc on the cross-section. dl indicates the equal interval along the bur axis (Fig. 7). A force computation is implemented if any cutting counter of the three line sets is not 0. Although any line set can be used, the one with the largest cutting counter is used to determine if a surface element contacts tissues. First, the element coordinate is transformed into the volume coordinate in which two (e.g. x and y) components of the element is compared with the same two (x and y) components of the (e.g. z) axis-parallel lines in the compared set to determine the line (e.g. L) closet to the element. If the two endpoints of the closet line (as p and q of L in Fig. 7) both cut the same tissue, the element is set as cutting this tissue. If only one endpoint cuts some tissue or the two endpoints cut different tissues, the third (e.g. z) coordinate components of the element and the endpoints are used to determine which endpoint is near the surface element (as p not q). The tissue removal load on a cutting element is represented as a tangential (F tang), a radial (Fradius), an axial (Faxial) and a trust (Ftrust) components that are calculated as follows (Fig. 7), Ftang(Fradius, Faxial or Ftrust) = Kh(Kr,Ka or Kt)dA frate. The tangential component is set against the rotation, the radial component toward this axis and the axial component along the bur axis. Different from other tools such as drills or saws that cut tissues only along the tool axis, most burs can cut tissues freely along any direction. Therefore, the trust component against the bur moving direction is added. Every component is set as proportional to the tissue removal rate that equals to (dA frate), the product of the element area (dA) with the feed rate (frate) of a haptic step", " The coefficients,Kh,Kr ,Ka andKt actually depend on many variables such as the cutting velocity (a product of the rotational speed with the cross-section radius that the surface element located on), the feed rate, the bur type and tissue types (cortical or cancellous bones), herein are determined by testing variousKh, Kr, Ka and Kt to bur frozen bones. The four force components are then transformed and summated into FX , FY and FZ , the forces along the X-, Y - and Z-axis as the following equations to be rendered by the haptic device, (Fig.7) FX = \u2211 (cos\u03b1Fradius \u2212 sin\u03b1Ftang \u2212 (f \u00b7 X)Ftrust). FY = \u2211 (sin\u03b1Fradius + cos\u03b1Ftang \u2212 (f \u00b7 Y )Ftrust). FZ = \u2211 (Faxial \u2212 (f \u00b7 Z)Ftrust). f indicates the bur moving direction vector. The trigonometric functions are also pre-computed to save computation time. Meanwhile, if only one endpoint in each of the three line sets cuts tissues, the real cutting area on the bur surface much differs in the case of initially burring tissues or the case of already immersed into tissues for tens of haptic steps. The force computations are then implemented three-times for all the three line sets respectively to render the average FX , FY and FZ " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001657_bfb0119410-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001657_bfb0119410-Figure5-1.png", "caption": "Figure 5: Three-Fingered End-Effector", "texts": [ " This highforce capability could be useful for manipulator-aided mobility or trap recovery. The base of the arm is mounted to a six degree-of-freedom force-torque sensor, which is used for control (see Section 4.1) [8]. which is used for control (see Section 4.1) [8]. Several end-effector concepts have been developed for handling rock samples. The most effective has proven to be a lightweight three-fingered end-effector. It utilizes flexural joints and relies on shape-memory alloy (SMA) actuation (see Figure 5). Each of the three fingers are formed from 1/8\" steel rod. A nylon mounting plate has integral flexures that allow motion without bearing surfaces, eliminating the need for lubrication and considerably simplifying design and fabrication. A 0.006\" diameter Flexinol SMA wire provides retracting force to each finger from its normally closed position. The wires are connected to the ends of the fingers, and run along the bottom of the mounting plate to increase their working length, and thus allow greater finger travel" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003232_978-3-642-15615-1_10-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003232_978-3-642-15615-1_10-Figure1-1.png", "caption": "Fig. 1. Mechanism of exoskeleton knee", "texts": [ " Secondly, a part of data is selected randomly as testing sample, and then is inputted into the trained neural network model. Thirdly, the predicted knee angle is outputted from the model. Fourthly, the correlation coefficient is used for evaluating the neural network. Finally, this output signal as control signal is delivered to the motor controller and can drive the exoskeleton knee to move in the same way. Exoskeleton knee is a part of the powered gait orthosis which is being developed which can move legs of a patient in a physiological way on the moving treadmill. The exoskeleton knee mechanism is shown in Fig. 1. It is composed of two connecting rods, bearings, lead screw and motor. A linear series elastic actuator with a precision ball screw is connected between a torque and the knee. There are many muscles in the human lower limb. Different muscles have different effects and roles on the gait movements. Biceps femoris (BF), semitendinosus (SEM), vastus medialis muscle (VMO), rectus femoris (RF) and vastus lateralis (VLO) are primary to knee joint and they have strong surface EMGs which can be detected and analyzed easily" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure4-1.png", "caption": "Fig. 4. Induced curvature.", "texts": [ " k\u00f01\u00den ; k\u00f01\u00deg ; k\u00f02\u00den ; k\u00f02\u00deg are normal curvature and geodesic curvature of two directrixes at conjugate point, k\u00f021\u00de g and k\u00f021\u00de n denote induced geodesic curvature and induced normal curvature. Obviously, induced principal curvature k\u00f021\u00de b1 of conjugate normal circular-arc surfaces can be come down to induced geodesic curvature and induced normal curvature of conjugate directrixes, which exposes the importance of the directrix. Referring to the diagrammatic method proposed in Section 3.1, here main properties of induced principal curvature are discussed. In Fig. 4, point P is an instantaneous conjugate point, and C2 is instantaneous contact generating circle. a2, a3 are common frame axes, a2 is on common tangent plane of two datum surfaces, a3 is unit common normal to datum surfaces. Geodesic curvature centers O\u00f01\u00deg ;O\u00f02\u00deg and normal curvature centers O\u00f01\u00den ;O\u00f02\u00den of two directrixes at a2-axis and a3-axis can be determined, respectively, so two curvature axes O\u00f01\u00den O\u00f01\u00deg Oc;O \u00f02\u00de n O\u00f02\u00deg Oc of two directrixes can be acquired, Oc is their intersection point", " It can be proved that the denominator in Eq. (17) is always positive, so noninterference condition may be judged by its numerator and noninterference condition can be represented as follows: k\u00f021\u00de g cos h\u00fe k\u00f021\u00de n sin h 6 0 \u00f0\u2018\u2018the Outward Rule\u2019\u2019\u00de P 0 \u00f0\u2018\u2018the Inward Rule\u2019\u2019\u00de \u00f018\u00de When choose the sign \u2018\u2018=\u2019\u2019 in Eq. (18), the following expression can be obtained: tan h \u00bc k\u00f021\u00de g k\u00f021\u00de n \u00f019\u00de h*, h* + p can be acquired from Eq. (19), which are angular condition of boundary points N(1), N(2). Now returning to Fig. 4, more relationships between the collocation of gear teeth entity and the noninterference condition can be obtained. Because the principal curvature k\u00f01\u00deb1 is nonnegative at every point of semicircle N(2)P(1)N(1), the body of gear teeth 1 must be chosen in the semicircle; for the principal curvature k\u00f01\u00deb1 is nonpositive at every point of semicircle N(1)P(2)N(2), the body of gear teeth 1 must be chosen out of the semicircle. This shows that under the condition of satisfying body collocation, two semicircles, whose endpoints are N(1), N(2), can be regarded as tooth profiles of normal circular-arc gears" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003491_1.3456118-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003491_1.3456118-Figure1-1.png", "caption": "Fig. 1. Forces on a rocking semicircular hoop.", "texts": [ " For the fixed point, we choose either the contact point of this object with the incline or else its center of mass. In this paper, we demonstrate that the equivalency of these two approaches for the wheel or hoop is an artifact of their azimuthally symmetric geometry. As a counterexample, consider the results from these two approaches for a uniform density semicircular hoop of radius R and mass M that is constrained to be within a vertical plane as it rocks without slipping on a horizontal surface. The hoop\u2019s point of contact with the surface is denoted by P, as shown in Fig. 1. The semicircular hoop is centered about point O, and its center of mass is located at point C, a distance 2R / from point O along the perpendicular bisector of the hoop. The distance OP is R. When the rocking hoop is oriented at an angle , as shown in Fig. 1, we can verify the following lengths that are needed in our calculations: OB = 2 cos R, BP = 1 \u2212 2 cos R , 1 BC = 2 sin R, CP = 1 + 4 2 \u2212 4 cos 1/2 R . 2 The virtue of choosing the contact point P for the evaluation of the torque is, as we emphasize to students, that this choice permits us to ignore the effects of the frictional and normal forces, Ff and Fn, because they exert no torque about the contact point. Only the force of gravity Mg, which can be thought of as acting on the center of mass C, exerts a torque about P", " 6 We next compare this equation of motion with that obtained from the alternative approach. Instead of using P, we evaluate the total torque about the center of mass C and use the moment of inertia about this point to derive the equation of motion for . This calculation entails the use of the frictional and normal forces of the surface on the hoop. To describe the motion of the center of mass, we use a Cartesian coordinate system. We choose its origin to be the location of O when the hoop is horizontal; that is, when =0. When the hoop is tilted, as shown in Fig. 1, the position of the center of mass is seen to be 905\u00a9 2010 American Association of Physics Teachers ense or copyright; see http://ajp.aapt.org/authors/copyright_permission xcm = 2 sin \u2212 R, ycm = \u2212 2 cos R . 7 The second contribution to xcm is due to the horizontal motion of point O as the disk rocks back and forth without slipping. We take the second time derivative of these coordinates to obtain the following Cartesian components of the center of mass\u2019s acceleration in the laboratory : x\u0308cm = \u2212 1 \u2212 2 cos \u0308 + 2 sin \u03072 R , 8a y\u0308cm = 2 sin \u0308 + cos \u03072 R . 8b Because the frictional force induces the horizontal acceleration of the center of mass and the normal and gravitational forces produce the vertical acceleration of the center of mass, we can write Ff = Mx\u0308cm, 9 Fn \u2212 Mg = My\u0308cm. 10 The total torque about the center of mass is comprised of only the torques due to the frictional and normal forces. We find the respective lever arms of these forces from Fig. 1 and write cm = BPFf \u2212 BCFn = \u2212 MR2 1 \u2212 4 cos + 4 2 \u0308 + 2 sin \u03072 + 2 sin g R , 11 and set this torque equal to Icm\u0308, MR2 1 \u2212 4 2 \u0308 = \u2212 MR2 1 \u2212 4 cos + 4 2 \u0308 + 2 sin \u03072 + 2 sin g R 12 or 906 Am. J. Phys., Vol. 78, No. 9, September 2010 Downloaded 28 Oct 2012 to 152.3.102.242. Redistribution subject to AAPT lic 1 \u2212 2 cos \u0308 = \u2212 sin \u03072 \u2212 sin g R . 13 This result differs from that of Eq. 6 . Which is correct? Using a Lagrangian formulation demonstrates the correctness of Eq. 13 . From the viewpoint of our earlier result, the first term on the right-hand side of Eq", "1 The angular momentum of a rigid object about any point, say S, is given by the sum over its constituent masses, imi ri\u2212rS vi\u2212vS . Thus, in addition to the customary r F terms, its time derivative has the additional phantom torque term, ph = \u2212 i mi ri \u2212 rS aS = \u2212 M rcm \u2212 rS aS, 14 where aS is the acceleration of point S. In the previous example, we denoted S by P. We now find the phantom torque for the rocking hoop. As the semicircular hoop rocks, the trajectory of point P, which is in contact with the horizontal surface in Fig. 1, can be described in terms of its displacement from the point of contact. We use the quantities exhibited in Fig. 2 to describe this displacement. The fixed angle specifies the angular displacement of the line intersecting the contact point from the line that intersects the center of mass C. The angle specifies the angular displacement of the center of mass from the vertical. We use these definitions and take into account the horizontal motion of point O as the hoop rocks to obtain the horizontal and vertical components of the displacement, velocity, and acceleration of P from its point of contact, xP = R \u2212 \u2212 sin \u2212 , 15 yP = R 1 \u2212 cos \u2212 , 16 x\u0307P = R cos \u2212 \u2212 1 \u0307 , 17 \u02d9 \u02d9 yP = \u2212 R sin \u2212 , 18 906L" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003845_1.4003270-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003845_1.4003270-Figure6-1.png", "caption": "Fig. 6 3RRR: state 1", "texts": [ "org/ on 02/03/2 3RRR mechanism that moves through three states. In the second example, a nine-bar linkage will be analyzed. The third example will focus on a planar metamorphic mechanism that was analyzed by Lan and Du in 16 . The last example will show how mechanism state matrices work well for topologically identical mechanisms with different configurations. 6.1 3RRR Mechanism. Consider the 3RRR2 mechanisms in Figs. 6\u20138. For this example, it is assumed that the mechanism transforms from state 1 in Fig. 6 to state 2 in Fig. 7 to state 3 in Fig. 8. In state 1, link 1 is the only fixed link. In state 2, both links 8 and 1 are fixed links, and in state 3 links 7 and 1 are designated as fixed links. The mechanism state matrix for the three states is given in Eq. 20 as MSM = 2Z R,6Z R,8Z R 3Z R 4Z R 5Z R,7Z R 6Z R 8Z R 2Z R,6Z R,8V X 3Z R 4Z R 5Z R,7Z R 6Z R 8Z R 2Z R,6Z R,8V X 3Z R 4Z R 5Z R,7Z R 6Z R 8V X 20 23RRR denotes a manipulator with three kinematic chains of the type revolute-revolute-revolute" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.44-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.44-1.png", "caption": "Fig. 14.44 Principle of operation of rotary type concrete pump", "texts": [ "5\u20135 m Charging hopper capacity: 350\u2013500 dm3 Rated pumping capacity: 36\u2013150 m3/h Concrete pumping pressure: 45\u2013130 bar Concrete mix pumping Cylinder diameter: 160\u2013230 mm Piston stroke: 1000\u20132100 mm Number of boom segments: 2\u20135 Self-propelled concrete pumps are also made in versions adapted for attaching a pipeline made from steel pipes for conveying concrete mix over considerable distances. Besides piston-type concrete pumps also rotarytype pumps, based on the principle of the peristaltic pump, are manufactured (Fig. 14.44). In the rotary-type concrete pump concrete mix is pumped as a result of squeezing it out of a reinforced rubber hose by two rollers attached to the rotor. The hose recovers its circular cross-section owing to elastic restoration or negative pressure inside the casing (vacuum restore). The pumping pressure in rotary-type pumps amounts to about 3 MPa, allowing concrete mix to be delivered over a distance of about 200 m in the horizontal plane and to an elevation of about 80 m. The design of the rotary-type concrete pump is simple but the conveying hose needs to be replaced quite often" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000057_s0069-8040(08)70029-3-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000057_s0069-8040(08)70029-3-Figure4-1.png", "caption": "Fig. 4. Mass transfer from an impinging jet electrode. (From ref. 48 by permission of the publisher, The Electrochemical Society, Inc.). I, Potential core region; 11, established flow region; 111, stagnation region; IV, wall-jet region.", "texts": [ " [38] and this expression for No was obtained by Braun [39] and subsequently experimentally verified [ 401 . 2.3.4 Electrodes based on impinging jets When a jet of fluid submerged in a medium of that fluid strikes a surface perpendicularly, it spreads out radially over that surface. Original interest in these systems was due to mass transfer investigations of downward directed jets of vertical-take-off aircraft [ 4 1 ] , though other applications such as electrochemical machining are important. We may identify several regions as shown in Fig. 4 . In particular, there is the stagnation region (Region 111) and the wall-jet region (Region IV) which give rise to the wall-tube and wall-jet electrodes, respectively. The relative sizes of electrode and impinging jet are thus most important. References pp. 434-441 374 (a ) Wall-jet electrodes The laminar flow hydrodynamics for a radial wall-jet were first considered by Glauert [41] and subsequently by Scholtz and Trass [42] . A more complete evaluation for electrochemical purposes has recently appeared [ 4 3 ] " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001064_j.conengprac.2006.05.002-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001064_j.conengprac.2006.05.002-Figure1-1.png", "caption": "Fig. 1. Shaft assembly, motor coupling and radial bearing stators for a five DOF magnetic bearing system.", "texts": [ " Finally, in Section 5 the test bench is described, and in Section 6, an experimental comparison of the controllers is provided. While modeling of AMBs is relatively established (Schweitzer et al., 1994), their dynamics are briefly reviewed here to establish notation which is used in control design. The system has five DOF and consists of a horizontal shaft onto which two journals and a disk are mounted. The journals form part of the radial magnetic bearings, and the disk is part of the axial magnetic bearing. The entire shaft assembly is coupled to a DC motor via a helical coupling. Fig. 1 shows the configuration of the shaft and magnetic bearings. Fig. 2 shows a schematic of a crosssection of the system. Assuming a rigid shaft the dynamic equations are (Matsumura & Yoshimoto, 1986): m \u20acx \u00bc Fx, m \u20acy \u00bc Fb;y \u00fe F f ;y \u00fe Fc;y \u00femgy, m\u20acz \u00bc Fb;z \u00fe Ff ;z \u00fe F c;z \u00femgz, Jz \u20acc \u00bc \u00f0lf ;a \u00fe x\u00deF f ;z \u00f0lb;a x\u00deFb;z Jxo_y\u00fe lcFc;z, Jy \u20acy \u00bc \u00f0lb;a x\u00deFb;y \u00f0lf ;a \u00fe x\u00deFf ;y \u00fe Jxo _c lcF c;y, (2) T.R. Grochmal, A.F. Lynch / Control Engineering Practice 15 (2007) 95\u2013107 97 where x; y; z denote the coordinates of the center of mass cm relative to the origin O of the inertial frame", " This system is available from SKF Magnetic Bearings (Calgary, AB) and its intended use is to study rotordynamics and control strategies under no-load operating conditions. Fig. 4 shows the test stand incorporated into the experimental setup. The system\u2019s specified range of shaft speed is 2000\u201310000 rpm and has a maximum shaft speed of approximately 15 000 rpm. The system allows several shaft configurations. For the following experiments a 304.8mm (1200) long shaft with a 9.5mm (3=800) diameter was used. Refer to Fig. 1 for a diagram of the assembly drawn to scale. The mountable radial journals have an outer diameter of 34.3mm, a length of 48.0mm, and span 178.0mm from center-to-center. The mountable axial disk has an outer diameter of 66.0mm. Table 3 provides additional radial and axial bearing specifications. Each bearing housing consists of the stator, rotor gap sensors (two-sided variable reluctance type), and the touchdown bearings. The touchdown bearings T.R. Grochmal, A.F. Lynch / Control Engineering Practice 15 (2007) 95\u2013107 101 prevent contact between the rotor and stator and rest the delevitated shaft assembly" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003772_6.2010-8363-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003772_6.2010-8363-Figure9-1.png", "caption": "Figure 9. 3D Guidance Logic Geometry.", "texts": [ " The logic used in Ref. [1] for trajectory tracking works by defining an imaginary point r moving along the desired current segment { , }a b of the path, like a pseudo-target. For small departures from the desired path this nonlinear guidance law behaves like a proportional derivative PD controller, minimizing the cross track error. Within the nonlinear space, the control logic formulation is better adjusted to a circular path. The guidance logic designed for this project minimize the angular error (see Figure 9) between the inertial velocity vector V of the UAV and the vector L subtended between the position of the UAV and the point r , ( r is located at a fixed distance from the closest point d toward point b ), for both longitudinal and lateral planes separately (see figure 10 where both planes were separated). Those errors will be zero only when both vectors are aliened and the UAV is flying zero cross distance in each plane. D ow nl oa de d by D E A K IN U N IV E R SI T Y o n A ug us t 1 1, 2 01 5 | h ttp :// ar c" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002952_j.proeng.2010.04.144-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002952_j.proeng.2010.04.144-Figure1-1.png", "caption": "Figure 1 Images from the FE model for a perpendicular impact at the GSC at 28.8 m\u00b7s-1.", "texts": [ " Non-spinning impacts normal to the face of the racket were simulated in the FE model at 5 locations on the stringbed, for each nominal velocity (Table 3). The inbound velocity of the ball was assigned in the FE simulations using INITIAL_VELOCITY _GENERATION. The initial ball velocity, impact location on the stringbed and racket properties were all defined using the Tennis Design Tool (TDT) [25]. Five FE simulations were undertaken for each nominal velocity, resulting in a total of 15 simulations. An example of an impact with a velocity of 28.8 m\u00b7s-1 at the geometric stringbed centre (GSC) is shown in Figure 1. In the M\u03a8O the rebound velocity of the ball was measured using light gates, which are located 1.5 m from the impact point [20]. Assuming the long axis of the racket is vertical at the point of impact, the light gates measure the rebound velocity of the ball in a direction normal to the face of the stringbed. Therefore, the perpendicular velocity from the FE model was compared with the M\u03a8O data, as opposed to the resultant velocity. A simple trajectory model was applied to the FE results to predict the velocity of the ball when it had travelled 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001872_physreve.76.051905-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001872_physreve.76.051905-Figure1-1.png", "caption": "FIG. 1. Color online a A fully inelastic collision of two microtubules mediated by a molecular motor. b Schematic representation: The molecular motor is attached symmetrically to two flexible microtubules at a distance s along the tubule from the midpoint.", "texts": [ " Since most of the in vitro experiments with purified microtubules and motors are quasi-two-dimensional i.e., the length of the microtubules typically exceeds the depth of the 1539-3755/2007/76 5 /051905 12 \u00a92007 The American Physical Society051905-1 container , we consider the two-dimensional case only. While our technique can be also applied to the fully threedimensional case with minor modifications, analysis of the potentially rich high-dimensional effects goes beyond the scope of this work. A two-dimensional collision of two microtubules is schematically illustrated in Fig. 1 a . Before the interaction, the microtubules are oriented at angles 1 b and 2 b. The simultaneous binding of the molecular motor to the two microtubules and the subsequent marching along them results in a complete alignment of the latter; after the interaction the microtubules are oriented at angles 1 a and 2 a, where 1 a= 2 a = 1 2 1 b+ 2 b . We refer to this type of interaction as a fully inelastic collision, by analogy with the physics of inelastically colliding grains see, for example, Ref. 13 ", " That is, we focus on the case where the endpoints of the microtubules are stabilized, for example with taxol, so that the polymerization and depolymerization processes, which may affect the lengths of the microtubules, are insignificant. Furthermore, for simplicity we assume that the molecular motor attaches symmetrically to the microtubules. Thus, the two attachment points are at the same position on each rod with respect to their respective midpoints, and the force exerted by the motor is perpendicular to the bisector of the microtubule pair; see Fig. 1 b . The last conclusion follows from the assumption that the motor, while moving along the filaments with a constant speed, acts as a strong spring bringing the two motor heads together. In Appendix A we show that the motor has a tendency to orient perpendicular to the bisector of the microtubule pair even if the motor has a nonzero length. Since the initial attachment may occur at a random position on the tubule, we are interested in the properties of the interaction in particular, the inelasticity coefficient averaged with respect to the initial attachment position", " This assumption is not significant and can be relaxed, although it does not affect the qualitative relation between filament flexibility and the inelasticity factor. To describe the motion of interlinked microtubules, we combine the theory of Refs. 15,16 for a semiflexible polymer with the analysis of the rigid case in Ref. 9 . We adopt a two-dimensional setting and model a microtubule as a semiflexible homogeneous inextensible elastic rod of length L and bending stiffness . We measure locations along the rod relative to the rod\u2019s midpoint, using the arclength s as the natural parameter, so \u2212 1 2L s 1 2L see Fig. 1 b . The inextensibility of the rod implies that its embedding in the Euclidean plane preserves the arclength element at all times. Thus, if r s is the position in the plane of the point s on the tubule and rs denotes its derivative with respect to s, then we have the local constraint rs \u00b7rs=1. A molecular motor attaches initially to the tubule at the point si and moves along the tubule with the constant velocity v, exerting a force f on the tubule. As long as the force does not exceed a critical value, we may assume that the velocity of the motor does not depend on the force f", " Since the velocity is fixed, the movement of the attachment point sa is subject to the kinematic constraint sa t = si + vt . 6 In a binary collision, the molecular motor attaches to and moves along two microtubules simultaneously, and the relative configuration of the tubules changes as a result of the motor force acting on both tubules. As explained above, we consider only symmetric interactions, where the force is normal to the bisector. Then we can select a Cartesian coordinate system where the y-axis is directed along the bisector see Fig. 1 b , so f= \u00b1f ,0 , where the magnitude f of f has to be deduced from the kinematic constraint. B. Governing equations The equations governing the motion of the microtubules are derived from the balance of forces. If the viscosity of the medium containing the mixture is large Stokes limit , then the viscous drag force is balanced by the force acting on the tubules. The latter is the variational derivative of the energy functional E measuring the bending energy of the tubule, together with the energy of the inextensibility and the motor attachment constraints E = 1 2 \u2212L/2 L/2 g rs \u00b7 rs \u2212 1 + rss \u00b7 rss + f \u00b7 r s \u2212 sa ds " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003570_j.triboint.2011.09.004-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003570_j.triboint.2011.09.004-Figure4-1.png", "caption": "Fig. 4. Schematic illustration of an elastomer shaft seal.", "texts": [ " In both cases during sliding the surfaces of the elastomer and the counterpart are separated by the local clearance h(x1, x3), given as: h\u00f0x1,x3\u00de \u00bc c\u00fect\u00feh0\u00f0x1,x3\u00de\u00fed\u00f0x1,x3\u00de \u00f02\u00de where h0(x1,x3) is the local depth of the dimple (see [18]) and ct is local additional clearance along the tapered edges: ct \u00bc ht x1 lt 1 for x1o lt ct \u00bc 0 for ltox1o l1 lt ct \u00bc ht x1\u00fe lt l1 lt for x14 l1 lt 8>>< >>: \u00f03\u00de It should be emphasized that the local film thickness h varies in both x1 and x3 directions since the dimples have the form of spherical segments distributed on the surface lying in the plane x1x3. Besides the film thickness varies in the x1 direction due to the tapered edges. Further clarification of the coordinate system will be given below during the discussion of Fig. 3. An additional scheme for the corresponding problem showing a shaft elastomer seal with the coordinate system is presented in Fig. 4. A stationary elastomer sleeve with the cross-section shown in Fig. 1 is fitted on a rigid smooth shaft moving axially in the x1 direction (Fig. 4(a)). A regular surface texturing is applied to the inner surface of the sleeve as shown schematically in Fig. 4(b). The x2 axis corresponds to the radial direction and the x3 axis corresponds to the circumferential direction. Due to the surface texturing a hydrodynamic pressure is generated in the radial gap (filled with lubricant) between the moving shaft and the stationary elastomer sleeve, causing deformation of the latter. Since the fluid film thickness is much smaller than the shaft radius, then curvature in the circumferential direction x3 can be neglected. This allows using Cartesian instead of cylindrical coordinates as in Figs" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.143-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.143-1.png", "caption": "Fig. 14.143 Small-sized crane-manipulator for transporting and laying aerated concrete blocks", "texts": [ " Forming of Concrete Structures for Sliding Forms The automation of concrete structure forming mainly applies to sliding forms used in the construction of dams and structures such as chimneys, towers, silos, and bridge piers. A few methods of automating the erection of structures by sliding forms are described in [14.53]. In European countries (Germany, the UK, and The Netherlands) research aimed at developing robots for masonry work has been conducted since 1991. This work has been focused on robots for erecting external and internal walls from aerated concrete and gypsum blocks. A crane-manipulator (Fig. 14.143) for transporting and assembling 0.6\u00d70.9\u00d71.0 m aerated concrete blocks has been developed in The Netherlands. The prototype robot shown in Fig. 14.144 was developed as part of the European Rocco project in 1995. Research on the mechanization and automation of block-work wall erection focuses on the following equipment: \u2022 Special bricklayer\u2019s platforms for adjusting the bricklayer\u2019s position to the height of the wall under construction\u2022 Special winches or air-driven arms for compensating for the gravity of the blocks so that the latter can be manipulated weightlessly Information on the numerous areas of research into robots for masonry work can be found in [14" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003271_s00170-011-3475-3-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003271_s00170-011-3475-3-Figure8-1.png", "caption": "Fig. 8 Casual two points on the involute curve", "texts": [ " In a related paper, it was seen that measurement results are in a good agreement with theoretically calculated values. As mentioned above, geometric accuracy of the gear manufactured by the end mill in the CNC milling machine according to the radial cutting method was merely measured. In this paper, the cutting errors of the tooth profile of the spur gear manufactured in the previous paper are investigated. Therefore, a n angle which describes coordinates in plane XY of points on the involute curve is considered (see Figs. 2 and 6). Points G and H show casually two points on the involute curve (see Fig. 8), where angles =G and =H show the involute functions corresponding to these points. In the radial cutting method, the end mill linearly cuts interval of these casual points on the tooth profile curve according to increment value (am) in angle an (see Fig. 5). In this case, a cutting error occurs on the involute tooth profile (see Fig. 9). In this study, Fig. 10 illustrates how to determine the quantity of this cutting error. Point L in Fig. 10 is also a point on the involute tooth profile, but this point is assumed as the midpoint of points G and H" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000704_s11768-006-5265-2-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000704_s11768-006-5265-2-Figure2-1.png", "caption": "Fig. 2 Illustration of stabilization by MLF.", "texts": [ "(5) consisting of two subsystems are asymptotically stable with initial condition x0 \u2208 \u03a6, if there exist Fi, Pi > 0, scalar \u03c1i > 0, \u03bb1 0, \u03bb2 0 such that, (A1 + B1H \u2217 1F1)TP1 + P1(A1 + B1H \u2217 1F1) < \u03bb1(P2 \u2212 P1), (14) (A2 + B2H \u2217 2F2)TP2 + P2(A2 + B2H \u2217 2F2) < \u03bb2(P1 \u2212 P2), (15)\u23a1 \u23a3 Pi hiF T i hiFi \u03c1iw 2 i \u23a4 \u23a6 0 (16) where the attractive domain \u03a6 = \u03a61 \u22c3 \u03a62. The switching rule \u03c3 = argmin{Vi(x)}, i = 1, 2. Proof From Eq.(14)\u223c(15), we can obtain that If xT(P1 \u2212 P2)x 0 and x \u2208 \u03a91, V\u03071(x) < 0; If xT(P2 \u2212 P1)x 0 and x \u2208 \u03a92, V\u03072(x) < 0. The switched system is asymptotically stable evidently for the given switching rule. This completes the proof. It can be seen the cone-type regions outside \u03a9 (such as the cone m-n-r in Fig. 2) is excluded from the domain of attraction that in Theorem 1 or 2. However, in some cases, we can take these cone-type regions into account to enlarge the attractive domain. Let \u03a8iaja denote the cone outside \u03a9 which is created by two lines lia, lja. \u03a8iaja can be represented as { F\u0304ix + f\u0304ia > 0 } \u22c2 { F\u0304jx + f\u0304ja > 0 } , where F\u0304ka equals either Fka or \u2212Fka; f\u0304ka equals either fka or \u2212fka, k = i, j. Equivalently, the region \u03a8iaja can be also written as Fiajax + fiaja > 0, where Fiaja = \u23a1 \u23a3 F\u0304ia F\u0304ja \u23a4 \u23a6 , fiaja = \u23a1 \u23a3 f\u0304ia f\u0304ja \u23a4 \u23a6 " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001690_s0022-0728(73)80135-4-Figure18-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001690_s0022-0728(73)80135-4-Figure18-1.png", "caption": "Fig. 18. Schematic diagram for the influence of a.c. voltage on the drop life time of the DME, according to ref. 37.", "texts": [ " INSTRUMENT FOR ELECTROCHEMICAL PROCESSES 277 Analysis of the polarograms of the square-wave, Fournier and RF techniques yields only complex curves of reoxidation and rereduction on the corresponding half-periods of the polarizing signals. These curves can be treated by a procedure analogous to that described for the anodic-cathodic irreversible d.c. polarograms 19\"11 by using diagrams in which the reversible system is represented by a straight line. More accurate results could be obtained by solving Koutecky's general equation for irreversible electrode reaction in a.c. polarography 17. 1 Tachi and Okuda 37 have investigated the influence of a.c. polarizing voltage on drop time. In Fig. 18 a schematic diagram of the dependence of drop time on the mean potential of the DME is presented. The amplitude of the polarizing voltage can be adjusted by using the instrument described here, in the range 0.1 to 100 mV. According to the results of Tachi and Okuda 37 the influence of polarizing potential on drop time is, for amplitudes of 0 < A E < 5 0 mV, similar to that for the d.c. technique. At higher amplitudes this is valid only for a certain potential range (for example at AE=100 mV the range is -0 " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003638_tiga.1969.4181029-Figure12-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003638_tiga.1969.4181029-Figure12-1.png", "caption": "Fig. 12. Overall system effects of springy shaft and possible remedies. (a) Basic system with stiff shaft. (b) Instability at fp due to springy shaft. (c) Stabilization by.f, reduction to f,'. (d) Stabilization by low-pass filter. (e) Notch-filter stabilization. (f) Notch filter not quite right.", "texts": [ " Since the damping and therefore v are so very small, this resonant peak can be very high (a factor of 1000 [60 dB] or more). To insure that this does not cause instability, we have to be sure that this peak does not raise the loop gain up to more than unity. The phase shift of the system at the normally relatively high frequency fp is almost certainly enough to cause a self-magnifying oscillation at that frequency if the loop gain there is greater than one, so there is a net amplification arounid the loop. Fig. 12 illustrates what happens and what can be done. Fig. 12(a) is basically the same plot as Fig. 3, except that more small time delays have been added in the highfrequency region (above the crossover frequency fG) to make the plot more realistic. Fig. 12(b) adds the two inertia and spring effects, producing instability in the fp frequency range because the resonant peak raises the loop gain to more than 1.0 in that range. Fig. 12(c) shows how slowing down the whole system by decreasing the crossover frequency from f, and f,' tends to lower the top of the resonant peak and stabilize the system. This is probably the most common stabilization method. Fig. 12(d) illustrates the use of a very sharp low-pass filter with its break point between fc and f,. The suitability of this method depends on the separation between fc and fp. The filter should have little effect at thef, frequency (otherwise the filter itself would cause instability) but enough effect at f, to pull the resonant peak down below the unity gain line. This technique is used quite often in conjunction with that of Fig. 12(c). Fig. 12(e) and (f) illustrates the use of a narrow bandreject notch filter to remove the resonant peak at f,. In Fig. 12(f) the filter notch does not quite coincide with the resonant peak, and instability still exists. The practicality of this method depends on how consistent the resonant point fp remains and how easily a drift-free field adjustable notch filter can be fabricated. Fig. 12 indicates that the height of the resonant peak at fp is a very critical quantity, and Fig. 11 indicates that this varies inversely with \u00a2, the mechanical system damping ratio. This ratio v is known to be very small since these mechanical systems are known to be highly oscillatory and very lightly damped. Experimentally, it may be possible to excite the system into oscillation and examine how much the freely oscillating system decays in amplitude per cycle of oscillation. Fig. 13 relates this percent decrease per cycle to the damping ratio t", " We will assume our system has a 286 CARTER: MECHANICAL FACTORS AFFECTING ELECTRICAL DRIVE PERFORMANCE % DECREASE/CYCLE 50 / 25- O *_ DAMPING .0 1.002 .004 .01 .02 04 .08 RATIO S Fig. 14. Minimum characteristic response time (torsional vibrationl limited). ADDITIONAL .LOOPGAIN FACTOR (LOG SCALE) AP 2 S FREQUENCY. HE ( LOG SCALE ) Fig. 15. Repeat of Fig. 11 with tachometer at load rather than at motor. characteristic time response as shown in Fig. 2, and that its gain drops off at an ever increasing rate at frequencies above its crossover frequency (in the manner quite common in high-performance feedback systems), as shown in Fig. 12(a). Based on these assumptions we can construct Fig. 14, which relates inertia distribution, mechanical damping, resonant frequency fp f 2r lJJM/(J+ JM)Hz and the resultant minimum response time TR. Assume, for instance, that we want a speed regulator to completely effect a change in speed in about 0.6 second. (Inherent in this statement is an assumption that the change in speed is small enough to avoid having the drive go into current limit.) The speed changes might be those called for by a tension regulator, for instance, or by a computer-driven preset draw scheme)", " Assuming a mechanical system damping on the order of a 5 percent decrease per cycle, our mechanical system would have to have stiff enough shafting to provide a resonant frequency f, in the 15-20-Hz range, depending on the relative motor and load inertias. Many paper industry drive trains have torsional damping in the 2.5-5 percent-per-cycle range and a resonant frequency in the 10-Hz area. Fig. 14 vividly points out the futility of trying to get a 0.2 second TR out of such a system without using techniques such as those shown in Fig. 12(d) or (e) (Fig. 14 assumes stabilization by crossover frequency reduction, as shown in Fig. 12(c)). It should, of course, be borne in mind that Fig. 14 attempts to give a very simple answer to a very complex question when it is actually applied to a practical system. This answer should, therefore, be considered only a crude approximation. So far in our analysis of Fig. 10, the speed we have concerned ourselves with is the speed of the inertia JM, that is, the speed of the motor. It seems logical to suggest that if we instead considered the speed of the inertia J, that is, the speed of the load, we would be bothered less by resonant speed oscillations because the inertia J is larger than JM" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000953_robot.2004.1307989-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000953_robot.2004.1307989-Figure8-1.png", "caption": "Fig. 8. Finger Control during Rotating Manipulation", "texts": [ " By considering the generated contact patterns as follows, the rhythmic rotating movements can be realized from the present simple neural model: Command of touching and rotating the object (\u201cTouch Command\u201d). Command of releasing and moving back to the initial point (\u201cRelease Command\u201d). The rotating manipulation can be performed by generating the touchlrelease command depending on the neuron\u2019s activities. When neurons are activated, the touch command is generated and when they are not, the release command is generated. Fig.8 shows the movements of each finger in the touch command and the release command. Under the touch command, two kinds of movements, \u201cApproach Mode\u201d and \u201cRotation Mode\u201d, are performed depending on the contact condition. Under the release command, two kinds of movements, \u201cRelease Mode\u201d and \u201cBack Mode\u201d, are also performed depending on the contact condition. In the rotation mode, the force Fp is generated on each finger by following equation : Fp = kpi(Pd - p ) - h i p + FC (4) where p is the current position of the object, p d is the target position of the object, p is the velocity at the current point, h,l and k,, are constant" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002336_iecon.2007.4460324-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002336_iecon.2007.4460324-Figure3-1.png", "caption": "Fig. 3. Flux linkages with respect at air gap ofPMSM with bearing damage.", "texts": [ " An harmonic pulsation of amplitude Tb and frequency Wb=24Tb were added to mechanical load torque ,TL [10]. TL =Tload +Tb coS(C)bt) (6) The value of fb it is given by equation (2) and depends on mechanical constructions parameters of the bearing. Therefore, mechanical speed and mechanical rotor angle consists both of a constant component and a sinusoidal one. Tb wo(t) = sin(cbt) +r (7) JcJbbJb The variations of the mechanical rotor angle have an influence on the rotor magnetomotive force, MMF. This could be seen in Fig. 3. Regardless of the air gap variations the amplitude of air gap MMF remains uniform [ 1]. Fmm = A sin(p,0 +2fTfs t) = A sin(pnO+ s t) (8) Where A is amplitude of MMF, t is time, 0 is angular position of rotor, Pn = up, u is number of stator harmonic, p is the pair of poles. The air gap MMF under bearing fault conditions could be expressed by combining expression (7) and (8). MMF b = A sin (Pnof +\u00b1 s t+ pAb sin(wobt)) (9) The relationship between the bearing vibration to the stator current spectra could be determined by remembering that any air-gap eccentricity produces anomalies in the air-gap flux density, as for example, the flux density distribution is not homogeneous in the stator tooth, as it is shown in Fig. 3. Since ball bearings support the rotor, any bearing defect will produce a radial motion between the rotor and stator of the machine. The mechanical displacement resulting from damaged bearing causes the machine air gap to vary in a manner that can be described by a combination of rotating eccentricities moving in both directions [2]. Fig. 4 shows frequency components and harmonics ofMMFb due to bearing damage. The air-gap flux density, Bb, resulting from the interaction of MMFb describe in equation (9) with the dynamic eccentricity component permeance generated by bearing fault, cos(O - ort) [2], it is given by equation (10)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001657_bfb0119410-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001657_bfb0119410-Figure6-1.png", "caption": "Figure 6: Reconfigurable Rocker Concept", "texts": [ " An SMA-actuated reconfigurable rocker-bogie suspension concept has been developed which allows the rover to modify its geometry to improve mobility and avoid failure situations. The mechanism allows the rover to squat one or both sides of its suspension, and thus increase its stability margin when required (for example, when the rover is on a transverse slope). Additionally, reconfigurability allows the rover to reposition its center of mass when performing traction control [5,6]. An illustration of a reconfigurable rocker mechanism is shown in Figure 6. A Flexinol SMA wire provides retracting force for the rocker, and a multi-jaw coupling locks the rocker links in place, fixing the rocker angle. The SMA wire is routed over Delrin wire guides to increase the working length. 3. Electronic and Power System The rover electronics system was designed to be compact, low-cost and lightweight (see Figure 7). A block diagram of the system can be seen in Figure 8. The system is based on a PC/104 486 computing platform, with additional modules for digital and analog IO, encoder reading, and interface to a JR3 six-axis force/torque sensor which is mounted underneath the manipulator (see Section 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002912_j.powtec.2009.07.004-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002912_j.powtec.2009.07.004-Figure1-1.png", "caption": "Fig. 1. Alignment of the ellipsoid.", "texts": [ " A model of the body's stiffness is required; this will vary with orientation, and this variation necessarily involves a torque. If the contact point stops sliding, tangential force is still required, because the centre of mass changes direction during rolling; if the coefficient of friction is insufficient, sliding can restart. As a final complication, the motion of the ellipsoid after a collision can lead to further collisions. The model for the sphere in Section 2 is extended here to the ellipsoid, using coordinates shown in Fig. 1. The model still neglects tangential distortion. Since the duration of the collision is important, the model must specify the normal force as a function of the compression distance. This was not necessary for the sphere in Section 2. In this work, the simplest model is used: the normal force is that of a perfectly elastic body. Assuming small deformations, elliptical contacts were considered by Hertz, and the model is well established [23]. In this work, it is stated as: FN = 4 3 GE4 ffiffiffiffiffi Re p \u03b43=2 \u00f08\u00de In this equation, the overall effective radius of curvature at the contact point, Re, is deduced from the dimensions and orientation of the ellipsoid; the dimensionless factor G is also a complicated function of orientation \u2014 see Appendix A for details", " (8) I Second moment of inertia about centre of gravity J Impulse vector K Gaussian curvature k Aspect ratio of ellipsoid, (b/a) L ln(\u03bb) m Mass n \u0302 Unit vector normal to plane, pointing into plane p, q Position of contact point relative to centre of gravity in axisaligned coordinates Q Rolling torque R Radius t Time U Velocity of contact point V Velocity of centre of gravity W Work done against elastic forces x Coordinate of centre of gravity y Coordinate of contact point Subscripts and superscripts 0 Dimensionless (0) Initial N Normal (n) Corresponding to n-th timestep T Tangential (\u221e) Corresponding to state after all impacts Greek letters \u03b1 Angle of orientation \u03b2 Angle of translational direction (from normal) \u03b4 Displacement of centre of gravity \u03b8 Parameter of position on ellipse \u03bb Ratio of principal radii of curvature of body at contact point v Poisson's ratio \u03bc Coefficient of friction \u03c1 Density \u03c4 Time \u03c9 Speed of angular rotation This appendix presents the detailed equations used in the current case, including a description of the \u2018rolling torque\u2019. As stated in the text, this case is the impact of a perfectly elastic ellipsoid onto a semiinfinite plane surface, where the motion of the ellipsoid is in one of its planes of symmetry, perpendicular to the surface. Standard equations [27] are used for the geometry and curvature of the ellipsoid. The first task is to deduce the position and attributes of the contact point from the coordinates of the ellipsoid. The coordinates are defined in Fig. 1. The orientation of the ellipsoid is defined by \u03b1, the angle anticlockwise from horizontal to the a-axis of the ellipsoid. The position of the centre of gravity is given by xN and xT, relative to a fixed position on the surface of the plane. The subscripts denote normal and tangential components respectively, with the normal component upwards. The position of the contact point (defined as the relevant tangent of the ellipsoid) is given by yN and yT, relative to the centre of gravity. The normal component is here downwards, so that both xN and yN are always positive" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000057_s0069-8040(08)70029-3-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000057_s0069-8040(08)70029-3-Figure9-1.png", "caption": "Fig. 9. Typical wall-jet cell, A, Disc electrode contact; B, ring electrode contact; C, Ag/AgCl reference electrode; D, Pt tube counter electrode; E, cell inlet; F, Kel-F cell body. (From ref. 44.)", "texts": [ " A note of caution is that, in this conformation, air can enter the solution relatively easily so that junctions must be well sealed. The arrangement of the flow systems is very similar to that employed in flow-injection analysis [ 135,1361. For use with tubular and channel electrodes (Fig. 8), no extra special 395 References pp. 434-441 difficulties arise so Iong as, if they are of the sandwich type, there are no leaks, especially round 0 ring joints. The same problem may occur with wall-jet cells (Fig. 9). If there are any small leaks caused by incorrectly fitting 0 rings, then this can usually be solved by a piece of Teflon tape on the screw thread. Additionally, for wall-jet cells, care needs to be exercised in filling the cell with solution so that trapping of air bubbles is avoided. The basic construction of a DME is as shown in Fig. 10. Mercury flows from a reservoir at height h above the top of a glass capillary, through the capillary and into the test solution. The connecting tubing would ideally be glass, but this is impractical and unnecessary: Tygon, Teflon, and various other plastics have been employed" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001127_s10015-005-0359-3-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001127_s10015-005-0359-3-Figure2-1.png", "caption": "Fig. 2. n-Link underwater robot model", "texts": [ " Therefore we propose a new discrete-time RAC method whose feedback gains can be set individually.12 Here, in the case of a singular configuration of manipulators, the control input generally cannot be calculated. In this paper, we propose a discrete-time RAC method considering a singular configuration of the manipulator. To avoid a singular configuration, the desired position of the vehicle is modified based on the determinant of the manipulator\u2019s Jacobian matrix. The effectiveness of the proposed method is shown by an experiment. 2 Modeling The UVMS model used in this paper is shown in Fig. 2. It has a robot base (vehicle) and an n-DOF manipulator. 2.1 Kinematics In this subsection, based on Yoshida and Umetani,13 kinematic and momentum equations are derived. First, from Fig. 2, a time derivative of the end-tip position vector pe is v v p r k p pe e i e i i i n = + \u2212( ) + \u2212( ){ } = \u22110 0 1 \u02dc \u02dc \u02d9w \u03c6 (1) On the other hand, the relationship between the end-tip angular velocity and the joint velocity is expressed by w w fe i i i n = + = \u22110 1 k \u02d9 (2) From Eqs. 1 and 2 the following equation is obtained: \u02d9 \u02d9 \u02d9c c fe = +A B0 (3) where A E p p E B b b b= \u2212 \u2212( ) = [ ]3 0 3 1 2 0 \u02dc \u02dc . . .r n, and bi = [{k\u0303i(pr \u2212 pi)}T, kT i ]T. Next, let h and m be the linear and angular momentum of the robot, including the hydrodynamic added mass tensor Mai and the added inertia tensor Iai of link i" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001782_j.molcatb.2007.08.005-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001782_j.molcatb.2007.08.005-Figure1-1.png", "caption": "Fig. 1. Principle of operation.", "texts": [ " Assembly of homocysteine biosensor and etermination l-Homocysteine desulfhydrase immobilized eggshell memrane was positioned on the surface of a Thermo-Orion mmonium selective electrode and kept in a steady position y an O-ring. Then, ammonium selective electrode and refernce electrode (Orion, model 90-02) were immersed together nto a stirred reaction media containing phosphate buffer soluion (10 mL). Standard dl-homocysteine or sample solution as injected into the reaction media and ammonium concenration signal was measured and processed by a Thermo-Orion 50A + pH/ionmeter (Fig. 1). . Results and discussion .1. Measuring procedure The ammonium selective electrode acting as an ammonium ransducer was employed to measure the rate of ammonium ormation in the enzymatic consumption of homocysteine. The nalytical signal of the homocysteine biosensor is the increase in he millivolts (mV) upon exposure to a homocysteine solution. he enzymatic oxidation of homocysteine by l-homocysteine lar Catalysis B: Enzymatic 49 (2007) 55\u201360 57 d t r n M a 3 3 d d o s b e w w 3 o p s w s o s m l t c F p F p m h c a 1 3 fi o s h b a t o S" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002534_s00285-008-0227-6-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002534_s00285-008-0227-6-Figure1-1.png", "caption": "Fig. 1 Sketch of two straight, infinitely long filaments and \u0303 with intersection angle \u03d5. At each binding site of with distance s from the intersection point there can be at most two cross-link connections of length \u03c1 to the other filament, with binding angles \u03b1 or \u03b1\u0302 at , and corresponding angles \u03b1\u0303 at \u0303", "texts": [ " Let us consider two (infinitely long) oriented stiff filaments and \u0303 in R 2 pointing into directions \u03b8 and \u03b8\u0303 \u2208 S 1, with uniquely defined intersection angle \u03d5 = (\u03b8\u0303 , \u03b8) satisfying cos\u03d5 = \u03b8 \u00b7 \u03b8\u0303 and sin \u03d5 = \u03b8 \u00b7 \u03b8\u0303\u22a5. Here we define the orientation of \u03b8\u22a5 by the convention (\u03b8, \u03b8\u22a5) = \u03c0 2 . As canonical arc length parameters let us choose the signed distances s and s\u0303 from the intersection point, so that any pair of positions (potential cross-linker binding sites) on the filaments is given by the points s\u03b8 and s\u0303\u03b8\u0303 , with the contact vector z = s\u03b8 \u2212 s\u0303\u03b8\u0303 pointing from filament \u0303 towards filament , see Fig. 1. Since for actin filaments with typical lengths L 1 \u00b5m the binding sites for myosin are regularly spaced by 2.7 nm (cf. [16]), for example, we propose that in a justified continuum limit the binding sites for cross-linkers are continuously and uniformly distributed along both filaments and that different cross-linkers do not conflict with each other, so that we can take the simple product measure \u00b5(s\u0303, s) = ds\u0303 \u00b7 ds in the energy integral (2). Furthermore, assuming a quasi-stationary situation for given intersection angle 0 < \u03d5 < \u03c0 , we propose that binding probability and averaged dynamical status, i.e. the energetics of any doubly bound cross-linker only depend on the geometric configuration of its contact vector z with respect to the two filaments. More specifically, for stiff cross-linkers that cannot bend but could be elongated by thermal fluctuations to a binding length \u03c1 = |z| \u2265 d > 0, the thermodynamic energy is assumed to depend only on the distance \u03c1 and the two binding angles \u03b1 = (\u03b8, z) and \u03b1\u0303 = (z,\u2212\u03b8\u0303 ), see Fig. 1. Notice that each binding angle \u03b1 at filament has a uniquely determined dual angle \u03b1\u0303 = \u03c0 \u2212 \u03d5 \u2212 \u03b1 at the other filament with the transformation properties s\u0303 = \u03c1 sin \u03b1 sin \u03d5 , (3) s = \u03c1 sin \u03b1\u0303 sin \u03d5 = \u03c1 sin(\u03b1 + \u03d5) sin \u03d5 . (4) Therefore, we can reparametrize the binding position coordinates (s, s\u0303) \u2208 \u00d7\u0303 by the coordinates (\u03c1, \u03b1) \u2208 [d,\u221e)\u00d7 (0, 2\u03c0) of binding length and angle with respect to one filament, here . Because the Jacobian of the transformation (\u03c1, \u03b1) \u2192 (s, s\u0303) simply equals J (\u03c1, \u03d5) = \u03c1/ sin \u03d5 > 0, this constitutes a diffeomorphism of [d,\u221e)\u00d7[0, 2\u03c0) onto the closed domain d = \u222a\u03c1\u2265dC\u03c1 \u2282 R 2, where the ellipse C\u03c1 = {(s, s\u0303) : |s\u03b8 \u2212 s\u0303\u03b8\u0303 |2 = 1 2 (1 \u2212 cos\u03d5)(s + s\u0303)2 + 1 2 (1 + cos\u03d5)(s \u2212 s\u0303)2 = \u03c12} is 2\u03c0 -periodically parametrized by the second argument \u03b1, see Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002090_9780470264003-Figure20.1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002090_9780470264003-Figure20.1-1.png", "caption": "Figure 20.1 Auto 3D schematic.", "texts": [ " Again the DPs are the \u201cwhats,\u201d and we decompose them into the \u201chows,\u201d the process variables (PVs). The automatic dissolving and dosing device (Auto 3D) is a device that comprises a system of hardware, disposable materials, and software that is used by a nurse, pharmacist, and/or caregiver at home to automatically reconstitute a prescribed dose of medication to be administered in one or multiple doses. The system goal is to consistently, safely, and effectively dissolve and dose the required pharmaceutical. Auto 3D is depicted in Figure 20.1. Early on, the design team realized that the actual customer base would include the corporate parent, looking for results; a local management team that wanted no increased cost; and the operational caregiver, who needed the device to service patients. The latter is closest to the end user, the patient. In this segment, the customer needs for the Auto 3D device were defined based on marketing research (through interviews laced with common sense), and the following list of needs was obtained: \u2022 Safety and effectiveness \u2022 Medication security \u2022 Device ease of use \u2022 Device reliability and availability \u2022 Device value cost \u2022 Device visual appearance The high-level wants were prioritized as shown in Table 8", " El-Haik and Khalid S. Mekki Copyright \u00a9 2008 John Wiley & Sons, Inc. In this chapter we present the development of an automatic dissolving and dosing device (Auto 3D), described briefly as a system of hardware, disposables, and software used by a nurse, pharmacist, and/or a caregiver at home to automatically dispense a prescribed dose of medication to be administered in one or multiple doses. The system goal is to dissolve and dose the required pharmaceutical consistently, safely, and effectively. Figure 20.1 is a schematic of the Auto 3D. Early on, the design team segmented the external customer base into the following groups: The pharmacists and nurses are the closest to the patient. In this segment, the customer needs for the Auto 3D device were defined based on marketing research (through interviews laced with common sense), and the following list of needs was obtained: \u2022 Safety and effectiveness \u2022 Medication security \u2022 Device ease of use \u2022 Device reliability and availability \u2022 Device value cost \u2022 Device visual appearance The high-level wants were prioritized using the analytical hierarchy process1 (AHP) shown in Figure 20", " 9 A ct ua to r A ss em bl y P F M E A ( Sh ow n P ar ti al ly ) It em / F un ct io n Po te nt ia l Fa ilu re M od e Po te nt ia l E ff ec t( s) o f Fa ilu re SE V C la ss Po te nt ia l C au se s( s) o r M ec ha ni sm (s ) of F ai lu re O C C C ur re nt D es ig n C on tr ol s P re ve nt io n an d D et ec ti on D E T R P N R ec om m en de d A ct io ns R es po ns ib ili ty an d Ta rg et C om pl et io n D at e 1 A ct ua to r as se m bl y: ac tu at or ex te nd s an d re tr ac ts a pr ec is e di st an ce D oe s no t ex te nd or re tr ac t D is po sa bl e ki t lo ss 8 SC D ef ec ti ve ac tu at or 3 Im pl em en t A P Q P pr oc es s on ac tu at or m an uf ac tu re ; de si gn /m fg re vi ew s 5 12 0 D efi ne a s a cr it ic al pr od uc t: ch ar ac te ri st ic s in dr aw in gs a nd k ey pr od uc t ch ar ac te ri st ic s sh ee t, pr ot ot yp e co nt ro l pl an c he ck , i ni ti at e as se m bl y D V P & R , in it ia te p ro ce ss F M E A a nd c on tr ol pl an , c on ti nu e A P Q P p ro ce ss , co nt in ue d es ig n/ m fg re vi ew s Te am a nd pl an t 1 m fg 9/ 8/ 07 P ro ce du re t im e de la ys (o pe ra ti on ) D ef ec ti ve di st ri bu ti on P C B 3 Im pl em en t A P Q P pr oc es s on P C B m an uf ac tu re ; de si gn /m fg re vi ew s 5 12 0 D efi ne a s a cr it ic al pr od uc t: ch ar ac te ri st ic s in dr aw in gs a nd ke y pr od uc t ch ar ac te ri st ic s sh ee t, pr ot ot yp e co nt ro l pl an c he ck , i ni ti at e as se m bl y D V P & R , in it ia te p ro ce ss F E M A a nd c on tr ol pl an , c on ti nu e A P Q P p ro ce ss , co nt in ue d es ig n/ m fg re vi ew s Te am a nd pl an t 1 m fg 9/ 8/ 07 484 485 It em / F un ct io n Po te nt ia l Fa ilu re M od e Po te nt ia l E ff ec t( s) o f Fa ilu re SE V C la ss Po te nt ia l C au se s( s) o r M ec ha ni sm (s ) of F ai lu re O C C C ur re nt D es ig n C on tr ol s P re ve nt io n an d D et ec ti on D E T R P N R ec om m en de d A ct io ns R es po ns ib ili ty an d Ta rg et C om pl et io n D at e R ed uc ti on in P I R B C s ui te pr od uc ti on th ro ug hp ut D ef ec ti ve pr ox im it y se ns or 3 D es ig n/ m fg re vi ew s 6 14 4 D efi ne a s a cr it ic al pr od uc t: ch ar ac te ri st ic s in dr aw in gs a nd ke y pr od uc t ch ar ac te ri st ic s sh ee t, im pl em en t A P Q P pr oc es s on s en so rs m an uf ac tu re r, pr ot ot yp e co nt ro l pl an c he ck , i ni ti at e as se m bl y D V P & R , in it ia te p ro ce ss F M E A a nd c on tr ol pl an , c on ti nu e A P Q P p ro ce ss , co nt in ue d es ig n/ m fg re vi ew s Te am a nd pl an t 1 m fg 9/ 8/ 07 C us to m er di ss at is fa ct io n D ef ec ti ve pl un ge r bu sh in g 2 D es ig n/ m fg re vi ew s 6 96 D efi ne a s a cr it ic al pr od uc t: ch ar ac te ri st ic s in dr aw in gs a nd ke y pr od uc t ch ar ac te ri st ic s sh ee t, pr ot ot yp e co nt ro l pl an c he ck , i ni ti at e A ss em bl y D V P & R , in it ia te p ro ce ss F E M A a nd c on tr ol pl an , c on ti nu e de si gn /m fg r ev ie w s Te am a nd pl an t 1 m fg 9/ 8/ 07 486 It em / F un ct io n Po te nt ia l Fa ilu re M od e Po te nt ia l E ff ec t( s) o f Fa ilu re SE V C la ss Po te nt ia l C au se s( s) o r M ec ha ni sm (s ) of F ai lu re O C C C ur re nt D es ig n C on tr ol s P re ve nt io n an d D et ec ti on D E T R P N R ec om m en de d A ct io ns R es po ns ib ili ty an d Ta rg et C om pl et io n D at e R eq ui re s fie ld se rv ic e D ef ec ti ve U -c up s ea l 3 D es ig n/ m fg re vi ew s 6 14 4 D efi ne a s a cr it ic al pr od uc t: ch ar ac te ri st ic s in dr aw in gs a nd ke y pr od uc t ch ar ac te ri st ic s sh ee t, pr ot ot yp e co nt ro l pl an c he ck , i ni ti at e as se m bl y D V P & R , in it ia te p ro ce ss F M E A a nd c on tr ol pl an , c on ti nu e de si gn /m fg r ev ie w s Te am a nd pl an t 1 m fg 9/ 8/ 07 TA B L E 2 0. 9 C on ti nu ed In this chapter we show an application of DFSS on a medical device, Auto 3D, depicted in Figure 20.1. The intention here is to provide a high-level understanding of the project rather than to document every step and tool application. SUMMARY 487 488 Medical Device Design for Six Sigma: A Road Map for Safety and Effectiveness, By Basem S. El-Haik and Khalid S. Mekki Copyright \u00a9 2008 John Wiley & Sons, Inc. 2k full factorial a specific type of design of experiments where two or more (k number of) factors, each with two levels, with all possible combinations between the factor levels, are studied" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003842_iros.2011.6095093-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003842_iros.2011.6095093-Figure1-1.png", "caption": "Fig. 1. Robot architecture and notations.", "texts": [ " Decentralized formulations for trajectory generation appear in [11] where the authors present a decentralized approach for multi-robot motion planning using potential field controllers based on navigation functions, [12] where the authors present a a decentralized nonlinear model predictive control (NMPC) formulation for trajectory generation of helicopters that includes a potential function that reflects state information of other moving vehicles into the cost function of the individual vehicle, and [13] where a decentralized algorithm for multiple aircraft coordination based on an iterative bargaining scheme is presented. Most closely related to our work is [14], where the authors present a decentralized receding horizon formulation for multiple aircraft path planning using mixed integer linear programming (MILP) to generate provably safe trajectories. Similar to [14], we use a sequential decision ordering mechanism as a part of our solution algorithm. Consider a two wheeled differential drive mobile robot shown in Fig. 1. The robot moves in a global (X, Y) Cartesian co-ordinate plane and is represented by the following kinematic model with associated non-holonomic constraint (that disallows the robot from sliding sideways). x\u0307 = scos(\u03b8); y\u0307 = ssin(\u03b8); \u03b8\u0307 = \u03c9 (1) x\u0307sin(\u03b8)\u2212 y\u0307cos(\u03b8) = 0. (2) Here s and \u03c9 are the linear and angular speeds of the robot respectively; x, y and \u03b8 are the coordinates of the robot with respect to the global (X, Y) coordinate system. Consider a group of n such mobile robots. Each robot i = 1,2," ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000524_gt2004-53297-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000524_gt2004-53297-Figure1-1.png", "caption": "Fig 1b Strip Seal Configuration (detail)", "texts": [ " Strip seals are easy to fabricate and install, are relatively inexpensive, and unlike machined labyrinth teeth, are easy to replace after many years of service. The most common configuration for the design consists of a thin length of strip steel, with a 90 degree bend at the base (forming the J profile), which is then formed around either the rotating or stationary part so that the sharp protruding tip acts as a knife-edge type seal. The seal is inserted into an accepting groove in the host part, and held into place by a retaining or caulking wire (Fig. 1a and 1b). 1 Copyright \u00a9 2004 by ASME rl=/data/conferences/gt2004/71154/ on 05/22/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use D Strip seal narrow profile and relatively small axial space requirements make them a popular choice in most reaction-type turbine designs. Moreover, their relative low cost and ease of installation and replacement make them particularly well suited for use on rotating components. One inherent deficiency of the strip seal design is its limited ability to survive a rub event", " Depending on the seal s location in the turbine and its distance from the thrust bearing, the relative axial movement between the seal strip and the adjacent sealing surface can exceed 12.5mm. In addition, machine transients such as startup and shutdown can result in radial movements that limit the minimum possible radial seal clearance. In the past, seal strips have been applied in locations radially opposite solid metal sealing surfaces (such as nozzle covers or bucket covers), as 2 Copyright \u00a9 2004 by ASME =/data/conferences/gt2004/71154/ on 05/22/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use D shown in Figure 1. However, use of abradable material on the sealing surfaces opposite the seal strips has allowed machines to be designed with tighter radial clearances, since the seal strips can cut into the abradable without incurring damage. At assembly, seal clearances are deliberately set such that the strips lightly rub the abradable material during the initial startup, forming small local trenches in the coating. The result is that the seals run with the minimum achievable radial clearances, thereby reducing secondary leakage in the turbine and improving overall turbine efficiency" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001893_s00170-008-1850-5-Figure23-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001893_s00170-008-1850-5-Figure23-1.png", "caption": "Fig. 23 Movement tool/workpiece elliptical trajectory", "texts": [ " Because of part elasticity intervening in the operation of turning, it is logical that the response in displacements of units BT\u2013BW is carried out with a certain shift compared to the efforts variation applied to the tool, variation induced by the lack of machined surface circularity in the preceding turn, which implies variations of the contact tool/workpiece (Fig. 17). Phase difference between the efforts and displacements thus remains a possible explanation to the self-excited vibrations appearance. Moreover, when the tool moves along the ellipse places e1, e2, e3 (Fig. 23), the cutting force carries out a positive work because its direction coincides with the cutting direction. On the other hand, on the side e3, e4, e1, the work produced by the cutting force is negative, since its direction is directly opposed to that of displacement. The comparison between these two ellipse parts shows that the effort on the trajectory e1, e2, e3 is higher than on the trajectory e3, e4, e1 because the cutting depth is more important. At the time of this process, work corresponding to an ellipse trajectory remains positive, and the increase of result energy thus contributes to maintaining the vibrations and to dissipating the energy in the form of heat by the assembly tool/workpiece" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000299_j.actaastro.2004.02.004-Figure13-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000299_j.actaastro.2004.02.004-Figure13-1.png", "caption": "Fig. 13. Scaled satellite dynamics model.", "texts": [ " Hardware experiments involving closed loop dynamics are carried out with a novel facility that displays system dynamics equivalent to the on-orbit dynamics of HALCA. This dynamic closed loop test (DCLT) facility accounts for satellite dynamics and structural 3exibility, and incorporates actual satellite hardware and software [7]. Fig. 12 shows the simpli7ed DCLT con7guration. In the DCLT, the attitude angle of the scale satellite dynamics model is measured by the tracking camera, instead of STT mounted on HALCA. The measurement signal by the camera is also used for the Gyro noise calibration. The scaled satellite dynamics model is shown in Fig. 13. It is developed to simulate single-axis dynamics of a satellite with 3exible structures [7]. In the actual attitude and orbit control subsystem (AOCS), the performance of the on-board computer is restricted, and the proposed control in Eq. (13) therefore could not be implemented fully. Because the sampling frequency (period of attitude measurements and calculation of control commands) is 8 Hz (125 msec), e8ective suppression of 3exible structure vibration (even for the solar array 7rst-mode) cannot be expected" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000699_jo00320a021-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000699_jo00320a021-Figure1-1.png", "caption": "Figure 1. pH-rate profiles for the hydrolysis of la in the absence (0) and in the presence of borate buffer (0).", "texts": [ " The final spectra of reaction mixtures coincide well with those of expected hydrolysis products (acid + thiol) except for mea where a nucleophile-like hydrazine was At higher pH (>lo), oxidation of thiol formed became significant to affect the infinite absorbance readings; therefore, a modified Guggenheim treatment14 of data was employed. Hydrolysis rates of hydroxy thiolesters 1 and 2, as well as a simple thiolester 3, increased linearly with [OH-] a t pH >9 and showed little buffer dependence in tertiary amine and carbonate buffers (not shown). However, the reactions of la were markedly accelerated in borate buffers. The hydrolysis rate of la in a dilute borate buffer, whose total concentration [B], = 0.04 M, buffer ratio = 1, and pH 9.05, is 78 times greater than that in the absence of borate buffer a t the same pH. Figure 1 shows pH-rate profiles of la in the absence and presence of borate buffer. Similar accelerations were found also for the other hydroxy thiolesters, lb, IC, and 2, but not for a simple thiolester, 3. The kinetic results are summarized in Table I. (11) Bruice, T. C.; Bruno, J. J.; Chou, W.-S. J. Am. Chem. SOC. 1963, (12) Fedor, L. R.; Bruice, T. C. J. Am. Chem. SOC. 1964,86,4117-4123. (13) Bruice, T. C.; Benkovic, S. \"Bioorganic Mechanisms\"; Benjamin. (14) Swinbourne, E. S. J. Chem. SOC. 1960, 2371-2372", " From these linearities (eq 31, the kinetic parameters, k,, and Ka, were calculated and are given in Table 11. To examine catalytic effects of boric acid in lower pH region, a second buffer was used to attain constant pH while keeping [B], = 0.04 M. Among the second buffers added, morpholine and N-methylmorpholine showed little buffer effect, whereas hydrazine, imidazole, and ammonia exhibited significant effects (Figure 4). The rate constants koW extrapolated to zero added buffer concentration while keeping [B], = 0.04 M are plotted against pH for la in Figure 1. The solid curve of Figure 1 is a theoretical one calculated with the parameters k, and K , obtained above. In order to characterize the effects of a second buffer found in the borate-catalyzed hydrolysis of la, the effect of hydrazine was examined in detail. The rates increased sharply with hydrazine concentration (Figure 4b) as found previously for the hydrazinolysis of thiolacetates and thio lactone^.^'-'^ Without added borate, eq 4 holds, as (4) illustrated in Figure 5, where [Hy], = [Hy] + [HyH+l and Hy and HyH+ stand for neutral and protonated hydrazine, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001327_s0019-9958(79)90088-3-Figure15-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001327_s0019-9958(79)90088-3-Figure15-1.png", "caption": "Fig. 15. Now C preserves all edges at each derivation step and moreover adds the", "texts": [], "surrounding_texts": [ "Consider G over (a, b, c, d}, {g, h) where S, P and C are given in Fig. 13. L(G) consists of the initial graph and the graphs of Fig. 14. Clearly G is not a PG00L system. Consider any PGOL system G' such that L(G') ~ L(G). Now the initial graph of G' must be a o_~ b, since this graph has the fewest nodes. Since we are dealing with propagating systems a and b each must give rise to at least one node, and in fact, one of them must give rise to two nodes only, and the other only one. Clearly, the productions for a and b in G can be interchanged and the connection rules modified appropriately to give L(G) once more in a nonPG00L fashion. Hence the only other possibility is that a (or b) gives either c--~c or d, and b (or a) gives -~ c --~. However, the connection rules must differentiate between the two possible daughter graphs of a to ensure that the appropriate connections are made. Hence this system is also non-PG00L. Therefore L(G) is a non-PG00L language.\nLet us now restrict our attention to graphs with a single-edge-label, in other words to the webs of Rosenfeld and Milgram (1972). Cheung (1979) has claimed that any bounded degree PGOL system can be \"simulated\" by a single-edge-", "labeled PGOL system, in the sense that the underlying graph structures obtained are the same. By underlying graph structure we mean the coding of the graph language which identifies all edge labels and all node labelS, that is, an unlabeled graph, tIowever, for arbitrary PGOL systems this \"simulation\" result does not hold.\nOn the other hand if we restrict attention to single-node-labeled graphs and PGOL systems the generative power of such systems is drastically reduced; in particular, the star graph language of Section 3 cannot be generated when only a single node label is allowed. Therefore the trade-off of edge labels and node labels is in one direction only. For bounded degree systems the edge label set can be reduced at the expense of the node label set.\nWe now give two examples which generate all complete graphs over Z, A and all graphs over Z', A.\nappropriate new edges to maintain completeness. For each complete graph it is straightforward to construct a derivation in G from S which generates it. Assume G generates at least one incomplete graph. Assume U is an incomplete graph in L(G) such that there is no incomplete graph V in L(G) with either fewer nodes or the same number of nodes and more edges. Consider all T such that T ~ U in G and T ,# U. Now T is complete, since either T has fewer nodes than U or T", "has the same number of nodes as U but more edges (fewer edges is impossible since edges cannot be created with C). But if T is complete then U cannot be incomplete by examination of C. Hence all U in L(G) are complete. This can be proved rigorously by induction on the number of nodes.\nWe now turn to a system generating all graphs. In the completeness example onlv one kind of connection was necessary, but in the following system all possible connections, including no connections, must be specified.\nLet G = (Z', A, p, C, S), where A = {h}, and S, P, and C are shown in\nWe close this section with two further examples. The first demonstrates a self-reproducing system, that is, one which generates multiple copies of the start graph.\nLet S in [22, A] ~ be the start graph.\nP contains the production given in Fig. 17a for all a in \u00a3; that is, each node duplicates itself with all possible hands over A (the labels arc not shown in the figure) and the hands shown for h, h not in A.\nC contains the two connection rules of Fig. 17b for all g in A and all a, b in 22.\nAt each derivation step two new copies of S are made from each old copy of S and each new copy is reconnected correctly by use of the extra hands labeled h in the productions. Since S does not contain a hand labeled h neither do any of its offspring.\nOur second and final example provides a simple PGOL system that generates the so-called spinning spider spiral. This was produced in response to challenge of Mayoh (1978), who provided the problem. This is a nontrivial example demonstrating the power of PGOL systems, since sequential context-free graph grammars are incapable of generating it and sequential non-context-free graph grammars for this problem are quite complex." ] }, { "image_filename": "designv11_20_0003710_s12541-011-0060-5-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003710_s12541-011-0060-5-Figure2-1.png", "caption": "Fig. 2 AmTec arm", "texts": [ " In this section, the comparison study is demonstrated in order to show that the proposed method is efficient than DLS method,8 MIN solution and RMA one,19 which are well known as traditional methods providing a good solution in which a manipulator is near the vicinity of the singularities or path through the position. The proposed method is applied to a 7-DOF redundant robot, ultra light-weight AmTec, which has kinematic structure like human arm. Its Denavit Hartenberg is given in Table 18 and the coordinate system is shown in Fig. 2. The maximum end effector translational velocities and rotational velocities are set at value of 0.28 [m] and 0.2793 [rad/s], respectively. Maximum allowable velocity in this redundant case is set at 0.6 [rad/s]. The maximum value of the damping factor, max ,\u03bb and the singular region, ,\u03b5 in Fig. 1 are same as those in a previous paper.8 The initial configuration and final one in this simulation are [0, 0.349, 0, 1.396, 0, 0.96, 0] initial T = \u2212q and = [0, /6, 0,finalq \u03c0\u2212 /2, 0, 0.61, 0] , T \u03c0 \u2212 respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003140_iros.2009.5354270-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003140_iros.2009.5354270-Figure5-1.png", "caption": "Fig. 5. State space of the giant-swing robot.", "texts": [], "surrounding_texts": [ "A. Experimental Condition and Method In order to verify whether the developed robot can achieve the giant-swing motion or not, a preliminary experiment was conducted by actually using the real robot; the dynamic simulator was not employed in this preliminary test. In this preliminary experiment, we allowed each enabled actuator to drive by 9 action patterns\u20140, \u00b110, \u00b120, \u00b130, and \u00b140 deg\u2014as shown in Fig. 6. Further, the reward based on the decrease in the height of a LED target attached on the arm was allocated for all the actions in each state; we supposed that the decrease in the swing height includes the potential energy, which may be effective for enhancing the swing motion. As for the \u03b5-greedy method, \u03b5 was decreased by 0.2 every 10000 learning steps from 1.0, i.e. it means that the frequency of greedy actions increases every 10000 learning steps. Hence, the leaning was finished when \u03b5 became 0. B. Experimental Results and Discussion Fig. 7 shows the transition in the swing angle for 60 s when using the action-value functions renewed in 20000 learning steps. In this graph, the swing angle over \u00b1180 deg means that the robot could rotate around the horizontal bar. As shown in Fig. 7, it should be noted that the robot could not perform the giant-swing motion. Fig. 8 shows a distribution of the action-value functions obtained in this learning, where the action-value functions are categorized by five levels. The result shows that the action-value functions in the bottom of the horizontal bar are relatively low and have no remarkable difference in the neighborhood of state 72. In most cases, the robot was easy to fall into a motion loop with few action patterns in the early stage of the greedy action. This motion loop caused the stagnation of the swing at the bottom of the horizontal bar. The stagnant state prevented the robot from shifting to a new state for increasing the swing angle. When the robot accidentally shifted to the new state from the stagnant states, there were cases where the robot was able to achieve the giant-swing motion. Actually, the robot could jump out of the stagnant state if we applied small forced oscillation due to the external force in the early stage. Fig. 9 shows the transition in the swing angle when applying the random actions for 10 s instead of the forced oscillation. In this experiment, the robot could select the best actions based on the learning result, which was the same with that in Fig. 7, after the robot performed the random actions for 10 s. The results demonstrate that the robot could rotate around the horizontal bar with some repeatability, as shown in Fig. 9; in this method, the giant-swing motion was almost regularly performed between 30 s and 70 s. Fig. 10 graphically shows the highlight. These results imply that the reward based on the decrease in the swing height was not so effective for generating the giant-swing motion; how to give the reward may break the problem related to the repeatability. However, it could be confirmed that the developed giant-swing robot has the ability to perform the giant-swing motion." ] }, { "image_filename": "designv11_20_0000980_1-4020-3559-4-Figure20-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000980_1-4020-3559-4-Figure20-1.png", "caption": "Figure 20. Localized deformations on a beam and a plastic hinge.", "texts": [ " Many applications of multibody dynamics require the description of the flexibility of its components. For structural crashworthiness it is La te ra l F la ng e F or c e [ N ] often unfeasible to use large nonlinear finite element models. The use of multibody dynamics with plastic hinges is an alternative formulation that allows building insightful models for crashworthiness. In many impact situations, the individual structural members are overloaded giving rise to plastic deformations in highly localized regions, called plastic hinges. These deformations, presented in Fig. 20 for structural bending, develop at points where maximum bending moments occur, load application points, joints or locally weak areas [19]. Multibody models obtained with this method are relatively simple, which makes the procedure adequate for the early phases of vehicle design. The methodology described herein is known in industry as conceptual modeling [20]. The plastic hinge concept has been developed by using generalized spring elements to represent constitutive characteristics of localized plastic deformation of beams and kinematic joints to control the deformation kinematics [21], as illustrated in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003845_1.4003270-Figure12-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003845_1.4003270-Figure12-1.png", "caption": "Fig. 12 Planar metamorphic mechanism: state 3", "texts": [], "surrounding_texts": [ "p i\nL t\nT l j\nw R t l m o T\n0\nDownloaded Fr\nIf this mechanism state matrix were input into the computer rogram, the augmented mechanism state matrix would be given\n1\n1 1 4\n2 3\n5\n6\n87\nY\nX\nFig. 7 3RRR: state 2\n1\n1 1 4\n2 3\n5\n6\n8 7\nY\nX\nFig. 8 3RRR: state 3\nn Eq. 21 as\nhe mechanism state matrix is given in Eq. 24 as\n11012-6 / Vol. 3, FEBRUARY 2011\nom: http://mechanismsrobotics.asmedigitalcollection.asme.org/ on 02/03/2\nMASM = 2Z R,6Z R,8Z R 3Z R 4Z R 5Z R,7Z R 6Z R 8Z R 3 2Z R,6Z R,8V X 3Z R 4Z R 5Z R,7Z R 6Z R 8Z R 2\n2Z R,6Z R,8V X 3Z R 4Z R 5Z R,7Z R 6Z R 8V X 1\nThe augmented mechanism state matrix identifies how the DOF of this mechanism changes as it moves from one state to another. Additionally, it is easy to track how a link changes as the mechanism changes states. For instance, link 8 changes from a movable link in state 1 to a fixed link in states 2 and 3.\n6.2 Nine-Bar Linkage. Mechanism state matrices can be particularly useful when analyzing complex mechanisms in which the DOF of the mechanism is not obvious upon inspection. Such a mechanism is the nine-bar linkage in Fig. 9. In this example only one state is considered, and the mechanism state matrix is given in\n1 2\n3 9 8 5\n4 6\n7\nY\nX\nFig. 9 Nine-bar linkage\n1 1\na d b\n2\n4 53 P2\nP1\nc\n2\nY\nX\nEq. 22\nMSM = 2Z R,8Z R,9Z R 3Z R,5Z R 4Z R 5Z R,6Z R 7Z R 8Z R,9Z R 22\nink 1 is specified to be a fixed link. It is unknown which links and joints, if any, become fixed as a result of link 1 being chosen as he fixed link. The output augmented mechanism state matrix is given in Eq. 23 as\nMASM = 2Z R,8Z R,9Z R 3Z R,5Z R 4Z R 5Z R,6Z R 7Z R 8Z R,9Z R 2 23\nhe augmented mechanism state matrix specifies that the nine-bar inkage has 2 DOF. Link 1 is the only fixed link, and all of the oints are movable.\n6.3 Planar Metamorphic Mechanism. The third example ill be that of the planar metamorphic mechanism analyzed in ef. 16 and shown in Figs. 10\u201312. This is a five-link mechanism\nhat oscillates between pins P1 and P2. The spring embedded in ink 2 pushes link 3 along the slot in link 2. In every state the\nechanism is a five-bar linkage with one DOF. However, the state f the mechanism changes as it oscillates between pins P1 and P2.\nMSM = 2Z R,5V X 3V P 4Z R 5Z R 2Z R,5Z R 3V X 4Z R 5Z R\n2Z R,5V X 3V P 4Z R 5Z R\n24\nThe augmented mechanism state matrix for this example is given in Eq. 25 as\nMASM = 2Z R,5V X 3V P 4Z R 5Z R 1 2Z R,5Z R 3V X 4Z R 5Z R 1\n2Z R,5V X 3V P 4Z R 5Z R 1\n25\nIn this case, the DOF does not change as the mechanism changes states. However, the augmented mechanism state matrix shows how the links and joints change as the mechanism moves from one state to another.\nTransactions of the ASME\n016 Terms of Use: http://www.asme.org/about-asme/terms-of-use", "u c a e f m g\na n\nJ\nDownloaded Fr\n6.4 2R-2P Mechanism. Mechanism state matrices can be sed to identify topologically identical mechanisms with different onfigurations. Consider the 2R-2P mechanisms shown in Figs. 13 nd 14. These are topologically identical mechanisms with differnt configurations. The difference between them is evident by orming the augmented mechanism state matrices. The augmented echanism state matrix for the 2R-2P mechanism in Fig. 13 is iven by\nMASM = 2Z R,4X P 3Z R 4Y P 1 26 nd the augmented mechanism state matrix for the 2R-2P mecha-\n1 1 a d\nc\nb\n2\n4\n53\nP2 P1\nY\nX\nism in Fig. 14 is given as\nournal of Mechanisms and Robotics\nom: http://mechanismsrobotics.asmedigitalcollection.asme.org/ on 02/03/2\nMASM = 2V X,4X P 3V X 4X P 1 27 Equation 26 shows that either links 2, 3, or 4 could be an input link. This is not the case for the 2R-2P mechanism with parallel prismatic joints. As shown in Eq. 27 , only link 4 can be an input link. The other three links are fixed. It should be noted that the 2R-2P mechanism with parallel prismatic joints is a special case in which the configuration of the mechanism causes three of the four links to be fixed.\n7 Conclusions This paper introduced mechanism state matrices as a novel way to represent the topological characteristics of reconfigurable mechanisms. It was shown that mechanism state matrices can be used to calculate the DOF of planar mechanisms containing only one DOF joints. This includes mechanisms that contain partially locked kinematic chains. The DOF matrix can then be combined with the mechanism state matrix to form an augmented mechanism state matrix. Augmented mechanism state matrices can be used to represent all five topological characteristics of reconfigurable mechanisms. Future work will involve applying mechanism state matrices to aid in both systematic analysis and synthesis of reconfigurable mechanisms.\nReferences 1 Liu, N., 2001, \u201cConfiguration Synthesis of Mechanisms With Variable\nChains,\u201d Ph.D. thesis, National Cheng Kung University, Tainan, Taiwan. 2 Yan, H., and Liu, N., 2000, \u201cFinite-State-Machine Representations for Mecha-\nnisms and Chains With Variable Topologies,\u201d ASME Paper No. DETC2000/ MECH-14054. 3 Kuo, C., 2004, \u201cStructural Characteristics of Mechanisms With Variable Topologies Taking Into Account the Configuration Singularity,\u201d MS thesis, National Cheng Kung University, Tainan, Taiwan. 4 Kuo, C.-H., and Yan, H.-S., 2007, \u201cOn the Mobility and Configuration Singularity of Mechanisms With Variable Topologies,\u201d ASME J. Mech. Des., 129 6 , pp. 617\u2013624. 5 Yan, H.-S., and Kuo, C.-H., 2009, \u201cReconfiguration Principles and Strategies for Reconfigurable Mechanisms,\u201d ASME/IFToMM International Conference on Reconfigurable Mechanisms and Robots, pp. 1\u20137. 6 Dai, J., and Qixian, Z., 2000, \u201cMetamorphic Mechanisms and Their Configuration Models,\u201d Chin. J. Mech. Eng., 13 03 , pp. 212\u2013218. 7 Dai, J. S., and Jones, J. R., 2005, \u201cMatrix Representation of Topological Changes in Metamorphic Mechanisms,\u201d ASME J. Mech. Des., 127 4 , pp. 837\u2013840. 8 Zhang, L., Wang, D., and Dai, J. S., 2008. \u201cBiological Modeling and Evolution Based Synthesis of Metamorphic Mechanisms,\u201d ASME J. Mech. Des., 130, p. 072303. 9 Parise, J., Howell, L., and Magley, S., 2000. \u201cOrtho-Planar Mechanisms,\u201d Proceedings of the 26th Biennial Mechanisms and Robotics Conference. 10 Carroll, D., Magleby, S., Howell, L., Todd, R., and Lusk, C., \u201cSimplified Manufacturing Through a Metamorphic Process for Compliant Ortho-Planar Mechanisms,\u201d Proceedings of IMECE2005, Paper No. IMECE2005-82093. 11 Yan, H.-S., and Kuo, C.-H., 2006, \u201cTopological Representations and Characteristics of Variable Kinematic Joints,\u201d ASME J. Mech. Des., 128 2 , pp. 384\u2013391. 12 Dai, J., and Wang, D., 2006, \u201cDifferential Geometry Based Analysis of Synthesis of a Multifingered Robotic Hand With Metamorphic Palm,\u201d ASME Paper No. DETC2006-99532. 13 Ziesmer, J. A., and Voglewede, P. A., 2009, \u201cDesign, Analysis and Testing of a Metamorphic Gripper,\u201d ASME Paper No. DETC2009-87512. 14 Tsai, L.-W., 2001, Mechanism Design: Enumeration of Kinematic Structures According to Function, CRC, Boca Raton, FL. 15 Dai, J., and Jones, J. R., 1999, \u201cMobility in Metamorphic Mechanisms of Foldable/Erectable Kinds,\u201d J. Mech. Des., 121 3 , pp. 375\u2013382. 16 Lan, Z., and Du, R., 2008, \u201cRepresentation of Topological Changes in Metamorphic Mechanisms With Matrices of the Same Dimension,\u201d ASME J. Mech. Des., 130, p. 074501. 17 Yan, H.-S., and Kuo, C.-H., 2006, \u201cRepresentations and Identifications of Structural and Motion State Characteristics of Mechanisms With Variable Topologies,\u201d Trans. Can. Soc. Mech. Eng., 30, pp. 19\u201340. 18 Yan, H.-S., and Kuo, C.-H., 2009. \u201cStructural Analysis and Configuration Synthesis of Mechanisms With Variable Topologies,\u201d ASME/IFToMM Interna-\ntional Conference on Reconfigurable Mechanisms and Robots, pp. 23\u201331.\nFEBRUARY 2011, Vol. 3 / 011012-7\n016 Terms of Use: http://www.asme.org/about-asme/terms-of-use" ] }, { "image_filename": "designv11_20_0003586_tencon.2011.6129266-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003586_tencon.2011.6129266-Figure9-1.png", "caption": "Figure 9. A closed up view of the BlackWidow 1.0 on the car with the battery pack below the microcontroller", "texts": [ " CONTROLLER-CAR COMMUNICATION AND EXPERIMENTATION OUTCOMES In order to enable communications in between the iPhone and the car, the default radio frequency receiver on the car was removed and replaced with the BlackWidow 1.0. The wiring works were then modified such that the servomotors that are used to accelerate and to steer the vehicle were connected to the output pins of the microcontroller, thus enabling wireless control of the car through the iPhone. The setup of the BlackWidow 1.0 on the remote car is shown in Figure 8 and Figure 9. Controller The BlackWidow 1.0 supports TCP/IP and UDP/IP protocols with encrypted or open Wi-Fi networks. It also supports a web server function. The connection could be established via 2 methods; ad-hoc and infrastructure. So the BlackWidow 1.0 was programmed to broadcast a SSID, also named \u2018\u2018RCApps\u2019\u2019 so that the iPhone could connect to it. Once the connection has been set up, the output generated by the iPhone through its sensors (touchscreen for touch-based interface and accelerometer for steer-based) could then be transmitted to the BlackWidow 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001431_cdc.2007.4435019-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001431_cdc.2007.4435019-Figure2-1.png", "caption": "Fig. 2. The figure shown positions, black circles, and velocities, of beads i and i + 1 at time t = T k i , t = t0 and t = t1 as described in Theorem 10. The squares are the centers Ci and Ci+1 of the dominance regions Di and Di+1.", "texts": [ " However, due to the switching nature of the dynamics of the beads, the asymptotic behavior of T k is more simple to analyze. On the other hand \u03b4 is a more suitable quantity to describe the asynchrony at time 0. Let us suppose that at time t the beads i and i + 1, with directions di(t) = \u2212di+1(t) = +1, are about to collide. We know that T k i and T k i+1, for some k, are the times at which they passed by the centers of their dominance regions. If T k i < T k i+1, that is bead i is early with respect to bead i + 1, the impact will occur in Di+1 as shown in Figure 2, otherwise it will occur in Di. Without loss of generality we suppose that the impact will occur in Di+1. Let \u03b7 = (T k i+1\u2212T k i )\u03bd. At t0 = T k i + (\u03c0/n) \u03bd bead i reaches the boundary of its dominance region (i.e., pi(t0) = Ui), and distcc(pi(t0), pi+1(t0)) = \u03b7. This is true because when traveling inside its dominance region vi(t) = di(t)\u03bdi(t), and by assumption \u03bdi(0) = \u03bd for all i and, therefore, for all t \u2265 0. Let t1 be the time at which the two beads collide and let \u00b5 = distcc(Ui, pi(t1)). Then we have that: \u03b7 + vi+1(t1 \u2212 t0) = vi(t1 \u2212 t0), \u00b5 = vi(t1 \u2212 t0)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000356_bf01170962-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000356_bf01170962-Figure2-1.png", "caption": "Fig. 2. Conical bearing gap", "texts": [ "7) ac~ 2 1 ~ ~ 1 h~, ~1 (~%) + ~-~ (~%) + ~ 7-3 (&~\" 0%) = o, (3.s) a~--~ ~'~-~ + ~ ~a~/ + ( / ~ I l \"L~a~l +~o~lj =dvV)T%\" (3.9) :For 0 < n < 1, the Eqs. (3.6)--(3.9) describe the flow of a non-~Newtonian power law lubricant through the curvilinear (in width direction) gap of a slide bearing. For n = 1, the equations listed above hold for a ~qewtonian gap flow. The unknown functions v~p p and T may be found by solving the Eqs. (3.6) to (3.9). 16\" (i = 1, 2, 3) become: ~1 ~ ~0, ~ ~ y, c~ 3 ~- x, respect ive ly , see Fig. 2. Thus, the L a m e ' s coefficients a re : h~ = x cos ~*, h v = 1, hx = 1, where ~* denotes the slope of the genera t ing l ine of conical surface. The components of the loeM lubricant veloci ty , v~, vy, vx, the h y d r o d y n a m i c pressure , p, and t empera tu re , T , are now assumed to be of the following forms : 9 (\u00a7 v~ = c o . l . c o s ~ * * v ~ , P =-~o Pl ; v~ = ~ f l \" T \" %,; T -=-- To q- Eo 9 P r 9 ToT1; }/ Vx ~\" 7 \"Ox~ ;\" (4.1) where: vvl , vyl, v~, are the dimensionless components of the local lubr ican t ve loc i ty in the ~, y and x direct ions, respect ively , p~ is the dimensionless hydrodynamic pressure, T1 is the dimensionless t empera tu re , l is the l eng th of the cone genera t ing line, g denotes the height of the gap, ~0 is the (dimensional) charac - ter is t ic v a l u e of lubr ican t densi ty , To denotes the a mb ie n t t empera tu re , co is the angu la r ve loc i ty of the journal , whereas Ee, P r are the E c k e r t and P r a n d t l numbers , respect ively " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003355_j.jeurceramsoc.2011.03.034-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003355_j.jeurceramsoc.2011.03.034-Figure1-1.png", "caption": "Fig. 1. UV laser microprocessing machine.", "texts": [ " ubsequently, chemical etching was used to clean the debris on he kerf surface and remove the thin recast layer, thus reducing Ra o under 0.2 m in order to satisfy the requirement for follow-up ilding on the kerf sidewall surface. . Experimental equipment, materials and methods .1. Experimental equipment and setup The equipment used for this study includes a 355 nm DPSS -switched UV laser with average laser power of 7.5 W at a ulse repetition frequency of 25 kHz for the ultrathin ceramic urface (Fig. 1). A 2D galvanometer scanner controls laser eam to carry out the microprocessing. The main parameters f the UV laser used in the experiments are listed in Table 1. he laser has high stability, with both average power and ulse to pulse energy fluctuations of less than 2%. An ultrahin ceramic plate is fixed on a vacuum table by a clamper Fig. 2). The focal plane is fixed on the surface of the samles and the laser beam spot size at focus is approximately The physical properties of the 125 m thick Al2O3 ceramics sed in the experiments are listed in Table 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002849_j.elecom.2010.10.021-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002849_j.elecom.2010.10.021-Figure1-1.png", "caption": "Fig. 1. Voltammetric responses. Scan rate: 50 mV s\u22121. A) 1-bromoadamantane (concentration: 14.9 mM) in DMF+TBABF4. First eight scans. B) 1-iodo-3-phenylpropane (concentration: 12.4 mM) in DMF+TBAPF6. First ten scans. C) Piece of commercial GC. Solvent: DMF+TBABF4.", "texts": [ " Work up consisted in an efficient sonication of electrodes (for 5 min), then rinsing by acetone before drying by hot stream (60 \u00b0C). Experiments shown hereafter during the reduction of RXs deal with GC surface changes obtained either with recurrent scans run within quite negative potential ranges (say between\u22121.5 V and\u22123 V vs. SCE) or fixed potential electrolyses at E\u226a\u22122 V. Thus, essentially in DMF containing TBABF4, a preliminary approach is based on the assumption that progressive modifications in the course of repetitive voltammetric sweeps recorded without interruption, could correspond to GC surface alterations. Fig. 1, parts A and B exemplify this:while the reduction peak of 1-bromoadamantane (tertiary halide) is absolutely not modified in the course of the first eight scans, that of 1-iodo-3-phenylpropane (primary halide) is progressively shifted towards more and more negative potentials. A thorough ultrasonic rinsing checks that the electrode in contact with the original solution does not recover its initial behaviour and the peak of the RI gets to be definitively shifted. At this point, it is proposed that the reduction process of primaryRXs leads to a changeof carbon interface", " Interestingly, voltammetries of 1-iodohexane have been already found to exhibit a current decay upon scans but any organic deposit could be characterized on GC [11]. On the contrary, voltammetries at copper and gold ofRIs (IC6-H13 and I-(CH2)2-C8F17)were reported to concomitantly exhibit progressive potential shifts and current decays. Evidences for grafting of carbon chains onto those metals were argued. Such immobilization was proposed to occur through a free radical addition. Specifically, the electroactivity of the used GC was checked to be effective. Thus, its voltammetry achieved on a fragment (Fig. 1, curve C) shows a large reversible step (E0\u2248\u22121.94 V) assigned to superficial graphitic areas with a charge threshold at \u22121.75 V. After running the SEM image of the resulting surface (Fig. 3, image 2), it appeared that a lone scan provokes a huge modification of the external structure of the GC, assigned to insertion of TBA+ cations into the material. These observations then close to that already obtained with highly oriented pyrolytic graphite (HOPG) under quite similar conditions [12]. The quasi-reversible system (threshold of the charge at about \u22121" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002349_s12239-009-0039-8-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002349_s12239-009-0039-8-Figure4-1.png", "caption": "Figure 4. Kinematics of the ISCVT.", "texts": [ " The ISCVT, as shown in Figure 2, has two rolling contact points between three rotors: the driving rotor, the counter rotor, and the driven rotor, as do most speed ratio changeable traction drive CVTs. The engine power flows successively through the driving shaft, the driving rotor, the counter rotor, the driven rotor, and the driven shaft. The driving shaft is connected to the engine output, and the driven shaft is connected to the rear wheel. The rolling bodies, the convex counter rotor and the concave driving and driven rotors, have exactly spherical surfaces with the functionally unnecessary parts eliminated, as shown in Figure 4. These spherical surfaces are located so that the contact points are tangential to each other, meaning the three points consisting of the two centers of the spheres and the contact point are collinear. This structure makes a larger circular contact area and helps ensure stable operation. The circular contact area improves the torque capacity and reduces the transmission size. It is also more effective in supporting the extremely high stresses on the contact point, which is an inescapable problem of the traction drive CVT", " There is a tilt handle under each of the counter rotors, which is guided by the spiral groove on the drumshaped cam. As this drum-cam rotates, the counter rotors are synchronously inclined. The pressure to produce the traction force on the contact point is generated by the coil spring and the loading cam. The coil spring is installed in the counter rotor assembly with the designed preloading, which improves the efficiency of the initial operation. The loading cams of the ISCVT are on the driving and driven rotors. A ball on a cone-grooved surface converts the torque on the shaft into the thrust force. Figure 4 shows the kinematics of the ISCVT, where the points O1 and O2 denote the center of the sphere of the driven rotor and driving rotor, respectively. The tilting angle \u03c6 determines the speed ratio \u03c1. As previously mentioned, the contact points Q1 and Q2, the tilting center of the counter rotor O\u03c6 , and the points O1 and O2 are on lines O1O\u03c6Q1 or O2O\u03c6Q2. When the speed ratio changes, the positions of the rolling contact points Q1 and Q2 on the counter rotor move slightly along the circumference of the counter rotor", " The kinematic design procedure determines the set of several consistent geometric parameters for basic operation under the given specifications and constraints. The kinetic design then optimizes the kinematic results according to the objective function, which implies the design requirements. As a result of the design procedure, the principle design parameters for the prototype are listed in Table 2. Having driving and driven rotor of the same diameter can be more efficient than different diameters. The angle of the center of the counter rotor denotes the angles Q1O1O2 and Q2O2O1 in Figure 4, which are the same value because the driving and driven rotors have the same diameter. These angles determine the distance of the counter rotor from the rotating axis of the main shaft of the ISCVT. A longer distance generates more efficiency but also forces a larger transmission size. The designed range of the tilting angle for the counter rotor is 0.17o~31.55o, which is not symmetric but necessary to allow for a more compact size. A more detailed structure design of the ISCVT and its associated parameters can be found in Ryu et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000319_095440605x8478-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000319_095440605x8478-Figure4-1.png", "caption": "Fig. 4 Multiple rigid bodies connected by multiple flexure hinges", "texts": [ " 2, the force vector consisting of force and moment components applied to the point oi by the small displacements Dri and Dui is Qi \u00bc \u00bdFij j Tij \u00fe dij Fij T (20) Substituting equation (19) into equation (20) results in Qi \u00bc ~DijK ji ~D T ij Dqi (21) In addition, the force vector applied to the point oji on the fixed frame is Q ji \u00bc \u00bdFij j Tij \u00fe h ji Fij T (22) wherehji is the distance vector from the point oji to the point oij and is given by h ji \u00bc \u00bdhxji hy ji hz ji T (23) Substituting equation (19) into equation (22) results in Q ji \u00bc ~H jiKij ~D T ij Dqi (24) where ~H ji \u00bc I 0 H ji I (25) and H ji \u00bc 0 hzji hy ji hz ji 0 hxji hyji hx ji 0 2 4 3 5 (26) Whena rigid body i is connected to a rigid body jwith a flexure hinge, as shown in Fig. 3, the force vector (consisting of force and moment components) applied to the point oi by the small displacements Dqi and Dqj is Qi \u00bc ~Dij K ji ~D T ij Dqi ~HijKij ~D T ji Dqj (27) Consider a rigid body i connected to multiple rigid bodies with multiple flexure hinges, as shown in Fig. 4. The rigid body i is connected to nb rigid bodies with flexure hinges. In addition, the rigid body i and Proc. IMechE Vol. 219 Part C: J. Mechanical Engineering Science C19203 # IMechE 2005 at NATIONAL UNIV SINGAPORE on June 27, 2015pic.sagepub.comDownloaded from another rigid body hold nh flexure hinges in common. The force vector applied to the point oi is Qi \u00bc Xnb j\u00bc0 j=i Xnh k\u00bc1 ~Dijk(K jik ~D T ijk Dqi ~HijkKijk ~D T jik Dqj) (28) where j \u00bc 0 indicates that the rigid body i is connected to a fixed body 0 with flexure hinges" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001821_bi00775a021-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001821_bi00775a021-Figure6-1.png", "caption": "FIGURE 6: Ionic strength dependency of Mgz+-isocitrate chelate dissociation constants with ionic strength. Sources of data: 0, Blair (1969); 0 , Londesborough and Dalziel (1970); 8, Duggleby and Dennis (1970); ellipse, our data from Figure 5B; the long axis of the ellipse indicates the range of ionic strength spanned by the addition of Mgz+ ion in the determination.", "texts": [], "surrounding_texts": [ "I S O C I T R A T E D E H Y D R O G E N A S E : S U B S T R A T E A D D I T I O N 0 R D E R\n6 .5 4 . 0 f 1 .2 2 . 0 f 2 . 7 6 . 3 & 0 . 4 13.7 =t 1 . 3 6 . 7 6 . 3 f 0 . 3 12.6 f 9.0 7 . 0 7 . 6 f 3.6 7.6 f 0 . 5 7 .2 8 . 3 f 9 . 8 6 . 8 * 5 . 4 9 . 2 + 5 . 4 13.9 =+ 8 . 2 7 . 5 5 . 2 f 1 . 5 4 . 6 * 4 . 1 6 . 4 + 0.6 11.5 * 0 . 4 8 . 0 8 . 0 =t 0.3 3 . 2 & 0 . 7 8 . 2 10.7 * 0 . 7 19.8 =t 1 . 3 8 . 5 10 .3 f 0 . 9 7.8 f 0 . 9 9 . 4 7 .5 f 0 . 3 10.5 0 . 7\n'The values are all expressed as M X lo6. 'Phosphate buffer was used for pH values 6.5, 7.0, and 7.2 with the ionic strength being maintained at 0.053. Tris-HC1 buffer was used at pH values 7.5, 8.0, and 8.5 and glycine-KOH at pH 9.4. The ionic strengths at these latter pH values were 0.053, 0.008, 0.029, and 0.044, respectively. Phosphate buffer, 0.025 M, was used from pH 6.5 through pH 7.2. Tris buffer, 0.025 M, was used at higher pH values.\ncurred at pH 6.8 on the acid side. A complete rate-pH profile was not obtained for the reverse reaction because of the difficulty in obtaining saturation by COZ at high pH values.\nThe values of the kinetic constants for the forward reaction obtained from secondary replots of initial velocities at various values of pH are listed in Table 11. Note that care was exercised to maintain a constant ionic strength of 0.053 from pH 6.5 to 7.5. It is evident that these constants bear little or no pH dependency. Under conditions where no precautions were taken to control the ionic strength, apparent variations of the kinetic constants with pH were observed, but the variations were irregular with changes in pH and only on the order of fourfold in magnitude at most. It could be argued that such apparent variability could have arisen from specific ion effects superimposed on ionic strength effects ; however, the data of Table I1 collected from experiments performed in three different buffer systems give no support to this contention. The Michaelis constants for TPNH and 2-ketoglutarate obtained from the reductive carboxylation reaction are also listed in Table I1 as a function of pH. A constant buffer concentration of 0.025 M was used in all these determinations with no special effort being made to control ionic strength. Again, little or no pH dependency of these constants is indicated.\nCOz L'S. HCOB- as the Carboxylation Substrate. The velocities of the reductive carboxylation reaction at pH 6.4 and 7.2, initiated by the addition of unbuffered sodium carbonate solutions, are graphed as functions of time in Figure 4. The inset of Figure 4 portrays the plots of -log (1 - (u/u,,)) us. time for the carboxylation reaction at these two pH values, along with an indication of the times at which the velocities reach 0.5 and 0.9 of their respective limiting values. The rate us. time curves obtained for reactions at pH 6.4, 7.2, and 8.0 initiated by the addition of unbuffered COz solutions are also shown in Figure 4. Significant amounts of the preequilibrium COZ were consumed in this reaction relative to the amounts of CO, remaining at or near equilibrium in the absence of reductive carboxylation. Hence, the data cannot be legitimately plotted in the logarithmic form which relates enzy-\nmatic reaction velocities to the rate constant for the spontaneous decay of CO2 to bicarbonate.\nIn Table 111, the initial velocities obtained for the reductive carboxylation at different pH values using COS as the initiating form are compared with the corresponding initial velocities for reactions using preequilibrated HC03--C02 solutions at an overall concentration of 4 X lop3 M. Notice that with increasing pH, the ratio of velocities in the unbuffered and preequilibrated reaction mixtures increases as does the ratio of COz concentrations in the unbuffered and buffered reaction mixtures. In the former mixture the initial ratio of CO, to HCOI- is constant though of uncertain magnitude.\nThe Michaelis constants for total HC03- (HC03- + COS), K H C O ~ - ( ~ ) , and for COS only, Kco,, are given at various pH levels in Table IV. The latter constants were obtained by multiplying the values of KHCO,-(t) by the term (1 + l o p H --6,35)-1 where the quantity 6.35 represents the negative logarithm of the equilibrium constant for the reaction COz + HzO e HC03- + H+ (Harned and Owen, 1958). In the region of maximum catalytic activity occurring between pH values 7.4 and 8.4, these constants exhibit the general behavior with respect to pH that is expected for a system wherein COS, rather than HC03-, is the true substrate, i.e., KCO, remains approximately constant throughout this pH range, whereas KHco3- (t) increases, reflecting the progressively diminishing C02-HC03-(t) ratios. In contrast to the Michaelis constants obtained for the other substrates of this enzyme, KCO, is a sensitive function of pH below 7.4. The implications of this variation are treated in the Discussion.\nMg2+ Zon Efects. Since the overall oxidative decarboxylation reaction has a requirement for Mgz+ ions (Chung and Franzen, 1969), the kinetic parameters describing the reaction were evaluated at different MgZ+ ion concentrations and are set forth in Table V. It is evident that, among these parameters, only the Michaelis constant for isocitrate is significantly affected by changing the concentration of MgClZ. This sort of behavior was also observed with the beef heart enzyme\nB I O C H E M I S T R Y , V O L . 1 1 , N O . 2 5 , 1 9 7 2 4771", "W I C K E N , C H U N G , A N D F R A N Z E N\n(Londesborough and Dalziel, 1970). It might be inferred from the variation of KIC with Mgz+ ion concentration that it is actually the Mgz+-isocitrate chelate which is the active substrate of this enzyme. If so, the Michaelis constant for this species should be independent of Mg2+ ion concentration. Some computations of K M ~ I C were made based on reported values of the dissoriation constants of the MgHP04 and Mgisocitrate chelates, corrected for ionic strength effects (Greenwald et al., 1940). The ionic strength correction applied to the isocitrate chelate dissociation constant was ~KM,IC = 3.5-\n4772 B I O C H E M I S T R Y , V O L . 1 1 , N O . 2 5 , 1 9 7 2\nTABLE v: Effects of Magnesium Concentration on the Turnover Numbers and Michaelis Constants for the Oxidative Decarboxylation of D-Isocitrate.'\n(VlIE) x MgC12 10-2 (mol of KTPN+ X lo3 TPNH/(min X lo5 KIC X l o 5 K M ~ I C X\n(M) mol of IDH)) (M) (M) 105 (M) _ _ _ _ _ _ _ _ ~ ~ ~\n5 88.5 f 1 . 2 2 . 0 f 0 . 1 1 . 7 f 0 . 1 0.873 1 91.2 f 1 . 1 2.1 f 0 .1 2 .8 f 0 . 1 0.488 0 . 4 94 f 10 2 . 2 i- 0 . 1 4 . 3 f 0 . 2 0.335 0 . 2 90.9 f 4 . 4 1 .4 f 0 . 1 4 .1 f 0 .4 0.166 0.1 89.3 i- 3 . 0 1 .6 f 0.1 6 .2 + 0 . 3 0.128 0.05 82.7 i- 7 .0 2 . 4 1 0 . 5 11 .5 i 0 .9 0.120\n' All experiments were conducted in potassium phosphate buffers at pH 7.3. The ionic strength was maintained a t approximately 0.13 M.\n3.1 (K) ' '~ . This relation was obtained by plotting the negative decadic logarithm of dissociation constants determined by a number of workers, including ourselves, as a function of (p)''?; see Figures 5 and 6 (Duggleby and Dennis, 1970;", "I S 0 C I T R A T E D E H Y D R O G E N A S E : S U B S T R A T E A D D I T I O N 0 R D E R\nBlair, 1969; Londesborough and Dalziel, 1970). The pK, for the dissociation of protons from H2POa-, also necessary for the computation of KM~IC, was estimated to be 6.7 in the ionic strength region of the experiments under consideration (Bernhard, 1956; Alberty et al., 1951). The K M ~ I C values so calculated appear in the last column of Table V. There is some uncertainty in the values of this column, since the ionic strength dependencies of the chelate and H2P04- dissociation constants are not precisely known. In addition, no correction was made for the amount of Mgz+ ion chelated by TPN (Colman, 1972a). As the total amount of TPN was relatively small in the kinetic experiments under consideration, the effect of Mg2+ ion binding by TPN on the availability of free Mg2+ was considered negligible. Furthermore, Mg2+ ion binding by TPN is presumably ionic strength dependent, and there is no information available as to how great this sensitivity is. Slight corrections, however, would not alter the observation that Knlg1c is not invariant with total Mg2+ ion concentration. The lack of constancy indicates that the postulated interaction of enzyme only with the Mg-isocitrate chelate is inadequate by itself to account for the observed behavior of the system. A mechanism invoking the existence of all the equilibria shown below is probably involved in the generation of the proposed active enzyme-substrate complex, E-Mg-isocitrate.\n, M s E + IC \\\n== E \\ + MgIC e p + H / IC + E / + Mg \\\n% / + Mg'\nAccording to this scheme, the increasing values of K M ~ I C with rising Mg*+ ion levels of Table V reflect the fact that at higher Mg2+ ion concentrations, there are fewer free sites to which the Mg-isocitrate chelate can bind, i.e., the chelate must displace a MgZf ion from the enzyme before it can bind. Clearly, a knowledge of the binding of Mgz+ ion to the enzyme would be useful here, but at the moment, this information is unavailable for the A . vinelandii enzyme.\nSubstrate Binding. The various enzyme-substrate interactions along with the corresponding dissociation constants, as determined by ultrafiltration methods, are listed in Table VI. Representative Scatchard plots from which some of the constants of Table VI were obtained are shown in Figure 7. Isocitrate, TPN+, and TPNH are seen to form strong complexes with IDH while 2-ketoglutarate binds much more weakly, which is in accord with the proposed random order of reactant addition and sequential order of product release.\nThe detection of the binding of isocitrate as monitored by optical rotation changes is illustrated in Figure 8. The saturation phenomenon observed in panel A indicates a differential molar rotation at 310 nm of about -7000 deg cm2 dmol-I, based on the molar concentration of protein. This is suggestive of the action of bound isocitrate being on the rotatory strength of a protein chromophore rather than vice versa, since the nearest Cotton effect for isocitrate is well below 240 nm; see Figure SA. That this is so is observed in the circular dichroic (CD) spectra of Figure 9. Even though C D spectra in the presence of DL-isocitrate, at concentrations which would\n~~ ~~~\nTABLE VI: Dissociation Constants for IDH-Substrate Complexes Determined in the Presence of MgC12 at pH 7.3 in 0.05 M Potassium Phosphate Buffer.\nLigand\nTPN +\nIC\nTPNH\nK G\nTPNH\nK G\nK G\nIC\nTPNH\nMgCh X 103 (MI Reaction\n5 0\"\n5 0\n5 OC\n10 0 10 0 10 0 10 0 20 0 20 0\nE-TPN+ + E + TPN +\nE-IC + E + IC\nE-TPNH + E + TPNH E-KG E + K G\nE-TPNH-KG + E-KG + TPNH E-TPNH-KG + E-TPNH + K G E-TPN +-KG E-TPNf + K G E-TPNH-IC E-TPNH + IC E-TPNH-IC t-r\nE-IC + TPNH\nx 105(M)\n2 . 7 f 0.6\n1 . 7 f 0 . 2 5 . 2 i 0 . 5\n5 . 1 f 0 . 7 4 . 1 i 0 . 4 2 . 6 + 0 . 3\n(1.8 f 0 .9 )*\n(2 .0 f 0.9)b\n118.0 + 21 74.0 i 26\n3 . 0 i 1 . 4 2 . 1 i 0 .2\n27.0 f 5 . 0 45.6 i 9 . 7 75.0 f 14.6\nNot determined Very large\n7 . 3 f 1 . 2 Very large\n7 . 9 i 1 . 3 -~ ~ ~\n\" In the absence of MgC12, sufficient KC1 was added to maintain the ionic strength constant. Values obtained from kinetic constants of Table I and the application of relations presented in the Discussion in this paper. This binding study was done at pH 8.0. In the presence of 5 x M MgClz, at pH 8.0, KTP\" = 3.1 f. 0.3 x M as determined by ultrafiltration binding studies.\ncause partial saturation, fall in between those shown in Figure 9, less than ideal signal-to-noise ratios prohibit the use of the difference in the C D spectra of bound and unbound enzyme for the measurement of binding constants. Nevertheless, there is indeed a negative CD perturbation of the the right order of magnitude, l o4 molar ellipticity units, to account for the rotation change observed at 310 nm. From panel B of\nB I O C H E M I S T R Y , V O L . 11, N O . 25 , 1 9 7 2 4773" ] }, { "image_filename": "designv11_20_0000326_1.1897410-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000326_1.1897410-Figure3-1.png", "caption": "Fig. 3 Possible traction-drive configuration for transmission application", "texts": [ " This will, in turn, create 632 / Vol. 127, JULY 2005 rom: http://mechanicaldesign.asmedigitalcollection.asme.org/pdfaccess.as sufficient friction between rolling elements to prevent gross slip. Furthermore, the angle of inclination can be designed for the expected coefficients of friction to ensure that slip will not occur no matter what the output load. Hence, the traction effect is automatically achieved at just the right level to transmit the instantaneous torque required for the application. Figure 3 shows an embodiment of the device that could be used for a transmission with two different speed ratios. Along a vertical line are two intermediate roller assemblies with outer rollers labeled A, and along a horizontal line are two other intermediate roller assemblies with outer rollers labeled B. The A set is shown engaged, whereas the B set is disengaged. The diameters associated with the A set of rollers would be different than those of the B set of rollers. This would produce different speed ratios for the two sets of rollers. Thus, the device could be shifted from one speed to another by disengaging the A set of rollers and engaging the B set. More than just two sets of rollers can be included in the device, allowing for more than two speed ratios. Also, similar units could be cascaded with the one shown to produce as many different speeds as desired. For the members labeled 2\u20137 in Fig. 3, the speed ratio is given by 7 2 = r2r4r6 r3r5r7 1 where rn is the radius of each roller with n=2,3 , . . . ,7. Members 3 and 4 rotate together as do members 5 and 6. In fact, 3 and 4 are likely to be different sections of the same roller with different diameters, and the same would be true of 5 and 6. On the other hand, if the device is to be used for a single speed reduction, as shown in Fig. 1, then the speed ratio would be 5 2 = r2 r5 2 It should be noted that for this configuration, the output member will rotate in the same direction as the input member" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000086_2004-01-1064-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000086_2004-01-1064-Figure1-1.png", "caption": "Figure 1 Kinematics of a tire during braking and cornering. Force vectors are also shown. (Left: top view; Right: side view)", "texts": [ " The main purpose of this work is to present the new model in a straightforward and application oriented formulation, suitable for direct implementation by the end user. Refer to [5] for technical details and discussions on background theory. In the remaining part of this section necessary definitions are introduced. In the next section the brush model is briefly recapitulated. After that the new model is proposed. The final section presents some results and a validation using real tire data. This section describes the relevant tire kinematics and introduces definitions that are used in the following. Refer to Figure 1 for illustration. Vectors have two components and are denoted by a bar as in u\u0304. The corresponding components and magnitude are denoted by ux, uy, and u. The wheel-travel velocity v\u0304 = (vx,vy ) deviates from the wheel heading by the slip angle \u03b1 tan(\u03b1) = vy vx (1) The circumferential velocity of the wheel is vc = \u2126Re (2) where \u2126 is the wheel angular velocity, and Re the effective rolling-radius of the tire. The slip velocity, or the relative motion of the tire in the contact patch to ground, is v\u0304s = (vx\u2212 vc,vy ) (3) The direction of the slip velocity is denoted by tan(\u03b2 ) = vsy vsx (4) The tire slip is defined by normalizing the slip-velocity with a reference velocity" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003589_ccca.2011.6031525-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003589_ccca.2011.6031525-Figure6-1.png", "caption": "Figure 6. The horizontal wind power generation system [5]", "texts": [ " The turbine under the wind action, drives the induction machine beyond its synchronous speed to operate as a generator. For a machine with two pairs of poles connected to grid of 50 Hz frequency, a gearbox is then necessary to fit the turbine speed to that of the asynchronous generator. The fixed speed wind generator model is shown in Figure 2. A. Gearbox modelling The gearbox has the task to transfer the aerodynamical power from the slow rotating rotor shaft to the fast rotating shaft, which drives the generator at the mechanical speed Qmec (Figure 6). It is mathematically described by the following equations [5]: G is the gear ratio. Cmec =Ct /G Q\",ec = G Q\",ec (8) (9) B. Mechanical shaft modelling The Modeling of the mechanical transmission can be easily determined using the dynamic equation. The mechanical system is represented by the following equation : (7) where JT(kg.m2) is the total inertia which appears on the drive shaft, including the inertia of the generator, turbine, two shafts and gearbox, Qmec (rad/s) is the mechanical speed on the generator axis, Cem (N" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003068_j.jbiomech.2009.04.048-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003068_j.jbiomech.2009.04.048-Figure5-1.png", "caption": "Fig. 5. Posture of zero Euler angles.", "texts": [ " The principle of transferring coupling equations from the Euler to the DH system is that the same posture can be represented by different orientation representation systems. The procedure is shown in Fig. 4. Both Euler and DH system have the same global coordinate system. However, DH local frames are different from Euler\u2019s frames. Figs. 5 and 6 show the postures corresponding to zero Euler angles and DH angles, respectively. The posture of zero Euler angles does not exist and it is used for calibration purpose. In Fig. 5, the local frames are same as those in Fig. 3. This section will illustrate the analytical inverse kinematics method by one example. By choosing ah \u00bc 451 and bh \u00bc 451, through Eqs. (1)\u2013(3), we obtain all Euler angles in Table 1 and the posture is in Fig. 7. Considering AC joint center for both systems, one has the following equations: R\u00f0ac;bc ; gc\u00de L2 0 0 0 B@ 1 CA 1 2 6664 3 7775 \u00bc T0 1\u00f0q1\u00de T 1 2\u00f0q2\u00de 0 0 0 1 2 6664 3 7775 (4) Bringing in all necessary terms in Eq. (4) yields sin\u00f0q2\u00de \u00bc cos\u00f0bc\u00de sin\u00f0ac\u00de (5) sin\u00f0bc\u00de \u00bc cos\u00f0q2\u00de sin\u00f0q1\u00de (6) Solving Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001583_1077546307078829-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001583_1077546307078829-Figure5-1.png", "caption": "Figure 5. Gear dynamic model with time-varying mesh parameters and sliding friction.", "texts": [ " The mesh locations increase and decrease in a quadratic manner for pinion and gear respectively, as shown in Figure 4(a). Further, Figure 4(b) shows that the sliding direction of the second teeth pair changes at the pitch point. at Bibliothekssystem der Universitaet Giessen on May 31, 2015jvc.sagepub.comDownloaded from The linear time-varying (LTV) model with sliding friction, that was developed by Vaishya and Singh (2001, 2003), is employed to describe the dynamic behavior of the gear system. The dynamic model is shown for a spur gear pair in Figure 5 where Ff and x are the sliding friction force and its moment arm respectively. The governing equations are briefly described below: Jp p [c t Rp p Rg g k t Rp p Rg g ] [Rp i i t x pi t ] Tp (24a) Jg g [c t Rg g Rp p k t Rg g Rp p ] [Rg i i t xgi t ] Tg (24b) Here, is the friction coefficient and i , j are tooth indices of pinion and gear respectively. The above equations assume rigid shafts and bearings and thus the effects of flexibility of such adjacent structures as shafts, gear casing and bearings are not included in this model" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002139_978-3-540-89393-6_15-Figure15.6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002139_978-3-540-89393-6_15-Figure15.6-1.png", "caption": "Fig. 15.6 A cardboard (a) and paper (b) wing, demonstrating how three flexion lines can allow and control supinatory twisting in the upstroke. Explanation in the text", "texts": [ " The amount of bending and angle of twist are related to the height of the wing camber proximally to the line of flexion, and this seems to be actively controllable by muscles at the wing base. Bending and torsion of a cambered or pleated wing involve elastic deformation, and it seems that wings too are resonant structures. Like the thoracic box they need to deform correctly at their working frequency, and one can identify morphological features that seem adapted to tune them to do so. These principles can readily be modelled physically and could certainly be used in designing wings for an MAV. Figure 15.6 shows one such model. Support is provided by the stippled area, which is made of thin card. The leading edge section is curved ventrally and is crossed by two oblique lines of flexibility, a\u2013a and b\u2013b, made by cutting part way through the card with a sharp blade. The broad basal stippled area is also cambered and is crossed by a longitudinal flexion line c\u2013c that crosses b\u2013b. The rest of the wing is made of paper. The wing has interesting and unexpected properties, best understood by making the model", " If the force is applied to the convex side, behind the wing\u2019s torsional axis, as it would be in an upstroke, the leading edge bends slightly about a\u2013a, and the wing twists readily towards the tip and could easily assume a positive angle of attack and generate useful upward force. Bending about b\u2013b greatly enhances the twisting, but this is controllable by varying the camber of the basal area around c\u2013c. When the base is nearly flat, bending at both a\u2013a and b\u2013b allows the distal part of the wing to twist dramatically (Fig. 15.6b). Steeper basal camber limits bending to a\u2013a and the wing twists far less. A wing so designed would twist automatically to some extent in the upstroke, but the extent could be controlled over a wide range by simple basal actuation. It seems possible that a complete system comprising a transmission with the properties of a five-bar linkage in a resonant, springy shell, together with a pair of smart wings with actively variable basal camber, could drive an MAV having mechanical control over all the kinematic variables that we have identified as essential for versatile flight, with a rather small number of appropriately designed and located actuators. The mechanism in Figs. 15.4 and 15.5 and the wing in Fig. 15.6 are na\u00efve examples; a sophisticated design would be a deformable monocoque, optimised using modern modelling software, with similar optimised wings. Such a system would have the added advantages of low weight and inertia and relative economy through cyclic elastic energy storage and release. It should moreover be fairly easy to build and replicate, and be capable in time of progressive miniaturisation, as smaller motors, power stores and control circuits become available. 1. Avadhanula, S., Wood, R" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000923_tasc.2005.849119-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000923_tasc.2005.849119-Figure3-1.png", "caption": "Fig. 3. Sketch of a fragment of one layer of a tilted coil wound with a circular cross-section conductor (wire).", "texts": [ " Perhaps there exist several ways to explain mathematically why transverse magnetic field generated by a wire-wound tilted coil can be so uniform [6]. Still this problem has not been properly treated in literature; at least, we failed to supply this paper 1051-8223/$20.00 \u00a9 2005 IEEE with any references. Here we suggest a rather simple method to prove that wire-wound circular tilted coils have a generalized \u201ccosine-theta\u201d distribution of axial current density. In order to do so, it is sufficient to consider a fragment of one layer of the winding (Fig. 3). Let us assume that the wire is wound in such a way that centers of cross-sections of the wire lie on the following parametric curve (a tilted cylindrical helix, Fig. 3): (1) , , and are constants here (see (9) below). An increase of by leads to a shift by a vector (independent of ): . A vector tangent to curve may be found by differentiation with respect to : (2) so the current vector equals (3) where is the current in the wire, , are the magnitudes of the relevant vectors. The average current density vector equals (4) where is the diameter of the wire, and is the distance between tangents to curve drawn in two points where angle differs by (Fig. 3). It should be remembered that the magnitude of the average current density, , depends on , although neither the current, , nor the wire cross-section depend on . This is due to the fact that the winding cannot be uniformly tight in this geometry, and the gap, , between two adjacent turns of the winding depends on : (5) The gap is minimal at , where is integer; in particular, in the vicinity of these points the adjacent turns can touch each other if the winding is closely packed. The gap is widest at and noticeably depends on the tilt angle (Fig. 3). Evidently, (6) where is the angle between vectors and at a given value of (see Fig. 3). Consequently, (7) and thus the axial component of the average current density turns out to be (8) It is noteworthy, albeit almost obvious, that the constant (and independent of tilt angle) component of the axial current arises due to the helical path of wire\u2014it vanishes if each turn is considered a flat ellipse. Technically, the presence of the constant component of axial current is a distinction, compared to classic cosine-theta coils, which basically are not supposed to have such a \u201cmakeweight" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003610_1830483.1830692-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003610_1830483.1830692-Figure1-1.png", "caption": "Figure 1: GENMAV in AVL", "texts": [], "surrounding_texts": [ "Traditional control techniques such as model-based PID control have been used successfully for many control problems including aircraft control [11, 20, 19, 13, 17]. These model based techniques perform very well for full size aircrafts where linear approximations of the system\u2019s dynamics produce good results. Micro Aerial Vehicle size renders them particularly susceptible to wind gusts and disturbances. These conditions make them notoriously difficult to control since many of the assumptions used in the controllers break down in such environments. Accurately modeling the MAV and its environment is also critical to design model based controllers which are typically not very robust to modeling inaccuracies. Moreover, PIDs typically involve tuning and optimization of the gains to achieve optimal results. This step can however be improved using neural networks [17]. Improvements can therefore be achieved by providing more flexible controllers that can adapt automatically to model error and changes in the environment. Learning based techniques are flexible, do not require a model of the system and can adapt to different platforms and dynamic environments. These techniques have been proven to be robust and efficient for complex control problems such as multi-rover control problems [3] where a large number of agents have to maximize the overall system level evaluation function that rates the performance of the full systems as well as their own evaluation functions. They are therefore well suited for MAV control where the system is highly non-linear, unstable and where obtaining an accurate model of the system is not a trivial task. The configuration of these platforms is flexible and provide many parameters that can be used by the learning based control system to stabilize the system and perform flight maneuvers. Another benefit of such a system would be its ability to control the system when noise and/or failures occur and adapt to the new structure of the system. A difficulty in applying a learning technique is the design of the evaluation function so that the correct mapping between MAV states (position, velocity) and MAV actions (actuator positions) can be achieved. Tuning of the learning parameters is also necessary in order to achieve optimal results. Sections 5.1, 5.2, and 5.3 show the results of learning based methods applied to the MAV control problem." ] }, { "image_filename": "designv11_20_0002982_10402000802105448-Figure12-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002982_10402000802105448-Figure12-1.png", "caption": "Fig. 12\u2014Speed gradients along z-axis at x = 0.", "texts": [ " This additional effect is enhanced because the distance between the center of rotation T ri bo lo gy T ra ns ac tio ns 2 00 9. 52 :1 71 -1 79 . of the spherical end specimen and the contact center (= RBsin\u03bb, see Fig. 2) is of the same order of magnitude as the contact diameter. In classical contact configuration the former distance is much longer and the latter is smaller, thus minimizing this effect. The speed gradients along two lines belonging to the two bounding surfaces in the y-z-plane (i.e., at x = 0) are schematically drawn in Fig. 12. Moreover, the speed vectors outside the contact transverse central plane and therefore at x = 0, gain a new component in the z-direction. This effect is clearly described in Fig. 13. The transverse speed motion increases when going away from the z-axis. This variation produces another additional shear stress in the direction perpendicular to the main entrainment one. Except for very low tilting angles and very large contact size, the transverse velocity is always smaller than the entrainment speed on the spherical end specimen surface" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002943_074683410x480276-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002943_074683410x480276-Figure3-1.png", "caption": "Figure 3. Two functions 1 \u2212 R and e\u2212\u03bbR intersect at fixed points (filled dot for stable, and hollow dot for unstable).", "texts": [ " The sudden increase of R\u221e, the fraction of the population that has been affected by the epidemic, at \u03bb = 1 is apparent. Algebraically, when S0 = 1, the numerator of the second term in (4), W (\u2212\u03bbe\u2212\u03bb), is simply \u2212\u03bb for \u03bb \u2264 1, because W is the inverse function of T (w) = wew for w \u2265 \u22121. Therefore, R\u221e = 0 for \u03bb \u2264 1. Furthermore, R\u221e is strictly increasing for \u03bb > 1 and tends to 1 as \u03bb \u2192 \u221e. Transcritical bifurcation The epidemic threshold can be viewed as a bifurcation point. A graphic analysis, using S0 = 1, is given in Figure 3 (following [6]). The two components of R\u2032 according to (2), namely the line y = 1 \u2212 R and the curve y = e\u2212\u03bbR , are plotted. The dynamics of R are indicated along the R-axis: by an arrow to the right if R\u2032 > 0 and an arrow to the left if R\u2032 < 0 For \u03bb = 1, the line and curve intersect twice, and one intersection is invariably R = 0. For \u03bb < 1, the zero root is a stable fixed point; for \u03bb > 1, the nonzero root is stable. As \u03bb increases from less than 1 to greater than 1, the fixed points pass each other (at R = 0, of course) and exchange stability" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002265_20070822-3-za-2920.00043-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002265_20070822-3-za-2920.00043-Figure2-1.png", "caption": "Fig. 2. The path \u03c3s in \u2126(\u0393) in R2 + can be chosen to be piecewise linear.", "texts": [ " An intermediate result is the following, that already implies a semi-global version of Theorem 5, where semi-global means \u201con arbitrarily large compact sets around the origin\u201d. Proposition 12. Let \u0393 \u2208 (K\u221e \u222a {0})n\u00d7n, \u0393 id, be such that \u0393 has no zero rows. For every s \u2208 \u2126 there exists a continuous and strictly increasing vector function \u03c3s : [0, 1] \u2192 (\u2126 \u222a {0}) \u2229B1(0, |s|) with \u03c3s(0) = 0 and \u03c3s(1) = s. Moreover, each component function is piecewise linear on every interval of the form [\u03b5, 1], \u03b5 > 0. Figure 2 shows what this looks like in two dimensional space. Proof. This construction was performed in the proof of Lemma 9 (iv), see Dashkovskiy et al. (2006a) for further details. 2 This gives one direction of the path, the other direction is given next. Theorem 13. Let \u0393 \u2208 (K\u221e \u222a {0})n\u00d7n, \u0393 id, be primitive. Then there exists a piecewise linear and strictly increasing vector function \u03c3 : R+ \u2192 \u2126 \u222a {0} with \u03c3(0) = 0 and limt\u2192\u221e \u03c3(t) = \u221e, i.e., the component functions are of class K\u221e. Proof. By Lemma 9 (vi) we have \u03a8\u221e \u2282 \u2126 \u222a {0}" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002036_fuzzy.2007.4295490-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002036_fuzzy.2007.4295490-Figure7-1.png", "caption": "Fig. 7. path planning result", "texts": [ " In all of simulations the underwater vehicle starts from the point (0, 0, 0) to the end point (60, 0, 70). Characteristics of moving obstacles are shown in Tables (V, VI, VII), respectively. Figures (4, 5, 6,) show the moving paths of underwater vehicle navigated by fuzzy controller. As the simulation results show, the introduced fuzzy controller can control the underwater vehicle to avoid moving obstacles, where the least detour in obstacle avoidance shows the robustness and efficiency of fuzzy controller performance. Figure 7 shows that when the vehicle encounter a static obstacle in the path, the controller acts with a robust and avoidance shows the robustness and efficiency of fuzzy stable performance. 40 0 20 40 60 Fig. 5. path planning result Underwater vehicles equipped with an obstacle avoidance controller and property determined regarding the real time navigation environment can safely control in a short way to the target, with least detour. The primary requirement is that the obstacle properties are detected and are sent to the fuzzy controller" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003629_iros.2011.6094853-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003629_iros.2011.6094853-Figure1-1.png", "caption": "Fig. 1. The simplest passive dynamic walker model with hip mass M , weightless legs, and point masses m at the feet.", "texts": [ " In Section III, we observe the escape-time structure of the transient walks in several Poincare\u0301 surfaces. In Section IV, we first introduce SVM and apply it to the escape-time distribution in order to construct the contours of the walking anomalies in each Poincare\u0301 surface. Then, in Section V, we apply CCA to extract a latent space for the bipedal walker. We conclude with a summary and discussion of future work in Section VI. The model explored in this paper is that of the \u201csimplest\u201d model for a biped walking introduced by Garcia et al.[9]. As shown in Fig. 1, it consists of two massless legs connected at the hip, a point mass at the hip, M , and a point mass at each foot, m. This model walks down a rigid surface at a slope \u03b3. The model has no control or actuation and it gains energy from the effect of gravity. The walking step starts with both feet on the slope, and the walking cycle is composed of two phases: a swing phase and a double-stance phase. During the swing phase, the standing foot does not slip, and behaves like a pin joint. When the swinging foot strikes the ground, the double-stance phase appears", " In this paper, we consider the limiting case, where the foot mass is negligibly small in comparison with the hip mass, so that m/M \u2192 0. 1) Equation of motion: swing phase: The equations of motion for the biped walker are written using two second order differential equations that describe the angular motion of the two legs about the foot and hip: \u03b8\u0308(t) \u2212 sin(\u03b8(t) \u2212 \u03b3) = 0, (1) \u03b8\u0308(t) \u2212 \u03c6\u0308(t) + \u03b8\u03072(t) sin \u03c6(t) \u2212 cos(\u03b8(t) \u2212 \u03b3) sin \u03c6(t) = 0, (2) where, \u03b8 is the angle of the stance leg relative to the perpendicular to the inclined plane, and \u03c6 is the angle of the swinging leg relative to the stance leg (see Fig. 1). We denote a point in the state space as x = (\u03b8, \u03b8\u0307, \u03c6, \u03c6\u0307)T. The motion of this model is such that the stance leg can rotate with respect to the floor around its point foot, and the swinging leg can rotate around the hip, i.e., the swing phase is simply a double pendulum. 2) Equation of motion: foot-strike transition: In the case where the floor is smooth and has a constant gradient, the foot-strike occurs when the swinging leg has passed in front of the stance leg, and the inter-leg angle is exactly twice the stance leg angle, i", " This transition rule, (4), provides the initial conditions for the next swing phase, where the names for the legs are exchanged with each other at the moment of this transition. B. Irregularity of Slope It has been shown in [9] that the simplest passive walking model (1)-(4) can walk in a stable cycle on a flat surface and be stable against small disturbances. In this study, we consider the case where the slope of the floor temporally changes irregularly and how the biped walker then becomes unstable, i.e., it eventually falls in time. As indicated in Fig.1, the irregularity for the i-th step is incorporated as \u03b3i = \u03b30 + \u03b4, where \u03b30 is a certain fixed value of a slope and \u03b4 is a Gaussian random variable with zero mean and a standard deviation \u03b7, i.e., \u03b4 \u223c N(0, \u03b72). This randomness adds a small perturbation \u03b4/2 to the foot-strike condition at each step as \u03c6 \u2212 2\u03b8 = \u03b4/2. These perturbations cause the foot-strike to occur either slightly early or slightly late. These perturbations also cause changes in swing speed, such that it is slightly faster or slower during the swing phase" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001108_1.2401215-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001108_1.2401215-Figure4-1.png", "caption": "Fig. 4 Center of gravity", "texts": [ " The carriage of the test bearing is driven at a 2007 by ASME Transactions of the ASME of Use: http://www.asme.org/about-asme/terms-of-use c b L T m s t G w r b t w p v E a t 3 b m b v o n g o s g s A C C C R R R B D N C O C C J Downloaded Fr ertain linear velocity. The pitching and yawing motion of the carriage was detected y using a laser autocollimator Chuo Precision Industrial Co., td: LAC-S and a mirror, and was stored in a personal computer. he positive and negative directions of the pitching and yawing otion of the carriage were defined as shown in Fig. 3. Figure 4 hows the mounting position of the mirror. Because the gravitaional center Gm of the mirror was right above that of the carriage c, the gravitational center G of the mirror attached the mirror as also right above Gc. In the measurement using the laser autocollimator and the miror, the motion caused by the deformation of the rail as well as the all passage vibrations were detected 11 . To avoid the effect of he motion caused by the deformation of the rail, a high pass filter as used. The cutoff frequency was 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002690_978-3-540-30301-5_19-Figure18.11-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002690_978-3-540-30301-5_19-Figure18.11-1.png", "caption": "Fig. 18.11 A cellular force sensor with orthogonal comb drives detailed", "texts": [ " A successful injection is determined greatly by injection speed and trajectory, and the forces applied to cells. To further improve the robotic system\u2019s performance, a multi-axial MEMS-based capacitive cellular force sensor is being designed and fabricated to provide realtime force feedback to the robotic system. The MEMS cellular force sensor also aids research in biomembrane mechanical property characterization. MEMS-Based Multi-Axis Capacitive Cellular Force Sensor The MEMS-based two-axis cellular force sensor [18.129] shown in Fig. 18.11 is capable of resolving normal forces applied to a cell as well as tangential forces generated by improperly aligned cell probes. A highyield microfabrication process was developed to form the 3-D high-aspect-ratio structure using deep reactive ion etching (DRIE) on silicon-on-insulator (SOI) wafers. The constrained outer frame and the inner movable structure are connected by four curved springs. A load applied to the probe causes the inner structure to move, changing the gap between each pair of interdigitated comb capacitors" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001456_s11460-007-0068-x-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001456_s11460-007-0068-x-Figure2-1.png", "caption": "Fig. 2 The characteristics curve of the wind turbine (a) Po = f (v, vw); (b) Tl = f (v, vw)", "texts": [ " The conclusions for the fixed-pitch wind turbine traits are as follows: 1) under rated wind speed, the increase of wind speed for a given wind turbine\u2019s angular velocity will increase its output power, and vice versa; 2) as the wind turbine\u2019s angular velocity changes for a given wind speed, the output power would vary. One specific angular velocity, which is called the optimal angular velocity, can achieve the maximum output power; 3) the optimal angular velocity is unique for a given wind speed. As the wind speed changes, the optimal angular velocity would vary correspondingly. The wind turbine\u2019s torque characteristics are shown in Fig. 2(b). Similarly, the optimal operation points of torque curves at different wind speeds form the optimum torque curve of the wind turbine, which is shown as Topt. Operating on this curve, the wind turbine can capture the maximum wind power. The wind turbine\u2019s torque characteristics are in essence the same as those of the wind turbine\u2019s power; they only reflect the operation characteristics of the wind turbine from different point of views. The maximum output power and the optimal torque of the wind turbine gained from Eqs", " uf and if are the terminal voltage of exciting coil and the exciting current, respectively. ea is electromotive force of armature winding. The steady-state mathematic model of the DC motor is given by Refs. [13\u201315] u i R e e C T C i a a a m dcm dcm e de t dcm = + = wv w= \u23a7 \u23a8 \u23aa \u23a9 \u23aa (4) Where Ce and Ct are the electromotive force constant and torque constant, respectively, and Ce = Ct. w is the main flux linkage. Tde is the electromagnetic torque. vm is the angular velocity. The power characteristics of the wind turbine at various wind speeds, which is shown in Fig. 2(a), can be obtained from Fig. 1 and Eq. (1). As seen in Fig. 2, the peak power for each wind speed occurs at the point where Cp is maximized. If various losses of the DC motor can be ignored, the electric power of the motor is equal to its mechanical output power as follows Pdo = Tdevm (5) From Eqs. (4) and (5) it can be given that P C R u C a m mdo e dcm e= w wv v( )\u2212 (6) If armature reaction can be ignored and w is constant, Pdo is only related to udcm and vm, that is Pdo = f (udcm, vm). And if udcm is a constant, then Pdo = f (vm) is conic. As udcm varies continuously, there is a bunch of conics about Pdo, which is shown in Fig", " By ignoring various losses, the electromagnetic torque of the DC motor is equal to its mechanical output torque as follows: T C R u C a mde t dcm e= w wv( )\u2212 (7) As can be seen in this equation, Tde is also related to udcm and vm, that is Tde = f (udcm, vm). Suppose udcm is a constant, so the relationship between Tde and vm is linear. As udcm varies continuously, a bunch of beelines about Tde can be obtained, as shown in Fig. 4(b). The conclusions contrasted the characteristics of the wind turbine (see in Fig. 2) with those of the DC motor (see in Fig. 4) as follows: 1) Both power characteristics and torque characteristics may be described by a bunch of curves with angular velocity as their abscissa. The change of v may give rise to a variety of wind turbine\u2019s power (torque) characteristics. Similarly, changing udcm can regulate the DC motor\u2019s power (torque) characteristics. 2) While v or udcm are constant, the output power of a wind turbine or a DC motor is directly related to their angular velocity. As a result, as their angular velocity changes, there is an optimal angular velocity at which the maximum output power can be attained" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000900_wcica.2006.1713769-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000900_wcica.2006.1713769-Figure5-1.png", "caption": "Fig. 5 Illustration of the desired angle of neck pan and tilt movements.", "texts": [ " (1) Solving (1), we get 3 5 4 5 4 2 l d Cz a S , and 3 4 4 5 4 2 r d Cz a S . (2) Therefore the coordinates of fixation point in the coordinate frame { 2Z } is 3 1 T L L p p pA A Z x y z , (3) where 3 3 4 5 4 3 3 3 5 4 2 p d C C Cx d S a C S , 3 3 4 5 4 3 3 3 5 4 2 p d S C Cy d C a S S , 3 5 4 5 4 p d S z S . After obtaining the coordinates of fixation point in coordinate frame { 2Z }, the desired neck joint angles can be calculated and [ , , ]T q q qx y z denotes the fixation point in { 0Z }. According to the geometry constraint of the robot head (Fig. 5), the desired neck pan angle is equal to the inclination angle between x2-axis and the line of O2 to fixation point. arctan( )p pan p z x . (4) To make eyes comfortable when head faces the target, eyes tilt angle 3 should be zero. Therefore the desired neck tilt angle can be calculated by 2 4 2 2 arcsin( ) arctan( ) 2 q tilt qq q xd d yx y . (5) After working out the desired neck joint angles, we make neck move towards the desired position within several control cycles. During the process of neck movements, two eyes should make appropriate movements to maintain fixation on the static red ball" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000843_taes.2005.1468753-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000843_taes.2005.1468753-Figure1-1.png", "caption": "Fig. 1. PMLSM servo drive. (a) Machine structure of PMLSM. (b) System configuration of field-oriented control.", "texts": [ " Moreover, the adaptive learning algorithms of the FNN controller and an adaptive bound estimation algorithm are derived in the sense of Lyapunov stability analysis, so that system-tracking stability can be guaranteed in the closed-loop system. In Section IV, simulated and experimental results due to periodic reference trajectories are provided to demonstrate the robust 622 IEEE TRANSACTIONS ON AEROSPACE AND ELECTRONIC SYSTEMS VOL. 41, NO. 2 APRIL 2005 control performance of the proposed control systems. Conclusions are drawn in Section V. The machine structure of the adopted PMLSM is shown in Fig. 1(a). Flexible cables are used to bring 3-phase power to the moving winding. The PMLSM comprises a long stationary tubular \u201csecondary\u201d that is supported at both ends housing a sequence of neodymium-iron-boron (NdFeB) permanent magnet with guidance rail and linear scale, and a moving short \u201cprimary\u201d which contains the core armature winding and Hall sensing elements [22]. The adopted PMLSM is 110 V, 2.9 A, 46 W, 57.8 N type. The configuration of a field-oriented PMLSM drive system is depicted in Fig. 1(b), where dm is the position command; d is the position of the motor; v\u00a4 is the velocity command; v is the linear velocity of the mover; i\u00a4q is the command of thrust current; i\u00a4d is the command of magnetic current; i \u00a4 a, i \u00a4 b, and i\u00a4c are the three-phase command currents; ia and ib are the A and B phase currents; Ta, Tb, and Tc are the switching signals of the inverter. The drive system consists of a PMLSM, a ramp comparison current-controlled pulsewidth modulated (PWM) voltage source inverter (VSI), a field-orientation mechanism, a coordinate translator, a speed control loop, a position control loop, a linear scale, and Hall sensors" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002657_6.2009-6274-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002657_6.2009-6274-Figure8-1.png", "caption": "Fig. 8 illustrates the geometry for the aircraft approaching a desired circular", "texts": [ " Once on a desired circular path, of any radius R 1 , the guidance law will maintain the vehicle on that path. Note that the guidance law does not have explicit knowledge of R 1 . ! horizontal plane, as distinguished from the commanded side acceleration ! a s cmd , ! V L is 10 aircraft velocity in the direction of the reference point and ! V T is the velocity of the reference point along the circular path. Also, as shown in the figure, the angle ! \" determines the position of the vehicle with respect to the desired path because ! L 1 is a fixed distance. Then, referring to Fig. 8 ! V L =V cos\" =V T cos# ! so ! V T = V cos\" cos# and also from Fig. 8 ! V 1 =V sin\" V 2 =V T sin# =V cos\" tan# Angular velocities for various vectors, as shown in the figure, can now be determined as ! \" L = V1 #V2 L1 = V L1 (sin$ # cos$ tan%) \" T = V T R = V cos$ Rcos% \" V = a s V The rates of change of the two angles ! \" and ! \" can be written as ! \u02d9 \" =# L $# V \u02d9 % =# L $# T and, upon substitution for the angular velocities, the following nonlinear differential equations for ! \" and ! \" are obtained ! \u02d9 \" = V L1 (sin# $ cos# tan\") $ V cos# Rcos\" \u02d9 # = V L1 (sin# $ cos# tan\") $ a s V The aircraft is assumed to have a highly effective lateral flight control system that causes it to respond quickly and accurately to lateral acceleration commands" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001487_s0022-0728(78)80377-5-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001487_s0022-0728(78)80377-5-Figure3-1.png", "caption": "Fig. 3. V a r i a t i o n s o f ~ v/iv] a t 25 \u00b0 f o r a c a t a l y t i c r e g e n e r a t i o n w i t h f > 1 ( f = 2, g = - -1 ) . T h e a b s o l u t e v a l u e o f ~, is i n d i c a t e d o n e a c h cu rve .", "texts": [ " (1--3) and (22--24) , we obta in : = Ct ~(1 + ~)x-2[X _ 1 + (g -- 1) ~(1 + ~)] (25) with g = 1/(1 -- f) = (1 \" f ) l = (1 - - f ) R T k / n F v Ct = 1 for the forward scan, and Ct = (1 + ~s)-2x for the backward scan. As f > 1, 1 - - f < 0, so that X is positive for the forward scan (l < 0) and negative for the backward scan (l > 0), g -- 1 is negative. Eqn. (25) is thus formally the same as eqn. 13, but the sign of ~ being opposite to that of l, and g -- 1 being negative instead of being positive, the shape of the curves is quite different (Fig. 3). R E F E R E N C E S 1 E. Laviron, J. Electroanal. Chem., 39 (1972) 1. 2 V. Plichon and E. Laviron, J. Eleetroanal. Chem., 71 (1976) 143. 3 E. Laviron and A. Vallat, J. Electroanal. Chem., 74.(1976) 297. 4 V. Plichon and G. Faure, J. Electr0anal. Chem., 44 (1973) 275. 5 E. Laviron and A. Vallat, J. Electroanal. Chem., 46 (1973) 421. 6 A. VaUat and E. Laviron, J. Electroanal. Chem., 74 (1976) 309. 7 L. Roull ier and E. Laviron, Electrochimr Acta, 21 (1976) 421. 8 J. Bad0z-Lambling and C. Guillaume, Advances in Polarography, Pergamon Press, Oxford, 1960, voL I, p" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000043_s0167-8922(08)70194-9-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000043_s0167-8922(08)70194-9-Figure9-1.png", "caption": "Figure 9 Plots of maximum orthogonal shear stress (T,).", "texts": [], "surrounding_texts": [ "The i n f l u e n c e of l u b r i c a n t c o n t a m i n a t i o n and subsequent s u r f a c e damage on r o l l i n g b e a r i n g f a t i g u e h a s formed t h e b a s i s of s e v e r a l s t u d i e s over r e c e n t y e a r s . Webster e t a 1 (1) c a l c u l a t e d t h e e l a s t i c s u b s u r f a c e stress f i e l d s from r e a l d e n t p r o f i l e s and used t h e s e a s i n p u t t o a f a t i g u e l i f e model t o de te rmine t h e r e d u c t i o n i n l i v e s . I n t h i s paper a s l i p l i n e f i e l d a n a l y s i s h a s been used t o c a l c u l a t e t h e s u b s u r f a c e r e s i d u a l stresses f o r d i f f e r e n t i d e a l i s e d d e n t / r o l l e r combina t ions . These stress f i e l d s were superimposed upon t h o s e c a l c u l a t e d from a d r y c o n t a c t a n a l y s i s of t h e o v e r r o l l i n g of t h e dent which used novel m u l t i - l e v e l t e c h n i q u e s t o a c c e l e r a t e convergence, and t h e r e s u l t a n t stress f i e l d s provided i n p u t t o t h e f a t i g u e l i f e model. The i n f l u e n c e of t h e d e n t i t s e l f on l i f e a p p e a r s t o be s m a l l and o n l y becomes s i g n i f i c a n t on i n c l u s i o n o f t h e r e s i d u a l stresses. These have a p a r t i c u l a r l y marked e f f e c t a s t h e r o l l e r r a d i u s and l o a d a r e reduced, s u g g e s t i n g e x p e c t e d l i v e s may not i n c r e a s e a s r a p i d l y with d e c r e a s i n g l o a d a s would be expec ted from convent iona l models. 1 INTRODUCTION Lubricant contaminat ion and subsequent s u r f a c e damage i s i n c r e a s i n g l y recognised a s having a s i g n i f i c a n t e f f e c t on b e a r i n g f a t i g u e l i v e s which would n o t normally be p r e d i c t e d by t h e t r a d i t i o n a l Lundberg and Palmgren model (2). I n ( 3 ) I o a n n i d e s and H a r r i s p r e s e n t e d a n i m p o r t a n t g e n e r a l i s a t i o n o f t h e c l a s s i c a l model i n which t h e s t ress p e r t u r b a t i o n s r e s u l t i n g from non-per fec t ly smooth c o n t a c t i n g s u r f a c e s can be inc luded . E s s e n t i a l l y t h e new model is a n e l e m e n t a l form o f t h e e a r l i e r a n a l y s i s where t h e p r o b a b i l i t y of f a i l u r e i s c a l c u l a t e d from t h e cumulat ive c o n t r i b u t i o n of stresses above a t h r e s h o l d v a l u e i n s m a l l volume e lements of t h e m a t e r i a l . Therefore t o a p p l y t h e model a comple te h i s t o r y of t h e stress f i e l d underneath t h e d e n t a s it p a s s e s through t h e c o n t a c t i s r e q u i r e d . T h i s f i e l d w i l l be e f f e c t i v e l y made up of two components t h e r e s i d u a l stress f i e l d from t h e i n d e n t a t i o n and t h e EHD stress f i e l d from t h e subsequent o v e r r o l l i n g of t h e d e n t . U n f o r t u n a t e l y t h e s e a r e n o t n e c e s s a r i l y independent as f u r t h e r p l a s t i c deformation of t h e d e n t edges and some shakedown w i l l p r o b a b l y occur , b u t a u s e f u l f i r s t approximat ion can b e made by t r e a t i n g them a s such . T h i s problem was f i r s t t a c k l e d by Webster, I o a n n i d e s and S a y l e s (l), where t h e y measured t h e p r o f i l e s o f d e n t s f rom a r e a l b e a r i n g s u r f a c e . This in format ion was t h e i n p u t t o a numerical c o n t a c t model coupled t o a f i n i t e e lement a n a l y s i s t o d e t e r m i n e t h e s t a t e o f stress i n a n i n d e n t e d b e a r i n g raceway. The stress i n f o r m a t i o n was used i n conjunct ion wi th t h e modif ied f a t i g u e model t o determine t h e r e d u c t i o n i n f a t i g u e l i f e . I n ( 4 ) a p p r o x i m a t e s o l u t i o n s w e r e p r e s e n t e d f o r t h e i n t e r f a c i a l p r e s s u r e s and d e f l e c t i o n s r e s u l t i n g from debris p a r t i c l e s b e i n g squashed i n t h e i n l e t t o an EHD c o n t a c t . A p l a n e s t r a i n approach was used t o a l l o w t h e a p p l i c a t i o n of 2 d imens iona l e x t r u s i o n t h e o r y . The problem i s s i m i l a r t o s t r i p r o l l i n g , b u t i n t h i s c a s e t h e r o l l e r s cannot b e e x p r e s s e d a s c i r c u l a r a r c s a s t h e e l a s t i c d e f l e c t i o n s due t o t h e EHD and e x t r u s i o n p r e s s u r e s a r e s i g n i f i c a n t . The s o l u t i o n was f o u n d by numer ica l ly i t e r a t i n g between t h e p r e s s u r e and d e f l e c t i o n e q u a t i o n s u n t i l convergence. As t h e r o l l i n g s u r f a c e s c o n t i n u e t o a p p r o a c h one a n o t h e r i n t h e i n l e t , t h e i n t e r f a c i a l p r e s s u r e s may i n c r e a s e s u f f i c i e n t l y such t h a t p l a s t i c deformat ion of t h e raceways can o c c u r . The c r i t i c a l v a l u e of p r e s s u r e may b e e s t i m a t e d from i n d e n t a t i o n exper iments or t h e o r e t i c a l c o n s i d e r a t i o n s such a s ( 4 ) . By a p p l y i n g t h e s e c r i t e r i a t o t h e model , t h e a p p r o x i m a t e d e n t s h a p e s a n d p r e s s u r e d i s t r i b u t i o n s may b e f o u n d a n d compared w i t h t h o s e measured e x p e r i m e n t a l l y ( 5 ) . The measured p r o f i l e s o f t h e s e d e n t s c l o s e l y approximate a c i r c u l a r a r c s i m i l a r t o t h e i m p r e s s i o n l e f t by a r i g id c y l i n d r i c a l i n d e n t e r . Dumas and Baronet ( 6 ) produced a f i n i t e e l e m e n t s o l u t i o n t o t h e c i r c u l a r i n d e n t e r problem and found t h a t a t s i g n i f i c a n t d e p t h s of d e p r e s s i o n , where t h e s u b s u r f a c e zone w a s e n t i r e l y p l a s t i c t h e c a l c u l a t e d p r e s s u r e p r o f i l e w a s f a i r l y c o n s t a n t o v e r most o f t h e i n t e r f a c e . A t s m a l l e r d e p t h s o f d e p r e s s i o n t h e p r e s s u r e p r o f i l e i s n o t f l a t , b u t r e a c h e s a maximum a t t h e c e n t r e where t h e p r e s s u r e i s approximate ly 50% g r e a t e r t h a n it i s n e a r t h e p e r i p h e r y o f t h e d e n t . T h i s s u g g e s t s t h a t f o r r e l a t i v e l y s m a l l i n d e n t a t i o n s t h e i n t e r f a c i a l p r e s s u r e i s s t i l l i n c r e a s i n g t o w a r d s t h e c e n t r e of t h e d e n t , t h u s g e n e r a t i n g a r a d i a l e x t r u s i o n f o r c e on t h e d e b r i s p a r t i c l e r e s u l t i n g i n a s u r f a c e t r a c t i v e f o r c e on t h e b e a r i n g raceway. A t g r e a t e r d e p t h s o f i n d e n t a t i o n t h e p r e s s u r e p r o f i l e f l a t t e n s and t h e r e s u l t a n t s u r f a c e t r a c t i o n f o r c e d i m i n i s h e s . O u t s i d e t h e i n d e n t e d zone where t h e d e f l e c t i o n s a r e e n t i r e l y e l a s t i c , d e b r i s d e f o r m a t i o n w i l l c o n t i n u e t o o c c u r and t h e p r e s s u r e p r o f i l e w i l l b e d e f i n e d by t h e e x t r u s i o n a n d e l a s t i c i t y e q u a t i o n s a s d i s c u s s e d i n ( 4 ) . The r e s i d u a l stresses r e s u l t i n g from t h e i n d e n t a t i o n e s p e c i a l l y i n t h e p r e s e n c e of a s u r f a c e t r a c t i o n f o r c e may encourage f a t i g u e c rack i n i t i a t i o n a s p o s t u l a t e d by Olver ( 7 ) . I n t h i s paper a s l i p l i n e f i e l d model based on t h a t deve loped by Olver was used t o e s t i m a t e t h e p l a s t i c stresses i n t h e s u b s u r f a c e and t h e r e s u l t a n t r e s i d u a l stresses on unloading. A f t e r i n d e n t a t i o n t h e s q u a s h e d d e b r i s p a r t i c l e i s l o s t and t h e damaged s u r f a c e r e p e a t e d l y p a s s e s t h r o u g h t h e EHD c o n t a c t . Model l ing t h i s p r o c e s s i s n o t p r a c t i c a b l e a t p r e s e n t , however a s t h e f i l m t h i c k n e s s i s s m a l l compared t o t h e e l a s t i c deformat ion of t h e s u r f a c e and w i l l p r o b a b l y b e f u r t h e r reduced around t h e edges of t h e d e n t , a d r y c o n t a c t a p p r o x i m a t i o n may n o t b e t o o u n r e a l i s t i c . C l e a r l y , a l t h o u g h t h e r e s i d u a l stress d i s t r i b u t i o n i s f i x e d r e l a t i v e t o t h e d e n t t h e super imposed c o n t a c t stress f i e l d w i l l b e changing d u r i n g o v e r r o l l i n g and t h e f u l l h i s t o r y w i l l be r e q u i r e d f o r t h e l i f e c a l c u l a t i o n s . 1.1 Notation A c o n s t a n t b ha l f -wid th of Her tz ian c o n t a c t , m dB incrementa l a r e a of t h e c r o s s - s e c t i o n of t h e r i n g , mm2 e l i f e exponent H d imens ionless f i l m t h i c k n e s s H O O c o n s t a n t i n d imens ionless f i l m k y i e l d s h e a r stress, N/m2 N f a t i g u e l i f e P dimens ionless p r e s s u r e R reduced r a d i u s of c u r v a t u r e , m S p r o b a b i l i t y of s u r v i v a l x , x ' c o o r d i n a t e s i n r o l l i n g d i r e c t i o n , m X , X ' d imens ionless c o o r d i n a t e s y c o o r d i n a t e i n a x i a l d i r e c t i o n , m z c o o r d i n a t e normal t o t h e s u r f a c e , m z ' stress weighted average depth , mm R domain a' extended domain t h i c k n e s s equat ion 2 SLIPLINE FIELD I n t h e s l i p l i n e f i e l d a n a l y s i s t h e d e b r i s i n d e n t a t i o n of t h e b e a r i n g s u r f a c e i s model led by a r i g i d d ie i m p r e s s i n g a r i g i d - p l a s t i c h a l f s p a c e c o u p l e d w i t h a t r a c t i o n f o r c e e x t e n d i n g from t h e c e n t r e l i n e . I f t h e t r a c t i o n f o r c e i s z e r o t h i s r e d u c e s t o t h e c l a s s i c a l f i e l d f o r a r i g i d i n d e n t e r a s p r e s e n t e d by H i l l ( 9 ) i n t h e deforming zone. I n ( 7 ) Olver p r e s e n t e d a s o l u t i o n f o r a s i n g l e u n i - d i r e c t i o n a l t r a c t i o n f o r c e a p p l i e d t o t h e i n t e r f a c e , f i g u r e 1. A s t h e s l i p l i n e s w i l l n o t m e e t t h e i n t e r f a c e a t 45' a n e x t r a f a n must be drawn i n ( r e g i o n BCF). I n t h e r e g i o n DCFE t h e l i n e s w e r e c a l c u l a t e d n u m e r i c a l l y by a f i n i t e d i f f e r e n c e t e c h n i q u e proposed by H i l l ( 9 ) and t h e e l a s t i c b o u n d a r y ( B E ) was d e r i v e d a n a l y t i c a l l y by Johnson ( 8 ) . I f t h e t r a c t i o n f o r c e i s made e q u a l t o z e r o t h e s l i p l i n e s m e e t t h e s u r f a c e a t 45O and t h e f i e l d reduces t o t h a t proposed f o r a p u r e l y normal i n d e n t a t i o n . If t h e t r a c t i o n f o r c e e x t e n d s r a d i a l l y f rom t h e c e n t r e t h e n t h e zone below t h e c e n t r e l i n e w i l l be p l a s t i c and t h e s l i p l i n e s m u s t m e e t a t 90\u00b0, f i g u r e 2 . From symmetry a s q u a r e s e c t i o n BEFG can be drawn and as b e f o r e a f a n i s c o n s t r u c t e d between BC and BE. The s l i p l i n e s i n r e g i o n s CDHE and EHIF c a n b e c a l c u l a t e d n u m e r i c a l l y by f i n i t e d i f f e r e n c e s . Although t h e s l i p l i n e f i e l d could be ex tended i n d e f i n i t e l y t h e a c t u a l zone o f p l a s t i c i t y w i l l o n l y e x i s t d i r e c t l y below t h e i n d e n t e r . D e f i n i n g t h e e l a s t i c / p l a s t i c boundary e x a c t l y i s n o t p o s s i b l e , b u t an approximat ion c a n b e made by a p p l y i n g t h e i n t e r f a c e p r e s s u r e s t o an e l a s t i c h a l f s p a c e and c a l c u l a t i n g where t h e y i e l d s h e a r stress is exceeded, a s s u g g e s t e d by Olver ( 7 ) . S i g n i f i c a n t p l a s t i c deformat ion w i l l p r o b a b l y o n l y o c c u r i n t h e r e g i o n OADCB, whereas i n t h e remainder o f t h e f i e l d t h e m a t e r i a l may be p l a s t i c b u t w i l l no t deform. Having c o n s t r u c t e d t h e f i e l d t h e stresses may b e c a l c u l a t e d i n t h e normal manner see ( 7 ) . During unloading , a t e n s i l e normal f o r c e e q u a l i n magnitude t o t h e p l a s t i c p r e s s u r e s i s e f f e c t i v e l y a p p l i e d t o t h e i n t e r f a c e . C o n s e q u e n t l y t h e r e s i d u a l stresses c a n b e found by summing t h e stress d i s t r i b u t i o n r e s u l t i n g from t h i s e l a s t i c t e n s i l e f o r c e t o t h e p l a s t i c stress d i s t r i b u t i o n . The e l a s t i c s tresses w e r e c a l c u l a t e d a n a l y t i c a l l y a s d e s c r i b e d by Ford ( 1 0 ) . A map of t h e p r i n c i p a l r e s i d u a l s tress d i s t r i b u t i o n s f o r t r a c t i o n c o e f f i c i e n t s of 0 and 0 . 2 are shown i n f i g u r e s 3 ( a ) a n d ( b ) a n d a c o n t o u r map o f t h e o r t h o g o n a l r e s i d u a l s h e a r stresses i n f i g u r e ( 4 ) . The e f f e c t of t h e s u r f a c e s h e a r stress is t o r e d u c e t h e t e n s i l e r e s i d u a l s tresses immediately below t h e s u r f a c e b u t t o i n c r e a s e t h o s e below t h e dent s h o u l d e r . 3 DRY CONTACT EQUATIONS To s o l v e t h e d r y c o n t a c t p r o b l e m i t i s n e c e s s a r y t o f i n d a p r e s s u r e d i s t r i b u t i o n P ( X ) , t h a t s a t i s f i e s t h e f i l m t h i c k n e s s e q u a t i o n ( l ) , i e s o l v e t h e i n t e g r a l e q u a t i o n ( 2 ) f o r t h e p r e s s u r e P . H ( X ) = 0 XE a (1) i n t h e one d i m e n s i o n a l c a s e , t h i s e q u a t i o n r e a d s : H, + - - - P(X') K(X , X') dX' = 0 XE a XL 2 I t ' I n (2) where K ( X , X ' )=In 1 X-X' I U n f o r t u n a t e l y t h e domain on which e q u a t i o n (1) h o l d s i s g e n e r a l l y n o t known i n a d v a n c e . Therefore (1) is extended t o ( 3 ) such t h a t (1) h o l d s i n t h e o r i g i n a l domain and P ( X ) = O X E ~ ' P ( X ) > 0 H ( X ) = 0 XE a P ( X ) = 0 H ( X ) > 0 X E R ' ( 3 ) 191 A second c o n d i t i o n t h a t has t o be s a t i s f i e d i s t h e f o r c e b a l a n c e e q u a t i o n : b ( X ) d x = f n ( 4 ) The i n t e r a c t i o n of t h e s e e q u a t i o n s i s i d e n t i c a l t o t h e l u b r i c a t e d c o n t a c t case , see f o r i n s t a n c e (11). I n t h i s d r y c o n t a c t a n a l y s i s t h e f i l m t h i c k n e s s i s set t o zero over t h e domain (a ) and t h e n o n l y t h e p r e s s u r e d i s t r i b u t i o n has t o b e c a l c u l a t e d . The problem h a s t r a d i t i o n a l l y b e e n s o l v e d b y d i r ec t m e t h o d s (Newton Raphson) , r e s u l t i n g i n l o n g computing t i m e s f o r l a r g e problems. AS a l a r g e number o f s o l u t i o n s i s r e q u i r e d i n a p p l y i n g t h e l i f e model t h e c o m p u t i n g t i m e c a n become e x c e s s i v e l y long . To a l l e v i a t e t h i s problem a n a l t e r n a t i v e i t e r a t i v e t y p e s o l u t i o n was used which h a s been d e s c r i b e d i n (12) . Convergence w a s f u r t h e r a c c e l e r a t e d by t h e a p p l i c a t i o n of novel m u l t i - l e v e l t e c h n i q u e s which have been a p p l i e d p r e v i o u s l y t o t h e s o l u t i o n o f d i f f e r e n t i a l e q u a t i o n s . The computing t i m e f o r t h e u s u a l i t e r a t i v e s o l u t i o n t e n d s t o be dominated by t h e i n t e g r a l c o m p u t a t i o n , s o t h e s o l u t i o n t i m e i s p r o p o r t i o n a l t o o r d e r n2 where n i s t h e number of p o i n t s , a s can b e seen i n Table 1. However it i s p o s s i b l e t o reduce t h e comput ing , t ime t o o r d e r n logn by t h e a p p l i c a t i o n of m u l t i l e v e l t e c h n i q u e s , s p e c i f i c a l l y M u l t i l e v e l M u l t i - I n t e g r a t i o n ( M L M I ) . The b a s i c i d e a was g i v e n i n (13) and worked out i n d e t a i l i n ( 1 2 ) . A s can be seen from Table 1, column 3 s i g n i f i c a n t t i m e s a v i n g s c a n be o b t a i n e d from l e v e l 6 onwards, s o t h e approach i s most u s e f u l f o r p r o b l e m s w i t h many g r i d p o i n t s . The c a l c u l a t i o n of t h e s u b s u r f a c e stresses i s a t a s k s i m i l a r t o t h e f i l m t h i c k n e s s s o l u t i o n , when one c o n s i d e r s o n l y one p a r t i c u l a r v a l u e of t h e d e p t h z, a t a t i m e . P l o t s of t h e d r y c o n t a c t s u r f a c e p r e s s u r e a n d a s s o c i a t e d s u b s u r f a c e o r t h o g o n a l s h e a r s t r e s s d i s t r i b u t i o n are shown i n f i g u r e 5 . 4 L I F E PREDICTIONS I n t h e t r a d i t i o n a l Lundberg a n d Pa lmgren b e a r i n g f a t i g u e l i f e model, t h e p r o b a b i l i t y of f a i l u r e c a n be e x p r e s s e d i n t e r m s o f t h e stressed volume and t h e magnitude and d e p t h below t h e s u r f a c e of t h e maximum o r t h o g o n a l s h e a r stress. I n ( 2 ) I o a n n i d e s and Har r i s p r o p o s e d a g e n e r a l i s e d model i n which t h e s t r e s s e d volume i s d i v i d e d i n t o d i s c r e t e volume e lements i n which t h e maximum stress i s c a l c u l a t e d a c c o r d i n g t o some stress r e l a t e d f a t i g u e c r i t e r i o n . I n common w i t h s t r u c t u r a l f a t i g u e l i f e p r e d i c t i o n s f o r s teels i n r e v e r s e d b e n d i n g or t o r s i o n , a t h r e s h o l d stress v a l u e i s d e f i n e d below which f a i l u r e w i l l n o t o c c u r . Each e l e m e n t i s w e i g h t e d a c c o r d i n g t o i t s depth below t h e s u r f a c e and t h e p r o b a b i l i t y o f f a i l u r e i s e x p r e s s e d i n terms o f t h e i n t e g r a l o f t h e e l e m e n t a l stresses over t h e e n t i r e volume. A modif ied l i f e c r i t e r i o n h a s been used t o compute t h e Ll0 b e a r i n g l i v e s i n t h e p r e s e n t c a s e . The maximum s h e a r stress ampl i tude z, i s c a l c u l a t e d d u r i n g t h e o v e r r o l l i n g of t h e d e n t . I n common w i t h s t r u c t u r a l f a t i g u e , t h e f a t i g u e stress t h r e s h o l d z, i s m o d i f i e d a c c o r d i n g t o t h e a b s o l u t e v a l u e of t h e s h e a r stress. z, i s assumed t o remain unchanged i f T,, d o e s n o t e x c e e d t h e y i e l d s t ress a n d t o d i m i n i s h l i n e a r l y t o z e r o f o r Tmax v a r y i n g between r e a n d t h e f r a c t u r e s t r e n g t h zf. A s t h e c rack might be e x p e c t e d t o b e c r e a t e d more e a s i l y i n t h e p r e s e n c e of a t e n s i l e r a t h e r t h a n compressive s t ress f i e l d , a n a d d i t i o n a l h y d r o s t a t i c weight ing was i n c l u d e d i n t h e model. I n t h i s t h e c r i t i c a l stress Ta was modi f ied to fa+ a.Hp, where Hp is e q u a l t o t h e h y d r o s t a t i c p r e s s u r e and a i s t a k e n as a=0.3.Using t h e s e v a l u e s , t h e p r o b a b i l i t y of s u r v i v a l of t h e i n n e r r i n g can be e x p r e s s e d a s : The e f f e c t i v e p e r t u r b a t i o n on t h e g l o b a l p r e s s u r e d i s t r i b u t i o n by t h e d e n t w i l l depend v e r y much on t h e r a t i o of t h e d e n t wid th t o t h e Her tz c o n t a c t s i z e . To a s s e s s t h i s e f f e c t f o u r d e n t / r o l l e r combinat ions w e r e chosen . An a r t i f i c i a l c i r c u l a r d e n t of 200 micron wid th and 3 micron d e p t h was o v e r r o l l e d by r o l l e r s of 2 , 4 , 8 and 16mm r a d i u s . The o v e r r o l l i n g of t h e d e n t is s i m u l a t e d u s i n g 9 d i f f e r e n t p o s i t i o n s o f t h e r o l l i n g element wi th r e s p e c t t o t h e d e f e c t , i n o r d e r t o p i c k up t h e maximum stresses. The p o s i t i o n of t h e c e n t r e x c of t h e r o l l i n g element i s g i v e n by: xc = b (n-5) /2 f o r n=1,2 ,..., 9 . The stress h i s t o r y i n e a c h p o i n t i s a n a l y s e d wi th r e s p e c t t o t h e s e n i n e p o s i t i o n s and t h e n t h e l i f e i n t e g r a l i s c a l c u l a t e d . A s t h e number of p o s i t i o n s i n t i m e and space a r e r e l a t i v e l y smal l , 9 and 49x17 r e s p e c t i v e l y , t h e v a l u e s of t h e l i f e i n t e g r a l s a re r a t h e r jumpy a n d c o n s e q u e n t l y t h e numer ica l r e s u l t s s h o u l d be i n t e r p r e t e d w i t h c a r e . 5 RESULTS The l i f e i n t e g r a l s were c a l c u l a t e d f o r f o u r d i f f e r e n t v a l u e s o f t h e r e d u c e d r a d i u s o f c u r v a t u r e a n d f o r s i x d i f f e r e n t l o a d s ( c o r r e s p o n d i n g t o H e r t z i a n p r e s s u r e s r a n g i n g f rom 2 . 0 t o 3 . 3 GPa) . T h r e e cases were examined, a smooth raceway, a raceway wi th one d e n t a n d a r a c e w a y w i t h o n e d e n t a n d a s s o c i a t e d r e s i d u a l stress f i e l d . The e f f e c t of t h e s e stress f i e l d s on p r e d i c t e d l i v e s c a n be g r a p h i c a l l y e x p r e s s e d i n t e r m s of r i s k maps. I n t h e s e a s e c t i o n of t h e x, z p l a n e i s drawn on a gr id w i t h t h e ' f a t i g u e c r i t e r i o n ' stress e x p r e s s e d as t h e y c o - o r d i n a t e . Each map i s n o r m a l i s e d t o t h e smooth c a s e by a s c a l i n g f a c t o r . The smooth c a s e r i s k map i s shown i n f i g u r e 6 ( a ) where a s e x p e c t e d t h e h i g h e s t r i s k o c c u r s a t t h e p o s i t i o n of t h e maximum o r t h o g o n a l s h e a r stress, 0 .8b below t h e b e a r i n g s u r f a c e . AS t h e d e n t ( f i g u r e 6 ( b ) ) and t h e d e n t p l u s r e s i d u a l stresses ( f i g u r e 6 ( c ) ) are i n c l u d e d t h e map i s m o d i f i e d , 192 p a r t i c u l a r l y a round t h e d e n t s h o u l d e r s . A s a r e s u l t t h e s c a l e f a c t o r , e f f e c t i v e l y a measure of t h e i n c r e a s e d r i s k , i n c r e a s e s d r a m a t i c a l l y , by a f a c t o r of a lmost 50 on t h e i n c l u s i o n of t h e r e s i d u a l stresses. The p r e d i c t e d l i v e s f o r each d e n t / r o l l e r combination a r e p l o t t e d i n f i g u r e 7 . From t h i s d a t a an approximate map of r e l a t i v e l i v e s can b e c o n s t r u c t e d i n t e r m s of t h e d e n t s i z e , c o n t a c t s i z e and r o l l e r r a d i u s o f c u r v a t u r e ( f i g u r e 8). A s can be s e e n t h e l i f e of t h e smooth r a c e w a y i n c r e a s e s r a p i d l y w i t h d e c r e a s i n g l o a d a n d t h e l i f e g e n e r a l l y i n c r e a s e s w i t h i n c r e a s i n g r a d i u s . The i n f l u e n c e of t h e dent wi thout t h e a s s o c i a t e d r e s i d u a l stresses on l i f e i s m i n i m a l . A s i g n i f i c a n t l i f e r e d u c t i o n o n l y occurs f o r t h e s m a l l e s t r a d i u s u n d e r t h e t h r e e l i g h t e s t l o a d s . T h i s c h a n g e s d r a m a t i c a l l y when t h e ( t e n s i l e ) r e s i d u a l stress f i e l d below t h e d e n t i s t a k e n i n t o a c c o u n t . The r e s i d u a l stresses were o b t a i n e d assuming no r a d i a l t r a c t i o n f o r c e a t t h e i n t e r f a c e , i e a f l a t p r e s s u r e d i s t r i b u t i o n . Under h i g h l o a d s t h e i n f l u e n c e o f t h e r e s i d u a l stress f i e l d s a r e r e l a t i v e l y s m a l l , b u t a s t h e l o a d i s reduced t h e l i v e s d e c r e a s e m a r k e d l y r e l a t i v e t o t h e smooth c a s e s . A s t h e r a d i u s o f t h e c o n t a c t i s i n c r e a s e d t h e i n f l u e n c e o f t h e r e s i d u a l s t resses a n d o f t h e d e n t g e o m e t r y i s diminished. I n s p e c t i o n o f t h e o r t h o g o n a l s h e a r stress c o n t o u r s ( f i g u r e 9) h e l p s e x p l a i n why t h i s might b e s o . A t h i g h e r l o a d s a l t h o u g h t h e o r t h o g o n a l s h e a r stresses of t h e m o d i f i e d H e r t z i a n f i e l d a r e of a h i g h e r magnitude, t h e maxima a r e s i t u a t e d w e l l below t h e s u r f a c e . The stress c o n c e n t r a t i o n s from t h e shoulder of t h e d e n t and p a r t i c u l a r l y t h e t e n s i l e stresses of t h e r e s i d u a l stress f i e l d l i e much more c l o s e l y t o t h e s u r f a c e and c a n n o t t h e r e f o r e combine w i t h them t o g e n e r a t e h i g h v a l u e s of t h e f a t i g u e c r i t e r i o n . A s t h e l o a d i s reduced t h e stress c o n t o u r s l i e more c l o s e l y t o t h e s u r f a c e a n d c a n combine w i t h t h e t e n s i l e r e s i d u a l stresses t o cause more damage. The r e s u l t s s h o u l d be i n t e r p r e t e d w i t h some c a u t i o n because of t h e s i m p l i f i c a t i o n s made and t h e c o a r s e g r i d numerics and s t r ic t q u a n t i t a t i v e c o n c l u s i o n s s h o u l d n o t b e made. However , t h e q u a l i t a t i v e r e s u l t s a r e i n t e r e s t i n g enough t o c o n t i n u e t h e r e s e a r c h i n t h i s d i r e c t i o n , i n c o r p o r a t i n g more r ea l i s t i c d e n t shapes , r e s i d u a l stress f i e l d s and more a c c u r a t e c a l c u l a t i o n s on f i n e r g r i d s . 6 CONCLUSIONS The e f f e c t of b o t h d e n t s i z e and s u b s u r f a c e r e s i d u a l stresses have been added t o t h e o r i g i n a l l i f e r e d u c t i o n work. The s l i p l i n e f i e l d a n a l y s i s c a n o n l y b e r e a l i s t i c a l l y a p p l i e d t o deep t r a n s v e r s e i n d e n t a t i o n s where t h e assumptions of p l a n e p l a s t i c s t r a i n can be j u s t i f i e d and t h e d r y c o n t a c t a n a l y s i s i s p r o b a b l y o n l y r e a s o n a b l e f o r r e l a t i v e l y t h i n f i l m c o n d i t i o n s . However t h e t r e n d s i n t h e l i f e r e d u c t i o n f a c t o r s are d i s t i n c t i v e and t h e e x p l a n a t i o n f o r them would s e e m reasonable and a p p l i c a b l e t o a n y d e n t p r o f i l e / r o l l e r combinat ion. The most s t r i k i n g outcome i s t h a t where f a i l u r e i s i n i t i a t e d t h r o u g h s u r f a c e i n d e n t a t i o n and a s s o c i a t e d r e s i d u a l stresses (and c o n s i d e r a b l e e v i d e n c e e x i s t s t o s u g g e s t t h i s i s s o ) t h e e x p e c t e d l i v e s may n o t i n c r e a s e w i t h d e c r e a s i n g l o a d a s r a p i d l y a s would b e p r e d i c t e d by convent iona l models. The r e d u c t i o n i n expec ted l i v e s i s v e r y s e n s i t i v e t o t h e s i z e of d e n t i n r e l a t i o n t o t h e r o l l e r r a d i u s a n d t h i s may w e l l have i m p o r t a n t consequences i n t e r m s o f c r i t i c a l p a r t i c l e s i z e and s a f e and u n s a f e l e v e l s of f i l t r a t i o n . To b e a b l e t o draw q u a n t i t a t i v e c o n c l u s i o n s f u r t h e r r e s e a r c h i s needed t h a t uses more r e a l i s t i c r e s i d u a l stress f i e l d s and f i n e r grids i n t h e l i f e c a l c u l a t i o n s . 7 ACKNOWLEDGEMENTS W e would l i k e t o thank D r Andrew Olver f o r h i s h e l p f u l a d v i c e i n t h i s work and t o register our g r a t i t u d e t o SKF-ERC, The Nether lands, who have s p o n s o r e d t h i s work, and t o D r I a n Leadbetter, Managing D i r e c t o r of SKF-ERC f o r permiss ion t o p u b l i s h . APPENDIX References 31 4 1 91 Webster, M . N . , Ioannides , E . and S a y l e s , R . S . , ( 1 9 8 5 1 , \"The E f f e c t o f T o p o g r a p h i c a l D e f e c t s on t h e C o n t a c t Stress and F a t i g u e L i f e i n R o l l i n g Element Bearings\" , Proceedings of t h e 1 2 t h LeedsLyon Symposium on T r i b o l o g y , Lyon, But te rwor ths , Vo1. 12, pp. 121-131. Lundberg, G . and Palmgren, A . , ( 1 9 4 7 ) , \"Dynamic C a p a c i t y o f R o l l i n g Bear ings \", Acta P o l y t e c h n i c a , Mechanical Engineer ing series , Royal Academy o f E n g i n e e r i n g Sc iences , Vol. 1, No 3, 7 . I o a n n i d e s , E . and H a r r i s , T . A. , (1985) , \"A N e w F a t i g u e L i f e Model f o r R o l l i n g B e a r i n g s \", ASME J o u r n a l of L u b r i c a t i o n Technology, , Vol. 107, pp. 367-378. Hamer, J . C . , Sayles , R. S . and Ioannides , E . , ( 1 9 8 5 ) , \"Deformation Mechanisms and S t r e s s e s C r e a t e d by 3 r d Body D e b r i s C o n t a c t s and T h e i r E f f e c t s on R o l l i n g Bear ing F a t i g u e \", Proceedings of t h e 1 4 t h Leeds-Lyon Symposium on Tr ibology, Lyon, But te rwor ths , Vol. 1 4 . Hamer, J. C . , Sayles , R. S. and Ioannides , E., \" P a r t i c l e Deformation and Counter face Damage When R e l a t i v e l y S o f t P a r t i c l e s a r e S q u a s h e d Between Hard A n v i l s \" , T r a n s ASME/STLE t o be p u b l i s h e d . Dumas, G . and B a r o n e t , C . N . , ( 1 9 7 1 ) , \" E l a s t o - p l a s t i c i n d e n t a t i o n of a h a l f - s p a c e b y a l o n g r i g i d c y l i n d e r \" , I n t e r n a t i o n a l J o u r n a l o f M e c h a n i c a l Sc iences , Vol. 13, 519. Olver , A . V . , (19861, \"Wear of Hard S t e e l i n L u b r i c a t e d , R o l l i n g C o n t a c t \" , Phd Thes is , I m p e r i a l Col lege . Olver , A . V . , Sp ikes , H . A . , Bower, A . and Johnson, K . L . , ( 1 9 8 6 ) , \"The R e s i d u a l S t r e s s D i s t r i b u t i o n i n a P l a s t i c a l l y Deformed Model Asper i ty\" , Wear, Vo1. 107, H i l l , R . , (1950) , \"The Mathematical Theory pp. 151-174. o f P l a s t i c i t y \", Oxford U n i v e r s i t y Press. 1 0 1 Ford , H . , ( 1 9 6 3 ) , \"Advance Mechanics of Mater ia l s\" , Longmans . 113 Lubrecht , A . A . , \"The Numerical S o l u t i o n of t h e Elas tohydrodynamica l ly L u b r i c a t e d L i n e and P o i n t C o n t a c t Problem, Using M u l t i g r i d Techniques\", Phd Thes is , Twente U n i v e r s i t y , l 9 8 7 , The Nether lands . 193 [121 Brandt, A. and Lubrecht, A. A., 'Multilevel Multi-Integration and Fast Solution of Integral Equations\", to be published in the Journal of Computational Physics. [131 Brandt, A. , \"Multigrid Techniques: 1984 Guide with Applications to Fluid Dynamics\", monograph available as GMD studien 85 from GMD postfach 1240 , Schloss Birlinghofen D5205 St. Augustin 1 BRD. coefficient (p) of: (a) p = 0 and (b) p = 0.2. %/A 000 ? Figure 4 Residual orthogonal shear stress (Q contour map. 194 shear stress (2,) distribution during the overrolling of the dent. 195 Figure 6(c) Risk map of indented surface plus residual stresses. Scale factor49.7 196 stresses. 197 n 9 17 33 65 129 257 513 1025 time 0.8 1.5 2.4 3.9 8.4 23.0 79.0 306.0 time * 6.2 12.3 24.3 48.4 Table 1: Computing time as a function of the level (L), the number of points n, in seconds on a VAX 785, for the smooth dry line contact problem, with (*) and without MLMI." ] }, { "image_filename": "designv11_20_0002904_1077546309104878-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002904_1077546309104878-Figure3-1.png", "caption": "Figure 3. Bridge irregularity AB.", "texts": [ " Assumptions In addition to the assumptions of Section 2.1.1, the following ones are considered: 1. We assume that on the deck of a single-span beam there is the irregularity AB of length e, that starts at point A (x xA) and ends at point B (x xA e). The above irregularity has a form given by equation 1. 2. A mass load like the one described in Section 2.1 moves on the beam with constant velocity. 3. The mass M, before its entering onto the beam, is moving on the level a\u2013a, from which are measured its displacements (Figure 3a). 4. Moreover, we assume that the mass load enters the irregularity normally, i.e. without the appearance of impact forces. 5. Finally, we assume that the irregularity does not affect the beam\u2019s characteristics (I and m), and that the critical speeds 1cr and 2cr can be determined by using equations 13 and 16, respectively. 2.2.2. Mathematical Formulation The equations of Section 2.1.2 are valid, considering that the deformations of the beam, compared to the form of the irregularity, are very small", "2, one can establish the governing differential equation of the motion for a slender beam. This equation, after neglecting the effect of longitudinal motion and using the Dirac function, is given by E Iy x t c x t m x t F x a (19) In equation 19, the prime denotes differentiation with respect to x, while the dot denotes differentiation with respect to t. The associated boundary and initial conditions are respectively 0 t t 0 t t 0 x 0 x 0 0 (20) Through a process like the one of Section 2.1.2, and taking into account Figure 3, we have F M g z mo[g o ] M g z mog mo (21a) at SIMON FRASER LIBRARY on November 17, 2014jvc.sagepub.comDownloaded from On the other hand, cutting at point G (Figure 3b) and taking into account the equilibrium of forces, we get M z ko[z o ] co[ z o ] koz ko ko o co z co which leads to z 2 o z 2 oz 2 o o 2 o (21b) at SIMON FRASER LIBRARY on November 17, 2014jvc.sagepub.comDownloaded from where o and o are given in equation 5, and o x 0 for 0 x xo or x xo e o x o for xo x xo e (21c) The solution of equation 21b, for initial conditions z 0 z 0 0, is given by Duhamel\u2019s integral: z t 1 o t 0 e o t 2 o x 2 o o 2 o x sin o t d (22a) The Leibnitz formula for G x x 0 Q x t dt gives dG x dx Q x x x 0 Q x t x dt Because of the above, equation 22a becomes z t 1 o t 0 e o t 2 o x 2 o o 2 o x [ o cos o t o sin o t ]d (22b) Thus, the load F according to equation 21a becomes F M mo g t2 t1 x t where t2 t1 ko o t2 t1 e o t 2 o x 2 o o 2 o x sin o t d co o t 0 e o t 2 o x 2 o o 2 o x [ o cos o t o sin o t ]d ko x t ko o x co x t mo x t (22c) Introducing the load F, from equation 22c, into equation 19, we get the following integral\u2013 differential equations: at SIMON FRASER LIBRARY on November 17, 2014jvc" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002859_20090630-4-es-2003.00151-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002859_20090630-4-es-2003.00151-Figure1-1.png", "caption": "Fig. 1. Measured signals of the vehicle suspension system for indirect tire pressure monitoring", "texts": [ " However, battery for power supply and wireless data transmission increase the complexity and costs of such systems. Due to extreme environmental conditions like a large operating range of temperature and high accelerations the sensor unit in the tire has to be very robust. Therefore, alternatives like indirectly measuring systems are desired. Indirectly measuring systems estimate the tire pressure using sensor signals of the wheel or the suspension which may be already available for other vehicle dynamic control systems and are influenced by the tire pressure. As shown in Figure 1 the wheel speed \u03c9 and the vertical wheel acceleration can be used for indirect tire pressure monitoring. First results obtained with vertical wheel acceleration were published in (B\u00f6rner et al. 2002, Weispfenning and Isermann 1995). Now a comparison of different indirect tire pressure monitoring methods was performed. Therefore, a test vehicle equipped with sensors for both signals was used to record data. The wheel speeds of the different wheels will be compared to detect changes in the tires\u2019 radii" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003307_j.proeng.2011.11.126-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003307_j.proeng.2011.11.126-Figure1-1.png", "caption": "Fig. 1. Oscillating face grinding (OFG) process", "texts": [ " vos \u2013\u2013 = vf 2+ vh \u2013\u20132 with vh \u2013\u2013 = As\u2219fs (1) The superposition of the grinding wheel circumferential speed vs and the average oscillation speed vos \u2013\u2013 leads to the resulting cutting speed vc for the oscillating face grinding process, which can be determined by using the feed rate vf of the workpiece, the average oscillation speed vos \u2013\u2013 of the grinding tool and the idealized structure angle \u03b1s (Equation 2). The idealized structure angle \u03b1s is geometrically described by using the feed rate vf of the workpiece as well as the average stroke speed vh \u2013\u2013. At the bottom and top dead centre the oscillation speed of the grinding tool has a minimum, resulting in a lower cutting speed at these turning points. The kinematic interaction for the oscillating face grinding process is shown in Figure 1, which is characterized by a superposition of three different speed vectors. vc = vs 2+vos \u2013\u20132+2\u2219 vs\u2219 vos \u2013\u2013 cos(\u03b1s) with Referring to the generated surface structures of the workpiece, the knowledge about the engagement of geometrically undefined cutting edges with the workpiece is important. Thereby the active cutting edge is on a defined path and is cutting due to the relative velocity between workpiece and tool. Therefore the surface structure of the workpiece will be generated by the path of the abrasive grain, which is unidirectional for conventional processes (Figure 2 A)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001675_00368790810902223-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001675_00368790810902223-Figure1-1.png", "caption": "Figure 1 Schematic view of the FZG gear test rig", "texts": [ " The mineral oil did not match the minimum requirements of 60 per cent biodegradability in 28 days, as shown in Table I. Thus, no toxicity tests were performed for this lubricant. The esterbased oil exceeded the minimum requirements of 60 per cent biodegradability in 28 days and passed both toxicity tests, OECD 201 \u201cAlga growth inhibition test\u201d and OECD 202 \u201cDaphnia Magna acute immobilization\u201d, as shown in Table I (Martins et al., 2005, 2006). All the gear tests were performed on the FZG back-to-back spur gear test rig, shown in Figure 1, and the gears used are similar to standard FZG typeC gears, having the geometric characteristics presented in Table II (Winter and Michaelis, 1985). The gears were manufactured in DIN 20 MnCr 5 steel, carburized to a depth of 0.8mm and a surface hardness 58\u201362 HRC. It has a very high surface hardness while keeping a very high toughness on the core, being a typical material used for manufacturing highly stressed parts such as gears, crankshafts, . . . The physical and mechanical properties (average values) of the gear material are presented in Table III" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003132_jmes_jour_1969_011_008_02-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003132_jmes_jour_1969_011_008_02-Figure6-1.png", "caption": "Fig. 6", "texts": [ " 1 0 ~ ~ ' 1 . 8 6 ~ ~ ' u ' I u 1 - [v. .] = - 0 . 0 6 9 ~ 2 ~ 0.402~2' 0.743~2' (30) T 0 . 0 2 3 ~ ~ ~ 0 . 1 4 7 ~ ~ ' 0 . 3 3 3 ~ ~ ' Hence the instability bands shown in Fig. 5 are calculated from equation (1 1). Although the bearings have radial symmetry, unstable bands are in evidence not only at the shaft natural frequencies, but also at half sums of these. The form of the half-sum instabilities is indicated in Fig. 5. Single-mass shaft on massive and jlexible foundations The idealized system, Fig. 6, is considered; the shaft principal stiffnesses on rigid bearings are A( 1 +a) and With /3 = 10, k, = 100 and k, = 500 the natural A( 1 - u). frequencies of the generating system are found to be wl,, = 0.9552, wcl = 0.9952, wy2 = 3.252 wZ2 = 7.0952 where The corresponding modal shapes, shown in Fig. 7, yield the inertia coefficients 1 (32) ayl = 1-098m, a;, = 1.005m, ayz = 846m a,, = 24OOOm The coefficients p, v are now determined from equations (5), including the 'Stieltje' terms due to shear at the shaft bearings, 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 K C E 5 Vol I1 No 1 I969 at DEAKIN UNIV LIBRARY on August 12, 2015jms" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002349_s12239-009-0039-8-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002349_s12239-009-0039-8-Figure2-1.png", "caption": "Figure 2. Assembly of the ISCVT.", "texts": [ " The design parameters were determined, and the transmission performances were evaluated by the optimal design procedure. The transmission efficiency, the life span, the work needed for changing the speed ratio, the internal forces, and the maximum severe stresses on each part were theoretically investigated, and the efficiency performances were experimentally measured. The manufactured prototype was installed in a motorcycle, which was fixed on a test-bench equipped with a dynamometer system. The parasitic loss and the cross-sectional road load performance were tested. The ISCVT, as shown in Figure 2, has two rolling contact points between three rotors: the driving rotor, the counter rotor, and the driven rotor, as do most speed ratio changeable traction drive CVTs. The engine power flows successively through the driving shaft, the driving rotor, the counter rotor, the driven rotor, and the driven shaft. The driving shaft is connected to the engine output, and the driven shaft is connected to the rear wheel. The rolling bodies, the convex counter rotor and the concave driving and driven rotors, have exactly spherical surfaces with the functionally unnecessary parts eliminated, as shown in Figure 4" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003167_iecon.2009.5415436-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003167_iecon.2009.5415436-Figure3-1.png", "caption": "Fig. 3. Top view of the mobile robot Auriga-\u03b1 (left) with a utility trailer (center) and a spraying trailer (right).", "texts": [ " Algorithm 1: Virtual Tractor Steering Limitation for i = 1 to n do (Steady analysis) Compute equilibrium limit \u03b3i m1 (12) Compute mechanical limit \u03b3i m2 (13) Compute forward-propagation limit \u03b3i m3 (14) \u03b3i m = min(\u03b3i m1 , \u03b3i m2 , \u03b3i m3) \u03b3v 0 m = \u03b3nm repeat (Transient analysis) \u03b3v 0 m \u2190 (\u03b3v 0 m \u2212 \u0394\u03b3) Evaluate step response from \u2213\u03b3v 0 m to \u00b1\u03b3v 0 m until equilibrium is reached without inter-unit collision Result: \u03b3v 0 m IV. IMPLEMENTATION The backward motion control with steering limitations presented in the previous section has been tested on the Auriga-\u03b1 mobile robot (see Fig. 3). Its dimensions are 1.24 m (l), 0.75 m (w) and 0.84 m (h), and it weights 258 kg. Auriga-\u03b1 is a tracked vehicle driven by two DC motors with gear-reduction and incremental shaft encoders for deadreckoning. The maximum speed of each track is 1 m/s. The vehicle top speed coincides with this limit in straight-line motion, but it decreases to zero according to the increase in the demanded curvature. The track speed controller runs in an on-board DSP every 10 ms, which also provides odometric data every 30 ms", " This also allows to define the virtual rear axle of this vehicle, which is relevant to determine hitch parameter L0b (see Fig. 1). A Sick LMS 200 time-of-flight laser scanner is mounted on the forward part of the vehicle at a distance of 0.5 m ahead of its coordinate center. To correct odometric estimations of the actual tractor, an accurate laser scan matching technique has been employed every 270 ms [14]. The dimensions of the two wheeled trailers are similar to the tractor. The first trailer is employed for carrying loads, while the second one is for spraying (see Fig. 3). Each angle \u03b8i is indirectly obtained with a draw-wire displacement sensor. Kinematic parameters and mechanical angle limits for this particular setup are the following: L0b = 0.71 m, L1f = 0.99 m, L1b = 0.61 m, L2f = 0.81 m, \u03b81 m = 68\u25e6 and \u03b82 m = 43.6\u25e6. To estimate the virtual tractor pose with respect to the global frame (step 1 of section II-B), \u03b8i measurements and geometric relations between units are employed: xi = xi\u22121 + Li\u22121 b sin(\u03c6i\u22121) + Li f sin(\u03c6i), (16) yi = yi\u22121 \u2212 Li\u22121 b cos(\u03c6i\u22121) \u2212 Li f cos(\u03c6i), (17) including (1) to propagate heading" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001370_s11044-007-9038-6-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001370_s11044-007-9038-6-Figure2-1.png", "caption": "Fig. 2 McPherson Strut suspension model", "texts": [ " (23) Using Equations (21)\u2013(23), the effective inertia matrix and effective force vector required in Equation (2) can also be obtained as the following equations: M c = Myy \u2212 Q\u2217T y (M\u2217)\u22121Q\u2217 y (24) P c = Py \u2212 Myqu R \u2212 Q\u2217T y (M\u2217)\u22121Q\u2217 q . (25) 3.1 Generation of approximate function using polynomials When developing real-time multibody vehicle dynamics models, it is necessary to make an accompanying vehicle model, using a general purpose program such as ADAMS [13], for verification purposes. Thus, to obtain the approximate function, ADAMS can be utilized to generate data, which represent the relationship between the dependent and the independent coordinates. Figure 2 shows the ADAMS model of the McPherson strut suspension. It consists of a LCA (lower control arm), an upper strut, a tie-rod, a knuckle, and a chassis frame. The LCA is connected to the chassis frame with a revolute joint. Spherical joints are used between the LCA and the knuckle and between the upper strut and the chassis frame. The upper strut is connected to the knuckle with a translational joint. The tie-rod connects the knuckle and the rack. Spring and damper elements are used. Bump and rebound stoppers are also modeled as nonlinear springs" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002361_gt2007-27314-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002361_gt2007-27314-Figure5-1.png", "caption": "Figure 5. Brush gear unit: (a) View on removable brush carriers; (b) Brush carrier with three brushes.", "texts": [ " The journal bearings and the hydrogen seal ring were excluded by variation of the lube oil temperatures and lube oil flows, which did not change the vibrational behavior. A test without carbon brushes revealed that the carbon brushes sliding on the slip rings were the location of the hot spot. The generator was brought to rated speed without brushes and operated for about one hour. No spiral vibration occurred. Immediately after insertion of the brushes, which can be done under operational conditions (Fig. 5), spiral vibration appeared with increasing magnitude. After brush removal the spiral vibration promptly stopped and the vibration vector returned straight to its original position, when the test was started. Fig. 4 shows the polar plot of the relative shaft vibration measured at the NDE-bearing during this run. The brush gear unit supplies the DC-current to the field winding of the generator rotor. The field current of hydrogencooled generators is in the range from 3000 to 7000 A. The slip ring shaft (SR) with two slip rings and a radial fan for cooling is supported by the NDE- and the end bearing. Carbon brushes sliding on the slip rings transfer the current from the stationary brush holders (Fig. 5) to the rotating slip rings. The slip rings are shrunk on the shaft via soft insulation material for electrical nloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 10/18/2017 Te insulation. Nevertheless the thermal bow of the slip rings is transmitted to the shaft to some extent. Slip ring shafts may be particularly prone to the spiral vibration phenomenon due to the friction heat that is inevitably generated by the brushes. During operation the generated friction heat may increase due to too low humidity (below 4", " The stability threshold is determined as a function of the rotor speed and the ratio of the added heat to the dissipated heat at the slip rings. A threshold at low ratio indicates that the shaft is sensitive to hot spots, i.e. little heat input is necessary to destabilize the system. The calculation method for generalized shaft systems according to Schmied [1] is described in ANNEX A. Three different heat input models In case of slip rings, the friction between the carbon brushes and the slip ring causes the hot spot. The brushes are pressed against the slip ring by roller springs (Fig. 5b). The pressure around the circumference may vary due to the shaft vibration. In case of a synchronous vibration it is always the same point on the shaft, which is submitted to an increased heat input. The pressure can vary due to the following three mechanisms (Fig. 6, case a to c). The brush contact force varies: a) Due to the stiffness of the roller springs, b) Due to the friction in the brush holder, which inhibits the movement of the brush in radial direction, c) Due to the inertia effects of the brushes", " The most suitable case will be selected by comparison of the analytical results to the practical experience. Case a: Heat input proportional to shaft displacement. For a rub against with a spring-contact force this would be the appropriate relation. It is a good approximation for the effect in a journal bearing or labyrinth seals. Kirk\u2019s method [8] to estimate the stability is limited to the effect in a journal bearing. A stability criterion is given although it does not provide a clear stability threshold. However, roller springs (Fig. 5b) are designed to provide a constant force not dependent on spring displacement or brush wear respectively and thus have very low spring stiffness. Therefore, it is highly likely, that one of the other two mechanisms prevails. Case b: Heat input proportional to shaft velocity. For a rub with a contact force varying due to Coulomb friction this would be the appropriate relation. This may be a good approximation for the brushes sliding within the brush holders. It is assumed that the heat input is proportional to the shaft velocity instead of the shaft displacement, since a friction force changes its direction with the direction of the velocity" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003596_icelmach.2010.5608115-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003596_icelmach.2010.5608115-Figure7-1.png", "caption": "Fig 7: Spatial distribution of the F created by the winding of each phase", "texts": [ " So from (9) and (10) a correction factor XBn for each spatial component can be defined as: cnt cntn m g cnt cntn m g n m m XB = = \u2260 \u2260 \u03a6 \u03a6 = \u03b2 \u03b1 \u03b2 \u03b1 (12) Then the correction factor can be applied to the spectrum of the flux linkage obtained from FEM2D simulations in order to calculate the results in 3D: ndndn XB\u22c5= 2_3_ \u03c8\u03c8 (13) After that, the temporal flux linkage is calculated applying the Inverse Fast Fourier Transformation (IFFT): ( ) ( )o n dnd tnpt \u03bb\u03c8\u03c8 \u2212\u03a9= \u2211 \u221e \u2212\u221e= sin3_3 (14) B. Analysis of the Magneto Motive Force per Current Unit The second effect is the one produced by F. In Fig 7 is shown the spatial distribution of this parameter. The expression for each harmonic of F is the following: k p ph k tk Nj F \u03be \u03c0 \u22c5\u22c5 \u22c5 = (15) Nph is the number of turns per phase. \u03be is the winding factor, which is the combination of the span factor \u03bes_k and pitch factor \u03bep_k: kpksk __ \u03be\u03be\u03be \u22c5= (16) The span factor is defined as: ( ) \u239f \u23a0 \u239e \u239c \u239d \u239b \u22c5\u22c5= 2 sin_ rtk w pks \u03b2\u03be (17) And the pitch factor is defined as: ( ) ( ) 2 2 sin _ r tk r tk w p w p kp \u03b1 \u03b1 \u03be \u22c5\u22c5 \u239f \u23a0 \u239e \u239c \u239d \u239b \u22c5\u22c5 = (18) \u03b2w and \u03b1w are the winding span angle and winding pitch angle respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000718_1.1829068-Figure15-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000718_1.1829068-Figure15-1.png", "caption": "Fig. 15 Modified Humpage trains used for modeling of wobble gear: \u201ea\u2026 train with tooth-number constraint satisfied \u201eS1 \u2026 and \u201eb\u2026 equivalent train with all geometric constraints satisfied \u201eS0 \u2026", "texts": [ " From Table 2 representing this constrained case of the Humpage set with links 3 and 4 being the input and output, respectively, the reduction ratio simplifies to v4 /v3512N2 /N4 (5) Recall that the wobble-gear speed relationship has been given in Eq. ~4!. This is identical to the analogous speed ratio given by the special case of the Humpage set in Eq. ~5!. In summary, for the physical construction of the train satisfying the tooth-number constraint mentioned earlier, the similarity index improves to S1, as shown in Fig. 15~a!. Note that link 5 has been relabeled in Fig. 15 as link 1 since link 1 was removed from the original Humpage 274 \u00d5 Vol. 127, MARCH 2005 rom: http://mechanicaldesign.asmedigitalcollection.asme.org/pdfaccess.a train. For the embodiment which satisfies both of the geometric constraints ~velocity vectors for link 2 align for both mechanisms!, the similarity index is improved to S0 ~indicating kinematic equivalence! as shown in Fig. 15~b!. It is interesting to note that three different kinds of mechanisms have been discussed in this section: those which nutate but are not pure gear trains ~such as the wobble gear!, those which deform ~the harmonic drive!, and those which are pure gear trains and whose wobbling motion is caused by the revolute axes changing position during motion ~epicyclic gear trains!. Despite their fundamental differences, the motion of these devices characterizes them as a class of kinematically similar or, in some cases, kinematically equivalent, mechanisms", "url=/data/journals/jmdedb/27802/ on 03/07/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F lent geared mechanism instead of the wobble gear itself, the process is greatly simplified, since automated techniques exist for kinematic analysis of gear trains. It is noteworthy that a simple, automated kinematic analysis technique for general bevel-gear trains has recently been proposed by Nelson and Cipra @13#. The equivalent gear train to be used is similar in appearance to the Humpage train and is shown in Fig. 15~b!. Note that the wobble gear is equivalent to both the Humpage train ~with geometric constraints satisfied! and the mechanism of Fig. 15~b!, but the equivalence of Fig. 15~b! is preferred since the number of links is the same as the wobble gear. Links 1 and 3 of the equivalent mechanism are considered the inputs ~known velocity, with link 1 fixed!. The tooth numbers are N1560, N2550, N28560, and N4549. Here N1 and N28 are arbitrarily chosen, with the only constraint being that N15N28 , as discussed in the previous section. From the geometry of the wobble gear, it can be seen that the bevel angle b25cos21~49/50!511.48 deg (6) Using all of this information, the automated method of Nelson and Cipra gives the kinematic solution HvW 2 vW 4 J 5F cos~11.48! 12cos~11.48! 50 49 21 49 G Hv1 v3 J zW 1F sin~11.48! 2sin~11.48! 0 0 G Hv1 v3 J rW (7) where the z and r vectors represent the principal directions shown in Fig. 15~a!. With link 1 fixed and link 3 given a unit input velocity HvW 2 vW 4 J 5H 12cos~11.48! 21 49 J zW1 H 2sin~11.48! 0 J rW 5 H 0.02 20.0204J zW1 H 20.199 0 J rW (8) Notice that the gear ratio appears in the expression for the output speed v4 of the wobble gear. More generally, since v2 is coupled to the numbers of teeth, with link 1 fixed HvW 2 vW 4 J 5 H 12cos b 12sec b J zW1 H 2sin b 0 J rW (9) Note that in this form, the reduction ratio is (12sec b)21 5cos b(cos b21)21. Also, for small values of b, the wobbling member has relatively small velocity components v2,r and v2,z " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002018_icarcv.2008.4795694-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002018_icarcv.2008.4795694-Figure10-1.png", "caption": "Fig. 10. Robot relative formation", "texts": [ " x = r \u2217 sin(\u03b8) y = r \u2217 cos(\u03b8) (2) The relative velocities in each Cartesian coordinate are then approximated as (\u0394x/t,\u0394y/t) using the relative positions history and the timestamp readings. Once the relative fire-fighter position and velocity are obtained, the robot calculates the estimated global position and velocity of the fire-fighter using its own estimated position and velocity. Then the desired relative position of the robot with respect to the fire-fighter is calculated according to the formation rules. For the demo the formation will be a simple isosceles triangle of approximately 1.5 meters high and 40 centimeters wide as seen in figure 10. This position is then transformed to a global position and passed to the local navigation module (ND) which will move the robot to this temporary goal while doing obstacle avoidance with laser and sonar data. The fact that the fire-fighter is in front of the robot does not conform a problem for the obstacle avoidance routine as the firefighter is far from the obstacle avoidance distance threshold. First, the experiment presented has been partially tested in simulation. The figure 11 shows the scheme of what we have perform" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001827_jjap.47.1203-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001827_jjap.47.1203-Figure2-1.png", "caption": "Fig. 2. Experimental setup for photoinduced birefringence and chirality measurement.", "texts": [ " The third material is poly{2-[4-(4-cyanophenylazo)phenoxy)]ethyl methacrylate-co-2-[2-methyl-4-(4-phenylazophenylazo)phenoxy]ethyl methacrylate} (PCDY), which is a copolymer consisting of two dyes: cyanoazobenzene with MMA and disperse yellow 7 (DY7) with MMA, as shown in Fig. 1(d). The molar ratio of the two dyes is 50 : 50 (PCDY50). The film thickness is 800 nm. The glass temperatures of these materials are about 120 C because the polymer main chains of these materials are the same. All experiments were performed at room temperature, and the thermal stability was not evaluated. The orientational stability of the azobenzene monomer or side chain was evaluated. The thermal stability can be controlled by changing the polymer main chain. Figure 2 shows the experimental setup in this study. Optical anisotropy was induced in the azobenzene materials by illumination with an Ar\u00fe laser beam (488 nm). The diameter of the beam was set to 3mm. The optical power was adjusted from 10 to 100mW using a neutral density (ND) filter. The photoinduced anisotropy was probed by measuring the polarization state of a diode laser beam (670 nm) through the azobenzene materials. The polarization state was calculated from Stokes parameters obtained with a polarimeter" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002821_j.sna.2010.07.002-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002821_j.sna.2010.07.002-Figure5-1.png", "caption": "Fig. 5. Deformation of", "texts": [ " In this frame which moves with the wave crest, the normal displacement of each point on the neutral plane of stator Eq. (1) becomes time independent as follows [13]: Uz(x) = A cos ( 2 x ) (2) Fig. 4 shows the deformation of stator and frictional layer of roller. In RIUSM the rollers are coated with the frictional layer. The frictional layer contacts with both stator and rotor. The properties of this frictional layer are given in Table 1. d frictional layer of roller. frictional layer. w p e w U c here \u201ca\u201d is the distance between surface of stator and neutral lane of stator (Fig. 5). The frictional layer is described by a viscolastic foundation model (Fig. 4). Normal and tangential stiffness hich result from shear deformation Uk x and normal deformation k z of the frictional layer is obtained as follows: N = Pb Uk z = Eb h , cT = b Uk x = Eb h2(1 + ) (4) e spri w t w l m w a t i i f m i T n here E, , , P, h and b are modulus of elasticity, Poisson ratio, angential stress, normal stress, thickness of the frictional layer and idth of frictional layer, respectively. The equivalent mass per unite ength of frictional layer is obtained as follows: = 0", " It is assumed that normal nd tangential damping (dN, dT) of frictional layer are equal and he values of them are estimated for frictional material. The rotor s assumed to be rigid and its speed (VR) is prescribed. The roller s pressed against stator by preload F, and stator penetrates into rictional layer. The stick\u2013slip phenomenon is considered in contact odel. The tangential contact stress (x) and the normal stress p(x) n both, side and down portion of roller will be calculated later. he computations are related to single roller. The origin lies in the eutral plane of the undeformed stator. According to Fig. 5 normal ng for lateral preload is not illustrated. deformation of frictional layer in down portion of roller relative to stator is obtained as follows: Uk z (x) = Uz(x) + a + \u221a r2 \u2212 x2 \u2212 H, \u2212 L \u2264 x \u2264 L Uk z = A cos ( 2 x ) + a + \u221a r2 \u2212 x2 \u2212 [ A cos ( 2 L ) + r cos + a ] , \u2212 L \u2264 x \u2264 L L = r sin (6) where \u201cH\u201d is vertical distance of the center of roller to x-axis (natural axes of undeformed stator). Velocity and acceleration of the frictional layer are obtained as follows: U\u0307k z = A 2 vw sin ( 2 x ) (7) F R U According to Fig", " In the stick one, the frictional layer velocity vr(x) is equal to the stator velocity s(x) using Eq. (19) Uk\u2032 x (x) can be formulated as bellow: k\u2032 x (x) = vR \u2212 vs(x) vw (20) By integration and differentiation, Uk x (x) and Uk\u2032\u2032 x (x) can be calulated: k x (x) = vR x + 2 Aavw sin(2 x/ ) vw + D2 (21) k\u2032\u2032 x (x) = \u22128 3aA sin (2 x/ ) 3 (22) Inserting Eqs. (20)\u2013(22) in Eq. (15) the shearing strength stick(x) n stick zone can be derived as follows: stick(x) = \u22128mvw 2 3Aa sin(2 x/ ) 3b \u2212 dT [vR 2 + 4 2Aavw cos (2 x/ )] 2b + E 2(1 + ) ( vR x + 2 Aavw sin(2 x/ ) bvw + D2 ) (23) According to Fig. 5 displacement, velocity and acceleration f normal element of frictional layer in side portion of roller is btained as follows: k x (z) = \u221a r2 \u2212 (z \u2212 H)2 \u2212 Ux(L\u2032), L\u2032 = r cos( \u2032) (24) tuators A 163 (2010) 304\u2013310 309 Uk x (z) = \u221a r2 \u2212 (z \u2212 H)2 \u2212 aA 2 sin ( 2 L\u2032 ) (25) U\u0307k x (z) = aA ( 2 )2 vw cos ( 2 L\u2032 ) (26) U\u0308k x (z) = aA ( 2 )3 vw 2 sin ( 2 L\u2032 ) (27) According to Fig. 8 Newton\u2018s second low can be applied for normal element of frictional layer in side portion of roller. mdzU\u0308k x (z) = p(z)bdz \u2212 cNdzUk x (z) \u2212 dNdzU\u0307k x (z) (28) Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.124-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.124-1.png", "caption": "Fig. 14.124 Scheme of pneumatic feeder for dense mortars", "texts": [ " The machines most frequently used for floor work are: \u2022 Pneumatic feeders for dense mortars (feeder of fresh concrete mix and mortar)\u2022 Vibrating beams (described in Sect. 14.3)\u2022 Floating machines for concrete (described in Sect. 14.3.8)\u2022 Grinders for stone and mineral floors\u2022 Sanding\u2013polishing machines for wooden floors (parquets) A pneumatic feeder for dense mortars is used for mixing the components of cement mortar and delivering the latter to the placement site. A scheme of such a feeder is shown in Fig. 14.124. The vessel filled with the appropriate amount of water can be charged with mortar dry components manually or by means of a charging bucket. Compressed air supplied to the vessel and dosing unit forces mortar out of the vessel, but in the dosing unit the stream of mortar is separated by compressed air. As a result, in the pressure hose it already has the form of a series of small portions separated by spaces filled with compressed air. This transport system ensures stable flow of mortar into the hose\u2019s tip placed on Part B 1 4 " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003887_19346182.2012.663534-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003887_19346182.2012.663534-Figure5-1.png", "caption": "Figure 5. (a) If a racquet approaches a ball at speed V with the head tilted forward, then the ball will be served with topspin even if the head is not rising. The same amount of spin is generated if the ball approaches the racquet at speed V and the racquet is initially at rest, as shown in (b).", "texts": [ " That angle could well be larger than the actual vertical approach angle of the racquet in a typical kick serve, in which case the ball toss would account for more than half of the topspin generated. From Equation (1), we find that v \u00bc 93 rad s21 (889 rpm) due to the ball toss alone. If the racquet head is tilted forward when the head strikes the ball, rather than being exactly vertical, then additional topspin is generated and the ball will pass lower over the net. The same effect occurs in a topspin groundstroke. The effect of racquet head tilt is illustrated in Figure 5. Suppose that the head is tilted forward by an angle A and approaches the ball in a horizontal direction at speed V, as in Figure 5(a). The physics of the collision is exactly the same if the racquet is at rest and the ball approaches at speed V as in Figure 5(b). Since the ball approaches the racquet at angle A, it will bounce off the racquet at an angle with the same topspin as that given by Equation (1). The result of the collision in Figure 5(a) is that the ball is served in a downward direction with topspin, even if the racquet head is not rising when it strikes the ball. The strings grip the ball during a serve. If the racquet rotates 108 while the ball is on the strings, then the ball also rotates by about 108, in the topspin direction. The same effect would occur if the ball was glued to the strings since the ball and the racquet would both rotate 108. The ball is not glued to the strings but it is squashed against the strings. The top end of the racquet is rotating faster than the bottom end, so the top side of the ball is pushed harder towards the net than the bottom side", " Despite the fact that players swing up to the ball in a kick serve, the ball comes off the strings in a downward direction, typically about 58 below the horizontal. If it didn\u2019t, the ball would land near the baseline. This result indicates that the upward motion of the racquet head during the serve is only one of several different effects contributing to the final result. Downward motion of the ball can be attributed to the initial D ow nl oa de d by [ U ni ve rs ity o f D el aw ar e] a t 1 7: 40 2 7 Ju ly 2 01 3 downward velocity resulting from the ball toss, or to the effect of racquet tilt shown in Figure 5, or the effect of rotation of the racquet during the impact, or a combination of all three effects. The analysis given in the previous section was simplified by ignoring the fact that the normal reaction force on the ball rotates in direction as the racquet rotates. Such an approximation can be justified on the basis that a small direction change in the force on the ball will have a negligible effect on the resulting torque on the ball. A more significant result is that the change in direction of the normal force will affect the outgoing angle of the ball" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.115-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.115-1.png", "caption": "Fig. 14.115 Diagram of two-piston mortar pump with mechanical cam drive for the pistons", "texts": [ " As the piston moves again to the right, mortar is sucked in again while the mortar on the piston\u2019s right side is forced into a pipeline. The pump makes it possible to minimize pressure fluctuations and therefore to maintain constant mortar spraying parameters and increase the fatigue strength of the mortar pipeline\u2019s flexible hoses. The travel of the pistons occurs as a result of the rotation of the cam pushing the roller during, respectively, the delivery stroke and the suction stroke. The operation of a popular two-piston mortar pump with a cam drive is illustrated in Fig. 14.115. The function of the compensating piston is to equalize the mortar forcing pressure during the suction stroke of working piston. The suction stroke of the working piston is aided by a spring and the return stroke of the compensating piston results from the mortar pressure. Besides piston pumps, screw pumps form another class of pumps. They are used for applying thin finishing coats of gypsum mortars, lime\u2013gypsum mortars, and mortars with plastics. In recent years they have also been used for conventional mortars, competing with piston pumps" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000937_iros.2003.1248806-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000937_iros.2003.1248806-Figure5-1.png", "caption": "Fig. 5 . Computing A(s/Lss)", "texts": [ " The area swept by the bar along a segment is called A(s1Lss). The computation of A(s1Lss) is done by drawing a circle centered& each one of the 2 end points of the segment with radius Lss (curve type 2) and 2 parallel lines a t Lss distance from the segment (curve type 1). The lines are drawn on both sides of the segment and are called plr and p l l . The regions inside curve type 2 are called ct and cb A(s]Lss) is equal to the union of the polygon defined by the end points of plr and pll with regions d and cb minus the intersection of ct and cb (see figure 5 ) . A(b1Lss) = UA(S(LSS)vjsEb. If cgr - A(b1Lss) = 0 then A(cgr1Lss) = A(b1Lss) (figure 6 b), else there is a hole inside A(b1Lss) and there are two cases. When all the curves type 2 related to the vertices of the cgr intersect, the boundary of the hole is composed by arcs of circle. This hole region is the intersection of all the curves type 2. In this case A(cgr1Lss) = A(b1Lss) as well (figure 6 c). However, if this hole is composed by straight line segments then A(cgr1Lss) = A(b1Lss) u c g r thus this hole will disappear (figure 6 a)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001970_robot.2008.4543773-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001970_robot.2008.4543773-Figure1-1.png", "caption": "Fig. 1. 4DOF upper-limb power-assist exoskeleton.", "texts": [ " If such an object is identified, then calculate the position of the new object by the image processing algorithm and trajectory modification is carried out by exoskeleton to guide the arm toward the new object. The proposed power-assist method is applied to a 4DOF upper-limb power-assist exoskeleton and effectiveness of the proposed intelligent exoskeleton has been evaluated by D 978-1-4244-1647-9/08/$25.00 \u00a92008 IEEE. 3666 experiment. II. 4DOF UPPER-LIMB POWER-ASSIST EXOSKELETON In order to assist 4DOF upper-limb motion, a power-assist exoskeleton (Fig. 1) which consists of a shoulder motion support part, an elbow motion support part, and a forearm motion support part was developed [12]. The shoulder motion support part, elbow motion assist part, and forearm motion support part are the main parts of the exoskeleton. The exoskeleton upper-limb is supposed to install in a wheel chair [13]. The stereo camera system is also supposed to install on the wheel chair as shown in figure 1. The stereo vision digital camera and sonar sensor is used for perception-assist. The camera is placed in a position high enough over the head of the user, hence visual images in front of the user\u2019s arm can be obtained. Therefore, it is possible to detect the positions of the objects in front of the user which are possible to grab or touch by the user. The sonar sensor is placed in the wrist arm holder, hence it moves with the arm of the user and can detect the object toward which the user is moving his/her arm" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001571_s1560354707030045-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001571_s1560354707030045-Figure4-1.png", "caption": "Fig. 4. Mathematical model for the Snakeboard.", "texts": [ " To propel the snakeboard the rider first turns both of his feet *E-mail: kuleshov@mech.math.msu.su 321 in (see Fig. 3). By moving his torso through an angle, the snakeboard moves through an arc defined by the wheel angles. The rider then turns both feet so that they point out, and moves his torso in the opposite direction. By continuing this process the snakeboard may be propelled in the forward direction without the rider having to touch the ground. The mathematical model of the snakeboard considered in this paper is represented in Fig. 4. We assume that the snakeboard moves on the xy plane, and let Oxy be a fixed coordinate system with origin at any point of this plane. Let x and y be the coordinates of the system center of mass (point G) and \u03b8 the angle between the central line of the snakeboard and the Ox-axis. In the basic model treated in [1] platforms could rotate through the same angle in opposite directions with respect to a central line of the snakeboard (by other words, for this model \u03d5f = \u2212\u03d5b = \u03d5, see Figs. 3\u20134). We suppose that platforms can rotate independently and their positions are defined by two independent variables \u03d5f and \u03d5b (Fig. 4). The motion of the rider is modeled by a rotor, represented in the form of a dumb-bell in Fig. 4. Its angle of rotation with respect to the crossbar is denoted by \u03b4. REGULAR AND CHAOTIC DYNAMICS Vol. 12 No. 3 2007 Let l be the distance from the system center of mass G to the location of the wheels (points A and B). We assume that |GA| = |GB| = l, see Fig. 4. The platforms of the snakeboard are assumed to move without lateral sliding. This condition is modeled by constraints which may be shown to be nonholonomic. For the front platform corresponding constraint has a form x\u0307 sin (\u03d5f + \u03b8) \u2212 y\u0307 cos (\u03d5f + \u03b8) \u2212 l\u03b8\u0307 cos \u03d5f = 0 (1) and for the rear platform it has a form x\u0307 sin (\u03d5b + \u03b8) \u2212 y\u0307 cos (\u03d5b + \u03b8) + l\u03b8\u0307 cos \u03d5b = 0. (2) We can solve equations (1) and (2) with respect to x\u0307 and y\u0307. Then x\u0307 = l\u03b8\u0307 sin (\u03d5f \u2212 \u03d5b) (cos \u03d5b cos (\u03d5f + \u03b8) + cos \u03d5f cos (\u03d5b + \u03b8)), y\u0307 = l\u03b8\u0307 sin (\u03d5f \u2212 \u03d5b) (cos \u03d5b sin (\u03d5f + \u03b8) + cos \u03d5f sin (\u03d5b + \u03b8))" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001889_ac60340a007-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001889_ac60340a007-Figure1-1.png", "caption": "Figure 1. A : Conventional impedance bridge. 8. Operational amplifier equivalent of ( A )", "texts": [ " One serious limitation of the bridge method is that obtaining a true null is a time-consuming and tedious task, making measurements a t IOU frequencies and measurements of rapid- 1 To u horn correspondence s h o u l d b e addressed ( 1 ) B Hague Alternating Current Bridge Methods Pitman London 1946 ly changing systems impractical or impossible. The problem of bridge balancing may be completely removed while retaining much of the high accuracy and precision of the bridge by using current feedback as shown in Figure 1. This circuit is an analog of the conventional impedance bridge where two arms have been replaced by active elements: one: a signal generator, and the other, the output of an operational amplifier (Keithley 301K). The cell and R, comprise the other two arms as in the conventional impedance bridge. The differential input of the operational amplifier (OA) serves as a null detector and keeps the bridge in continuous balance even as cell impedance changes with frequency or chemical composition. Since the bridge is continuously in balance and the OA output is the direct analog of the series RC arm in the conventional impedance bridge, one need only resolve the output of the OA into in-phase and quadrature components (with respect to the signal generator) to ohtain the current voltage phase relationship in the cell" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003503_pime_conf_1969_184_137_02-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003503_pime_conf_1969_184_137_02-Figure1-1.png", "caption": "Fig. 1 . I . Disc machine", "texts": [ " of this paper was received at the Institution on 6th May * ShelZ Research Lid, Thornton Research Centre, P.O. Box 1, * References are given in Appendix 1.2. 1969 and accepted for publication on 17th July 1969. 22 Chester, CHI 3SH. made for the investigation, for reasons explained later, and were somewhat simpler in their formulation than normal commercial greases, there is no reason to suppose that their mechanical properties differed essentially from those of greases formulated along more conventional lines. The disc machine shown in Fig. 1.1 has been described elsewhere (I), and is essentially similar to that used by Crook (4). The discs were of case-hardened En 34 steel, 76.2 mm (3 in) in diameter and 25-4 mm (1 in) in width. To accommodate small relative displacements in the axial direction, the edges of one disc were rounded to give an effective contact width of 22.2 mm (0.875 in). The discs were ground to an eccentricity of less than 2.5 pm (0.0001 in) and to a surface roughness of 0.0375-0*05 pm (1.5-2 pin) c.1.a. The finish was improved to better than 0", " During the course of the work it became necessary to admit the possibility that the inlet film would not always be completely full. The extreme case of an empty inlet film would be when its configuration was the same as that previously assumed for the outlet film, and this is referred to as the \u2018empty inlet\u2019 condition. In practice the film thickness, h,, would be expected to lie between the two extreme values calculated for the \u2018full inlet\u2019 and \u2018empty inlet\u2019 conditions. An example of the effect on the estimated value of film thickness of the assumption made regarding the filling of the inlet section is shown in Fig. 1.2. Vol184 Pt 3F Proc lnstn Mech Engrs 1969-70 at TUFTS UNIV on December 8, 2015pcp.sagepub.comDownloaded from FILM THICKNESSES IN ELASTOHYDRODYNAMIC LUBRICATION OF ROLLERS BY GREASES 3 2 7 2 6 c z u) z 3 2 5 8 2 4 0 cr W L 2 3 separated from a lower electrode of much greater area by three Tufnol pins of 6-35 mm (0.25 in) diameter. These pins, arranged at the points of an isosceles triangle, protruded 1.02 mm (0.040 in) from the face of the lower electrode. Before use the condenser was calibrated with fluids of known dielectric constant. No attempt was made to measure the dielectric constant of the greases at high pressures owing to experimental difficulties and to an uncertainty in the application of the results that will now be discussed. GNrn-' 0 10 20 30 40 50 60 70 80 90 100 PRESSURE, lb i f2x103 Fig. 1.3. Effect of pressure on the dielectric constant of test oils at 60\u00b0C and 19 kHz The interpretation of the capacitance measurements in terms of film thickness depends not only on the degree of filling by the lubricant as discussed above, but also on the dielectric constant of the lubricant at the temperature and pressures existing in the disc machine. The dielectric constants of the component oils at atmospheric pressure were measured in a modified Henley dielectric cell, while measurements at pressures up to 345 MN/m2 (50 000 lbf/in2) were made with a special cell, as described in reference (5). For the interpretation of capacitance measurements made at the higher of the two loads used, the dielectric constant was required at the mean Hertzian pressure of approximately 689 MN/m2 (100 000 lbf/in2), and this was estimated by graphical extrapolation from the results covering the range of pressure from atmospheric to 345 MN/m2 (50 000 Ibf/in2). The extrapolation is shown in Fig. 1.3. The extrapolation relating to oil 0, in Fig. 1.3 must be regarded as speculative, but the mechanism by which the dielectric constants of viscous polar liquids are thought to be affected by pressure (5) demands an increase in dielectric constant over the range of extrapolation. The assumed value of dielectric constant at 689 MN/m2 (100 000 lbf/in2) may be in error by a few per cent, and this may introduce a corresponding, though rather smaller, error into the conversion of capacitance measurements to values of film thickness for this oil. It will not, however, affect determination of the ratio of the thickness of a film formed by grease made from this oil to that formed by the oil alone, since the error in dielectric constant is common to both lubricants", " The plan was that a comparison between the behaviour of greases G, and G3 would show the effect of a change in the soap content with a constant base oil viscosity, while a comparison of GI, G,, and G4 would show the effect of base oil viscosity at a constant soap content. The calculation of the volume fraction of soap in the finished grease is explained in Appendix 1.1. The uncertainty in the dielectric constant must now AMIC LUBRICATION OF ROLLERS BY GREASES 5 be discussed in greater detail. The original fibrous structure of the soap in the grease, depicted in Fig. 1.4, will be broken down under the conditions of high shear stress encountered in the inlet zone of the disc machine, and the extreme end condition would be a dispersion of uniform spherical particles of soap in the oil medium. Fig. 1.5, which shows the condition of soap particles on one of the discs after many revolutions through the loaded contact, suggests that this end condition is probably Vo1184 Pt 3F Proc lnstn Mech Engrs 1969-70 at TUFTS UNIV on December 8, 2015pcp.sagepub.comDownloaded from 6 A. DYSON AND A. R. WILSON Lubricant attained. The dielectric constant, e3, of such a suspension would be (9) c3 = el+Ael Dielectric constant at 1 (50 000 lbf/in2) 1 ( l o \" o \" o \" ~ ~ ~ ~ ~ z ) ~ Atmospheric 345 MN/ma pressure* where 0, 0 4 GI Gz G3 and el is the dielectric constant of the oil medium, e2 is the dielectric constant of the soap, and c is the volume fraction of soap in the suspension", " Not all of the lubricants, however, were examined under all conditions. The film thicknesses deduced from capacitance measurements are shown in Figs 1.6-1.16. Results obtained in continuous tests at constant speed and load (Procedure 1) are shown in Figs 1.6-1.10 and in each figure the film thicknesses of one of the greases, plotted as a function of time, are compared with those GI and base oil 0, given by the oil from which the grease was made. Figs 1.6-1.9 relate to greases GI, G,, GB, and G4 respectively, while Fig. 1.10 shows results obtained with grease G, at the lower load. In all cases the films formed by the greases were initially thicker than those of their component oils by an amount that appears to be consistent with the Vol 184 Pt 3F Proc lnstn Mech Engrs 1969-70 at TUFTS UNIV on December 8, 2015pcp.sagepub.comDownloaded from 8 A. DYSON AND A. R. WILSON D-c OIL 01 thickening power expected for a uniform suspension of rigid soap particles. The greater the viscosity of the base oil, the rotational speed, and the soap content of the grease, the thicker the initial film. The difference in penetration numbers of greases G, and G, was not reflected by any marked difference in behaviour. A comparison of Fig. 1.7 with Fig. 1.10 shows that, as would be expected on theoretical grounds, the effect of load on film thickness was not large. Whereas the thickness of the oil films was, in general, approximately constant throughout a test, the grease films generally became thinner with time at both speeds and loads and the thickness rapidly became less than that of the corresponding oil film. With one exception the grease films seemed to reach equilibrium at a thickness approximately 40 per cent of the initial value. At any stage during an experiment a further application of grease would bring the capacity down to the initial value, i.e. the film thickness up to the initial value. An exception to the general pattern of behaviour of the greases is shown in Fig. 1.6 by grease GI. The initial rate of decrease in film thickness was smaller than that observed with the other greases at the same rotational speed and the film thickness subsequently increased to approximately the initial value. The significance of this increase, which was also observed with the oil on which the grease was based, is discussed later in this paper. As shown in Figs 1.7-1.10, the thickness of the films formed by the oils generally remained approximately constant with an apparent tendency to increase slightly with time", " This tendency, however, probably reflects a reduction in capacity caused by a progressive partial emptying of the inlet section of the Hertzian contact rather than a real change in thickness. At the start of a test the inlet must be full, but it was observed that the film thickness calculated at the end of a test on the \u2018empty inlet\u2019 hypothesis was approximately equal to that calculated at the start of the test on the \u2018full inlet\u2019 assumption. There were two exceptions to the general observation that oils maintain films of approximately constant thickness. First, Fig. 1.9 shows that the most viscous oil, 0,, at the higher speed formed a film which decreased in thickness with time to half the initial value. This may be related to the effect of centrifugal force in throwing off the oil against the surface tension. It occurs when the initial thickness, h,/2, of the film on the circumference of each disc is of the same order as the theoretical maximum equilibrium film thickness, h,, that can be retained by the surface tension against centrifugal force. This thickness is given by the relation where T is the surface tension, p the density of the lubricant, and u is the peripheral speed of the disc. The second exception is seen in Fig. 1.6 with an oil of low viscosity, 0,. With this oil the decrease in capacity with time was five times greater than that attributable to emptying of the inlet section, and an explanation must be sought for an apparently real increase in film thickness. A tentative suggestion is that the most volatile constituents of a light oil may evaporate from a film 0.25 pm (10 pin) thick on the surface of a disc at 60\u00b0C rotating in air at a peripheral speed of 6.4 m/s (21 ft/s). Such evaporation would lead to a gradual increase in viscosity of the reserve of oil in the inlet to the Hertzian contact and hence to an increase in film thickness. Conversely, an oil from which the most volatile constituents had been removed previously would not be expected to behave in the same way. This was confirmed by a test run on oil 0,. This is a mineral oil from which the light ends have been stripped and it is intended for use in high-vacuum systems. Its vapour pressure at 60\u00b0C is 2.1 x torr compared with 2 . 3 ~ torr for O,, although the viscosities of the two oils are similar. Fig. 1.11 shows that oil O4 maintained a film of constant thickness, and this result supports the suggestion that the increase of the film thickness with time, experienced with 0,, was due to selective evaporation of the lighter fractions. Results obtained under procedure 2, i.e. with speed cycled between 1600 and 400rev/min, are shown for hl N Tlpu2 Proc lnstn Mech Engrs 1369-70 at TUFTS UNIV on December 8, 2015pcp.sagepub.comDownloaded from FILM THICKNESSES IN ELASTOHYDRODYNAMIC LUBRICATION OF ROLLERS BY GREASES 9 ( a ) GREASEG4 *-4 4 0 0 r p m. H 1600 r.p rn. b. --%.. 100pzzzy--l C-4 4 0 0 r p m 7(~)o1~~~ b-4 400rpm. 2.5 2.0 E ? 15 8 W z Y C) t 10 5 3 0 5 G. 0 2 0 E 4 3 0 5 G: 0 3 100 150 TIME, minutes 60\u00b0C; 2200 lbf load. (a) grease G, and ( b ) base oil 0, Fig. 1.12. Effect of variation of speed on film thickness of Fig. 1.13. Effect of variation of speed on film thickness of (a) grease G4 and (b) base oil O3 greases G, and G, in Figs 1.12 and 1.13. The thickness of the oil films always follows the change in speed with a delay that increases with increasing oil viscosity. The effect of changes of speed on the thickness of the grease films gradually decays with time, the final thickness tending towards an equilibrium value which is approximately equal to the value that would have been reached if the machine had been operated continuously at the lower speed. At the start of the second and third periods at the higher speed in the run on grease G, (Fig. 1.12), the film thickness recovered temporarily to almost its initial value at the beginning of the first period, but rapidly declined towards the end of the periods. The explanation for this occurrence is not known. In this work it was noted that the effect of stopping and restarting the disc machine in the unloaded condition was very small both for oils and greases. However, the effect of stopping and restarting under load, as in procedure 3, was more marked and is shown for greases G, and G3 in Figs 1.14 and 1.15. The general effect of repeatedly stopping and restarting under load was to accelerate the reduction in film thickness with time that would occur in continuous running. In one case, with Y) m 2 0 E 5 75 0 4 - E 15 2\" w 2 x50 u w z 10 0 3 0 5 i; GI I + 25 5 0 0 1 2 3 4 5 6 7 8 9 1 0 I 2 3 4 5 6 7 8 9 1 0 CYCLE NUMBER 60\u00b0C; 2200 Ibf load; 1600 revlmin; 15 min running periods. film thickness of (a) grease G, and ( b ) base oil O2 Fig. 1.14. Effect of repeated shut-down under load on 100 (0 5 .E 75 E mw 50 0 D .- m 2 x I k- 25 0 u I 2 3 4 5 6 7 8 9 1 0 60\u00b0C; 2200 Ibf load; 1600 rev/min; 15 min running periods. film thickness of (a) grease G, and (6) base oil O3 Fig. 1.15. Effect of repeated shut-down under load on grease G,, the decrease in film thickness caused an increase in frequency of electrical contact between the discs such that measurements of capacitance became impracticable after the ninth shut-down. An exception to the general pattern of behaviour is observed in Fig. 1.14 with oil 0, which recovered its initial value of film thickness after a delay of up to 5 min after a restart. Many bearings run for years on the same charge of grease, but the maximum duration of the tests reported here was 5 hours. The results, therefore, refer only to a very small fraction of the service life. Furthermore, the quantity of grease applied to the discs was much smaller than that contained in a bearing. Within these limitations it is believed that the work gives some evidence about the nature and behaviour of the lubricant in the greaselubricated bearing", " The fact that grease films are initially thicker than the corresponding oil films, and the very different time behaviour, show that the lubricating agent is something other than the oil alone. The fair agreement between the observed and expected thickening effect of the soap suggests Proc lnstn Mech Engrs 1969-70 2 Vol 184 Pt 3F at TUFTS UNIV on December 8, 2015pcp.sagepub.comDownloaded from that the lubricating medium consists of a suspension of soap in the oil, the initial fibrous structure of the soap having been broken down by shear into a suspension of particles which are approximately spherical in shape. Fig. 1.5 shows an electron micrograph, at a magnification of x 2 5 000, of a carbon replica of soap particles on the surface of a disc after 15 min running in the disc machine. The soap in the original grease was arranged in a fibrous structure, as shown in Fig. 1.4, but this has been almost completely destroyed by working in the machine. The initial thinning of the grease films with continued running may be due to the progressive breakdown with shear of the structure of the grease, but this cannot be the complete explanation since the films formed by greases eventually become thinner than those formed by the corresponding oils. An explanation of this observation is offered along the following lines. The maximum shear rate in the inlet zone gradually increases as the lubricant approaches the conjunction", " In the disc machine this migration would cause a partial emptying of the inlet section by leakage towards test oil o f 5%wt Oppanol B50 the sides of the discs, and a subsequent decrease in film thickness. Evidence in support of the importance of normal stress differences was obtained by testing a solution of 5%wt polyisobutene (Oppanol B50 weight, average molecular weight 365 000) in oil O2 with l%wt anti-oxidant additive. Such a solution is viscoelastic and would show normal stress differences in shear. The thickness of the film formed in the disc machine by this lubricant decreased with time in a manner similar to that observed for greases, as shown in Fig. 1.16. The observation that the films rapidly become thinner than those given by the base oil alone again shows that shear breakdown of the polymer cannot be a complete explanation. The authors wish to thank Mr H. D. Moore for his expert guidance on the characteristics of greases, Messrs K. Greenway and G. W. Turner for preparing the greases and for a lot of unreported but essential preliminary work, Messrs R. B. Bird and G. Rooney for the preparation of the electron micrographs, and, particularly, Mr W" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003570_j.triboint.2011.09.004-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003570_j.triboint.2011.09.004-Figure3-1.png", "caption": "Fig. 3. Elastomer model: (a) a single elastomer slice associated with a half column of di", "texts": [ " In both cases during sliding the surfaces of the elastomer and the counterpart are separated by the local clearance h(x1, x3), given as: h\u00f0x1,x3\u00de \u00bc c\u00fect\u00feh0\u00f0x1,x3\u00de\u00fed\u00f0x1,x3\u00de \u00f02\u00de where h0(x1,x3) is the local depth of the dimple (see [18]) and ct is local additional clearance along the tapered edges: ct \u00bc ht x1 lt 1 for x1o lt ct \u00bc 0 for ltox1o l1 lt ct \u00bc ht x1\u00fe lt l1 lt for x14 l1 lt 8>>< >>: \u00f03\u00de It should be emphasized that the local film thickness h varies in both x1 and x3 directions since the dimples have the form of spherical segments distributed on the surface lying in the plane x1x3. Besides the film thickness varies in the x1 direction due to the tapered edges. Further clarification of the coordinate system will be given below during the discussion of Fig. 3. An additional scheme for the corresponding problem showing a shaft elastomer seal with the coordinate system is presented in Fig. 4. A stationary elastomer sleeve with the cross-section shown in Fig. 1 is fitted on a rigid smooth shaft moving axially in the x1 direction (Fig. 4(a)). A regular surface texturing is applied to the inner surface of the sleeve as shown schematically in Fig. 4(b). The x2 axis corresponds to the radial direction and the x3 axis corresponds to the circumferential direction", " The cavitation pressure in potential cavitation regions is pa (Reynolds cavitation condition). Reynolds condition is used in cavitation zones due to its simple incorporation into most numerical schemes of hydrodynamic lubrication [25] and good correlation with experimental data (see e.g. [10,26]). Similar to the pressure distribution the elastic deformations are also periodic and symmetric in the x3 direction with the same half period of r1. Hence, it is sufficient to solve Eq. (5) for a single elastomer slice of width r1 associated with a half column of dimples (see Fig. 3(a)). The boundary conditions for Eq. (5) are therefore: 1. The normal and shear stresses on the elastomer interface with the fluid film (textured surface of the elastomer) are equal to the hydrodynamic pressure and viscous shear stress, respectively. Expressing these stresses in terms of displacements, the corresponding boundary conditions may be written as follows (see [18\u201320]): u2,2\u00fe n 1 2n uk,k X2 \u00bc 0 \u00bc 1 E \u00f01\u00fen\u00deP\u00f0X1,X3\u00de \u00f09a\u00de \u00f0u1,2\u00feu2,1\u00de X2 \u00bc 0 \u00bc 2 E \u00f01\u00fen\u00de 1 6H\u00f0X1,X3\u00de a \u00f09b\u00de where a\u00bc0 accounts for the vanishing stresses in cavitation zones and a\u00bc1 relates to full fluid film zones", " For the elastomer deformation problem with well defined boundary conditions finite element method is very effective and popular (compared to the cumbersome procedure of solving the problem of hydrodynamic lubrication for regions including the cavitation zones). For the purpose of the elastomer deformation problem solution the commercial software ANSYS 11.0 was used. The elastomer was meshed using three dimensional 10-node finite elements (SOLID186) with non uniform grid. Relatively coarse mesh was utilized near the built-in bottom plane of the elastomer and the finest mesh was near the textured plane (see Fig. 3(b)). Accurate convergence was tested by mesh refining. The mesh for the hydrodynamic problem does not exactly map to the surface mesh of the elastomer deformation problem. This required interpolation of hydrodynamic lubrication and elastomer deformation problems results and mapping between the two meshes, which leads to numerical errors stemming from interpolation. However, the choice of sufficiently detailed grids for both problems minimizes this error. The prediction-correction approach for consecutive numerical solution of Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002622_cdc.2007.4434589-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002622_cdc.2007.4434589-Figure2-1.png", "caption": "Fig. 2. Instrumented experimental platform at UNSW@ADFA - YAMAHA RMAX helicopter", "texts": [ " This paper discusses the backstepping algorithm as a result of varying the tuning variables to control the system overshoot and damping. The details of the helicopter model are not included in this paper. Only the important equations of the helicopter model are included. These equations provide a connection between the rigid body dynamics and helicopter aerodynamic forces. Interested readers may refer to [1], [2], [3] for a detailed understanding of the helicopter model. The experimental platform [4] (RMAX helicopter) at UNSW@ADFA is shown in Figure 2. A detailed platform descriptions and sensor enhancements are given in [4]. The platforms have the capability to fly autonomously. For this purpose, platforms are well equipped with state-of-theart sensors and instruments to measure position, velocity, angles and rates in the body coordinates. In this paper we assumed that the platforms under consideration can measure position, velocity, angles, rates and gravity vector in the body coordinate frame in real-time. This capability of the onboard sensors and instruments plays a vital role in achieving our overall objective", "005) demonstrate that carefully chosen \u03b1 and \u03b2, enable the UAV to hover at the equilibrium point, starting from a nearby initial position. It is clearly shown in the results that one can achieve the desired position in 2.5\u223c3.0 seconds depending on the values of \u03b1 and \u03b2. The controller was successful in stabilizing the system. We used LevenbergMarquardt method to solve four equations in four unknowns given in equations (7)\u2013(8). V. PLATFORM AND SIMULATION PARAMETERS The values of physical parameters of the helicopter used in this simulation are given in Table I. Parameters Mx,Mz ,Tx,Tz are shown in Figure 2. VI. CONCLUSION In this paper position control of a UAV using backstepping is presented. For nonlinear control design, backstepping technique has several advantages for underactuated mechanical systems. This paper presents an insight into the dynamics of the system while at the same presenting a controller without adaptation but the algorithm has the capability for extension to adaptive control. The future work is to include the model uncertainty for a UAV such that the environmental factors are taken into account for practical applications" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003887_19346182.2012.663534-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003887_19346182.2012.663534-Figure2-1.png", "caption": "Figure 2. (a) Racquet head approaches a stationary ball at speed V and angle A. (b) The ball emerges at speed v and with topspin. In a kick serve, the angle A is only a few degrees, but is shown here as a relatively large angle for clarity.", "texts": [ " There is only one spin axis and it is tilted away from the vertical. If the axis in Figure 1 is vertical then there is no topspin, just sidespin. If the axis is horizontal then there is no sidespin, just topspin. If the axis is tilted then any point on the ball will rotate in a circle around the axis, and it rotates simultaneously in vertical and horizontal directions. The amount of topspin generated in a serve due to vertical motion of the racquet head can be estimated by considering the situation shown in Figure 2. The racquet head is vertical and is approaching the ball rapidly at speed V and angle A. The ball can be regarded as being stationary, although it may in fact be falling slowly as a result of the ball toss. We can ignore sideways motion of the racquet head in order to calculate the amount of topspin. Sideways motion is a separate issue and it generates sidespin, in the same way that vertical motion of the racquet head generates topspin. After the racquet head strikes the ball, the ball will emerge at speed v and with topspin as in indicated in Figure 2(b). The ball emerges at high speedtowards the net and is shown heading upward since that is the effect of the upward friction force of the strings acting on the back of the ball. The friction force must act in an upward direction to generate topspin. In order to estimate the amount of spin in Figure 2, we can consider the collision in a reference frame where the racquet is initially at rest and the ball approaches the racquet at speed V, and angle A as indicated in Figure 3. Measurements of ball spin have previously been made for a hand-held racquet (Cross, 2003a; 2005) and it was found that Rv is typically about equal to vx2 where R is the ball radius and vx2 is the tangential component of the ball speed after the collision. The condition Rv \u00bc vx2 corresponds to a rolling ball condition. It was also found that vx2 \u00bc 0:7vx1 to a good approximation and that vy2 is typically about 0.4vy1 depending on the impact point on the strings. The outgoing ball spin is therefore given to a good approximation by: Rv \u00bc vx2 \u00bc 0:7vx1 \u00bc 0:7V sinA \u00f01\u00de The outgoing ball spin in Figure 3 is the same as that in Figure 2 and increases with both the speed of the racquet head and the approach angle. For example, if V \u00bc 40 m s21, A \u00bc 58 and R \u00bc 0.033 m then v \u00bc 74 rad s21 (706 rpm). If the approach angle is zero then the outgoing ball spin is zero. Hitting up at a greater approach angle generates more topspin, but the ball is then launched at a higher angle over the net and may land beyond the service line. The latter result follows from the fact that the rebound angle B in Figure 3 increases when the angle of incidence increases. The result, when transformed back to the reference frame in Figure 2, is an increase in the launch angle as the approach angle, A, increases. Figure 2 also describes the result when the racquet is moving sideways across the back of the ball, and is approaching the ball at a sideways angle A. In that case, the ball acquires sidespin, and the amount of sidespin is given by the same expression. In practice, the racquet head usually approaches the ball as shown in Figure 1, with a large sideways approach angle and a relatively small vertical approach angle. D ow nl oa de d by [ U ni ve rs ity o f D el aw ar e] a t 1 7: 40 2 7 Ju ly 2 01 3 As a result, the ball is usually served with about 4000 rpm of spin in a kick serve, but the spin is mostly sidespin and the amount of topspin is relatively small. That is, the spin axis is almost vertical, as indicated in Figure 1. Suppose a racquet approaches a ball in a horizontal direction at speed V and the ball is falling vertically at speed v just prior to impact, as shown in Figure 4(a). In a reference frame where the ball is at rest, the racquet is rising vertically at speed v while simultaneously moving horizontally at speed V, as shown in Figure 4(b). The situation is then the same as that shown in Figure 2 and the spin is given by Equation (1). If the ball falls say one metre before it is struck then it will be falling at 4.43 m s21 when it is struck. In a kick serve, V is typically about 40 m s21. The racquet approaches the ball at an angle A given by tan A \u00bc v=V \u00bc 0:11 in this case, so A \u00bc 6.38. That angle could well be larger than the actual vertical approach angle of the racquet in a typical kick serve, in which case the ball toss would account for more than half of the topspin generated. From Equation (1), we find that v \u00bc 93 rad s21 (889 rpm) due to the ball toss alone" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001059_1.1839922-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001059_1.1839922-Figure2-1.png", "caption": "Fig. 2 Test rotor supported on three lobe test bearings showing eddy current sensors and infrared tachometer", "texts": [ " A K-type thermocouple monitors the temperature of the motor armature. A rapid rise in motor temperature is a good indicator of sustained solid contact between the rotor and the test bearings. Each bearing housing has two O rings that effectively seal the bearing section and allow that section to be pressurized. The test chamber therefore has six sections with supply air inlet and outlet fittings into each section. Pressure gauges and flow meters record the air flow conditions into each bearing plenum chamber. Figure 2 shows the end view of the test rig displaying two orthogonal positioned highly sensitive, eddy-current sensors that measure the displacement of the shaft at each end.5 The displacement sensors have a sensitivity of 39.4 mV/ m 1 V/mil with a linear range of 400 m. The sensor recorded voltages are conditioned to remove the large dc bias offset before connection to two separate oscilloscopes and/or the data acquisition system. The oscilloscopes display the rotor orbit at the end monitored. Force piezoelectric sensors mounted between the bearing housing and the test chamber alignment bolts measure the load transmitted through the bearings as depicted in Fig", " The static and dynamic response of the rotor supported on the three lobe bearings characterizes the overall performance of the rotor/bearing system. The feed pressure required to lift the rotor characterizes the static performance. Analysis of the rotor coast 3The difficulty with obtaining the appropriate terminal voltage is the speed limiting factor in this application, not the performance of the gas journal bearings. 4Note that balancing the rotor to reduce remnant imbalance proved very difficult. 5Experimental results refer to the left L and right R bearings with respect to the view in Fig. 2, having vertical V and horizontal H sensors at each bearing, i.e., LV refers to the rotor response recorded with the left vertical displacement sensor, for example. Journal of Engineering for Gas Turbines and Power JULY 2006, Vol. 128 \u00d5 627 Downloaded From: http://gasturbinespower.asmedigitalcollection.asme.org/ on 01/08/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use down response to calibrated imbalance masses, for increasing supply pressures, determines several performance characteristics" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003360_1.3610462-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003360_1.3610462-Figure1-1.png", "caption": "FIG. 1. (a) Configuration for electrowetting on a lipid bilayer on hafnium oxide. A silicon electrode with a hafnium oxide film is submersed in dodecane with sorbitan trioleate. A water drop is placed on the hafnium oxide and contacted by a platinum electrode. A voltage is applied across the hafnium oxide and a spontaneously formed lipid bilayer. (b) Schematic depicting the lipid bilayer composed of sorbitan trioleate. The lipid bilayer acts as a capacitor in series with the hafnium oxide.", "texts": [ " Spontaneously formed in the presence of appropriate oil surfactants, these bilayers facilitate high contact angle changes at low voltages, comparable to those observed in mercury/oil/water electrocapillary systems without dielectric layers.11 Electrowetting alters the contact angle of a liquid drop on a solid surface via voltage application across the solid-liquid interface. A typical electrowetting configuration consists of a water drop on a dielectric stack covering a planar electrode, as shown in Fig. 1(a). The Lippmann-Young equation predicts the contact angle change in response to voltage,10 cosh\u00f0V\u00de \u00bc coshY \u00fe CV2 2cwo ; (1) where h\u00f0V\u00de is the contact angle of the water drop as a function of voltage, hY is the angle at zero voltage (known as Young\u2019s angle), C is the capacitance per area of the dielectric, V is the voltage applied between the electrode and the water drop, and cwo is the surface energy between the water and the nonconductive ambient phase, commonly air or oil. Most electrowetting systems rely on spin-coated amor- phous fluoropolymers such as Teflon AF TM or Cytop TM on top of inorganic dielectrics to obtain reversible wetting", " When a water drop is deposited onto hafnium oxide, sorbitan trioleate adsorbs at the oil-water and oil-hafnium oxide interfaces, a)Author to whom correspondence should be addressed. Electronic mail: iguha@alum.mit.edu. 0003-6951/2011/99(2)/024105/3/$30.00 VC 2011 American Institute of Physics99, 024105-1 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: 138.237.77.209 On: Thu, 11 Dec 2014 21:20:06 forming a lipid bilayer, depicted in Fig. 1(b). The disjoining pressure of the bilayer buffers the water from the hafnium oxide. Surfactant bilayers are commonly used to stabilize waterin-oil emulsions17,18 and have even been used to electrowet two water drops against each other.19 However, this effect has never been exploited for reversible electrowetting of a water drop on a solid, hydrophilic dielectric surface. The bilayer is surprisingly stable and the film does not fracture into smaller droplets upon electrowetting, as has been observed for oil films trapped between aqueous drops and fluoropolymer dielectrics", " The TEMAH precursor was held at 80 C and delivered using 300 sccm Ar carrier gas. Ellipsometry measurements indicated that the resulting hafnium oxide film was approximately 9 nm thick. The dielectric constant for the film measured approximately 17. To characterize the electrowetting behavior, a degenerately boron doped single crystal silicon electrode coated with a 9 nm hafnium oxide dielectric was submersed in dodecane containing sorbitan trioleate, and a water drop ( 3 to 4 mm diameter) was deposited on the hafnium oxide surface, as shown in Fig. 1(a). A platinum electrode contacted the water drop and a voltage was applied to the water drop while the silicon remained grounded. Applied voltage in the form of an AC sine wave from V to V at 300 Hz was incremented from 0 V to a maximum voltage (forward sweep), resulting in wetting, and then decremented back to 0 V (reverse sweep), resulting in dewetting. The water drop contact angle (h), the capacitance between the grounded substrate and the water drop (C), and the diameter of the contact area between the water drop and the hafnium oxide were recorded for each applied voltage", " The low oil-water surface tension contributes to the distortion of the drop shape, potentially resulting in underestimations of the contact angle, particularly at high angles. It is possible to measure the thickness of the lipid bilayer using capacitance measurements. The entire dielectric can be 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: 138.237.77.209 On: Thu, 11 Dec 2014 21:20:06 modeled as the capacitance of the hafnium oxide in series with the capacitance of the bilayer, as shown in Fig. 1(b). The capacitance per area of a 0.1 N NaCl water drop on hafnium oxide measures higher in air (without the bilayer) than in dodecane with sorbitan trioleate (with the bilayer). Accounting for the difference between these two capacitances, the effective bilayer thickness can be estimated. For the partial wetting system, the capacitance per area of the bilayer (Coil), and the effective bilayer thickness (toil) at different voltages reveal that the bilayer thins as the voltage increases, as seen in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002332_j.mechmachtheory.2009.05.007-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002332_j.mechmachtheory.2009.05.007-Figure7-1.png", "caption": "Fig. 7. (a) Book facing page of Redtenbacher\u2019s \u2018\u2018Motion Mechanisms\u201d. (b) Drawing of mechanism from \u2018\u2018Motion Mechanisms\u201d.", "texts": [ " The novelty and complexity of this technical system required scientific penetration and from March 1854, he declared in several letters that specifically this technology driven by the English with their long empirical knowledge would require a scientific basis. He thereby entered the new field of inertia forces, vibrations and impact loading, and was not able to cope with the complex driving system called a locomotive, in all its detail. Critical comments by experts in this field could not be ignored. His series of books was extended in 1857 by the contribution \u2018\u2018Motion Mechanisms\u201d [29] (in 1861 it was expanded by a new issue [31] to be collected in 1866 as one volume, see Fig. 7), which essentially presented and explained the teaching models aids collection of the mechanical engineering section constructed in the workshops of the Polytechnic school under Redtenbacher. These models supported his lectures to the professional community from models of a simple pair of wheels or a slider crank mechanism through kinematic models of the complex non-uniformly transmitting gear boxes. 9 its in Box 1. The Redtenbacher Machine Models Collection at Karlsruhe In contrast to the present use of computer simulation and animation in teaching kinematics of machines, engineering professors developed precise 3D mechanical models to help students visualize complex motions in machines as well as illustrate the mathematical curves and functions necessary for scientific design of machines" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000948_detc2005-84109-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000948_detc2005-84109-Figure4-1.png", "caption": "Figure 4. Double slider-crank in a) permanent and b) instantaneous critical form", "texts": [], "surrounding_texts": [ "In this section it is hinted the use of logical functions for d.o.f. computation of mechanisms with variable kinematic structure or intermittent motion. We assume that the matrix method is adopted. Logical functions are an useful mathematical tool for modeling the kinematics and the dynamics of intermittent and variable kinematic structure mechanisms. For our purposes, the occurrence of discontinuities of kinematic structure can be treated by introducing ad hoc logic conditions that regulate the type and the number of kinematic constraints that must be taken into account. Let L(x) be a continuous function 6 nloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx? L(x) = 1 2 |x|2n+1 + x2n+1 |x|2n+1 + 1 2 [ |x\u2212 \u03b5|2n+1 \u2212 (x\u2212 \u03b5)2n+1 ] = 0 x \u2264 0 1 2 y = \u03b5 2 1 x \u2265 \u03b5 (20) where \u03b5 > 0 is the amplitude of the transition interval (see Figure 2) from one state to another and n is chosen so as to assure continuity of any derivative, of order d, which will be true if 2n + 1 > d . (21) Equation (20) approximates the ideal Heaviside step function H(x) H(x) = { 0 se x < 0 1 se x > 0 . (22) The first and second derivatives of (20) give an approximation of the \u03b4 Dirac\u2019s and doublet functions, respectively. The step function at abscissa x = xa is obtained substituting in (20) (x\u2212 xa) at x. Several investigations confirmed the reliability and accuracy of dynamic analysis results through the use of logical functions." ] }, { "image_filename": "designv11_20_0001679_gt2008-50305-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001679_gt2008-50305-Figure6-1.png", "caption": "Figure 6. Turbine rotor, as separated from rotor group, side view of cast portion (A), aft view (B), and side view of compound machined assembly (C).", "texts": [ " The investigation team interpreted this to indicate that the problem was not inherent in the design of the engine and its aerodynamic components. Engines that were apparently identical had nonidentical results. The team initiated an extensive review of the turbine rotor fleet and made comparisons of the good and bad turbine rotors. The turbine rotor of the microturbine is similar to other small turbomachinery products, as seen in Figures 5 and 6. There is a Mar-M 247 cast radial turbine wheel with an Inconel 718 tiebolt inertia welded to the backface, visible in Figure 6. The final machining is completed on this compound, welded assembly. The good and bad rotors, with respect to sub-synchronous vibration, were consistent for meeting the existing drawing tolerances, with the critical machined features held to within 5 \u00b5m (~.0002 in.) and the cast features conforming to typical aerospace tolerances for investment castings. It was then noted that the majority of the bad rotors were received with a high initial unbalance. It must be emphasized that the quality of the balance work and the amount of residual unbalance at time of assembly was identical for all rotors, good or bad" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002495_icca.2009.5410535-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002495_icca.2009.5410535-Figure6-1.png", "caption": "Fig. 6. Sweep coverage with autonomous vehicles (n = 6, r = 1, s = 0.5) along a curve: initial positions (circles) and vehicle trajectories (solid lines).", "texts": [ " One can also see that the distance between vehicles converges to 0.5 which equals to the parameter s. Similar results have also been obtained for \u03c6\u0304 = \u03c0/4 and \u03c6\u0304 = 0, and they are shown in Figure 4\u20135, respectively. In reality, the line W that we wish to let the vehicles sweep along may not be straight, for instance, the coastline of a nation is not generally straight. To show that our proposed algorithm can also provide sweep coverage along a curved line, we tested our algorithm using a curved line as shown in Figure 6 and the results demonstrate the flexibility of our control algorithm. Besides curved lines, we also examined the effectiveness of the proposed algorithm on a right-angled corner. This scenario may arise when one wants a group of security robots to patrol around the perimeter of a building or structure. The results of sweep coverage around a rightangled corner are illustrated in Figure 7. Figure 6 and 7 show that the straight-line assumption can be relaxed in some situations. In this paper, we have proposed a set of decentralized control laws for the coordination of a group of autonomous vehicles to address sweep coverage along a given line or path. The coordination or control rules were developed using the information consensus approach, that is simple and can be easily implemented in real-life situations. To illustrate the effectiveness of the control algorithm, a number of numerical simulations were performed" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002690_978-3-540-30301-5_19-Figure18.9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002690_978-3-540-30301-5_19-Figure18.9-1.png", "caption": "Fig. 18.9 Robotic biomanipulation system with vision and force feedback", "texts": [ " However, it lacks a means to hold the cell in place for further manipulation, such as injection, since the magnitude of the electric fields has to be kept low to ensure the viability of cells. The limits of noncontact biomanipulation in the laser trapping and electrorotation techniques make mechanical micromanipulation desirable. The damage caused by laser beams in the laser trapping technique and the lack of a holding mechanism in the electrorotation technique can be overcome by mechanical micromanipulation. To improve the low success rate of manual operation, and to eliminate contamination, an autonomous robotic system (shown in Fig. 18.9) has been developed to deposit DNA into one of the two nuclei of a mouse embryo without inducing cell lysis [18.11,128]. The laboratory\u2019s experimental results show that the success rate for the autonomous embryo pronuclei DNA injection is dramatically improved over conventional manual injection methods. The autonomous robotic system features a hybrid controller that combines visual servoing and precision position control, pattern recognition for detecting nuclei, and a precise autofocusing scheme" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000390_robot.2004.1308068-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000390_robot.2004.1308068-Figure8-1.png", "caption": "Fig. 8. Experimental Setup: a pair of ZWF finger robots grasping a rectangle abject", "texts": [], "surrounding_texts": [ "In this case control input is given by U ; = u,i + u g i and the observed object angle 6 is used in us,. Desired angle is given as -5 [deg]. It is found that 6 converges to the desired values. Yl - Yz also converges to zero so that the stable grasping and the absolute angle control are established simultaneously.\n1 \" -\nC. Control of relative object angle\nFigure 6 shows the relative angle control of the grasped object. Control parameters at that time shows in Table I. Until l[s] the finger robots are controlled by ui = U,<, and after that, u g i is added in U$. In this case Eq.(37) is used instead of the observed object angle. The initial joint angle gor is selected when the time is 0.99: [SI, just before ugi is added and the desired relative aigle 6 d = -5 [deg]; The results shows that the relative angle control is realized and the stable grasping is also established since Yl - Yz converges to zero.\n1 is 1024 pules per revolution. The joint velocity can be gained by numerical differentiation. The obiect angle is measured by a A\nV. Experimental results\n-3 4 .\n-5 - 2 6 - g -7\n-8\n-9\nw\n0 0.5 1 1.5 2 2.5 3 3.5 4 tISl\nThis paper constructed new sensory feedback inputs for stable grasping, and absolute and relative angle conhul of an object with parallel surfaces in horizontal plane. This - - I\n\" J pa of laser deflection sensors. The-bar wcth the object shown in Fig. 7 is attached to reflect the lasers. - = 'I/\\ AW-1\nIf stable grasping has already been realized, relative object angle control can be realized without object sensing . And the effectiveness of these controls were shown through the Fi 5 -k irad] , M = 0.1 Lkgl With sensor information of the object angle E fd = 1.01N1. Ed =", "151 S. Arimoto. M. Yoshida. J.-H. Bae. and K. Tahara. Dynamic fordtorque balance of ?D polygonal objects by a rolling canfacts and sensory-motor\n161 J . Baumgane. Stabilization of con~trilnt~ and integrals of motion in dynamical systems. Computer rnerhods in applied mechanics orid enginerring. l(lhI-l6, 1972. 171 A. Biechi. Hands for dexterous manipulation and robust grasping: A difficult mad towards simplicity. lEEE Trmsocriortr on Robotics ond Automotion, 16(6):652-662, 2000. 181 A. 8. A. Cole. J. E. HBUSIT, and S. S. Sastry. Kinematics a d control of multifingered hands with rolling ~ontacl. IEEE Tmmsctionr on Aulornalic Cortrml. 3 4 ( 4 ) : 3 9 8 4 , 1989. 191 I. Kao and M. R. Cuikorky. Quasistatic manipulation with compliance and sliding. The lnrernorioriol Journol of Roboricr Research. I l ( l ) :2& 4m 1992\n'coordination. Jounial ofRoboric SYstemr, 20(9):517-537. 2CHX\nsimulations and :he experiments\n.. 1. Kerr and B. Rath. Analysis of multifingered hands. Tlw Inlernational J o u m l ofRoboricr Resenmh. 41413-17, 1986. H. Maekawa. K. Tank, and K. Komoriya. Tactile sensor based manipulation of an unknown abjectmultifingered hand with mlling contact. In Pmc. of IEEE Internotional Conference on Roborics and Aurmwtion.\nReferences\n[I] S. Arimoto. Contml ,heow of norr~lirieor mechonicol rysrrrns L?\npossivih-bosed and circuir-rhroreric oppmmh. Oxford Science Publications. Oxford Univeniiy Press. 1996. 121 S. Anmoto. H. Hashiguchi, and R. Ozawa. A simple contml m e a d coping with a hemat ica l l y ill-posed invene problem of redundant rnbols: Stability on a consiraim manifold. Arioa Joounrol of Conrml, submitted. 131 S. A\"rnoto, K. Tahara, P.T.A. Nguyen. and H.-Y. Ha\". Ptinnciple of superposition for controlling pinch motions by means of robot fingen with soft lips. Roborico. 18. 21m. 141 S. Arimoto, K . Tahara, M . Yamaguchi, P.T.A. Nguyen. and H. Y. Ha\". Principle of superposition for controlling pinch motions by means of robot fingers with soft tips. Roborico. 19(1):21-28. 2uOI.\npages 743-750, 1995. Dynamic gnsping farcc\ncontrol using tactile feedbackmulitfinzered hand. In Pmc. of IEEE 112J H . Maekawa. K. Tanie. and K. Komoriya.\nImemariorid Confermce on RoDoiic$ ond Aumnorion, pages 2 6 2 - 2469. Minneapolis. Minnesota 1996. 1131 K. B. Shimoga. Robot grasp synthesis algorithms: A survey. The hzrarnorionai Joirnial of Roborics Rereorch. 15(31:230-236, 1996." ] }, { "image_filename": "designv11_20_0003147_ijmee.38.2.5-Figure11-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003147_ijmee.38.2.5-Figure11-1.png", "caption": "Fig. 11 Prony-brake dynamometer set-up. RPMs were measured using an optical tachometer pointing at the fl ywheel.", "texts": [ " Or, the smoothness of running was at USD & Wegner Health Science Information Center on April 9, 2015ijj.sagepub.comDownloaded from International Journal of Mechanical Engineering Education 38/2 at USD & Wegner Health Science Information Center on April 9, 2015ijj.sagepub.comDownloaded from International Journal of Mechanical Engineering Education 38/2 occasionally hampered by a non-perpendicular cylinder bore. Students then learned various techniques for overcoming such diffi culties and, in the end, all engines ran well. Engines were tested on a simple Prony-brake dynamometer (Fig. 11) for minimum pressure required to run (typically 1\u20132 psi), minimum speed at that minimum pressure (200\u2013300 RPM), speed at 30 psi (1200\u20133000 RPM), and shaft output torque and rpm at 30 psi, from which the students calculated power (3\u20136 watts). The kinematics of the inverted slider\u2013crank mechanism are analyzed in the Dynamics as well as Machine Design courses. Power calculations build on concepts in Physics and Dynamics. Power output and valve port fl ow were directly related to theoretical concepts in Thermodynamics and Fluid Mechanics" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001000_tmag.2004.840315-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001000_tmag.2004.840315-Figure4-1.png", "caption": "Fig. 4. Analysis model.", "texts": [ " The flow path resistance in the circulation port and adjacent narrow gap parts is substantially lower than that in the gap of the radial bearing, with the result that the pressure at the circulation port is near atmospheric pressure. The pressure indicated by \u201cP\u201d in the Fig. 3 makes the circulation pressure, caused by the asymmetrical herringbone grooves creating strong axial pressure and axial flow through the bearing to purge the air bubbles. It is possible to obtain the flow rate by calculating the circulation oil pressure and flow path resistance using the analysis model in Fig. 4. Incidentally, the gap in the radial bearing is not always constant and may be inclined or tapered, which changes the circulation pressure. The broken black line graph in Fig. 3 is the result of assuming that the front and rear bore of the sleeves has a taper of 1 m, which results in the increase of pumping pressure. Fig. 5 shows the results of the circulation for flow rate. The flow rate (q) is obtained by dividing the pressure (P) indicated in Fig. 3 by the flow path resistance (r). In this paper, we use a general expression that assumes that the flow path resistance is proportional to the flow path length, the cube of the flow path gap, and inversely proportional to the flow path width. At low temperatures, the circulation pressure rises due to the increased oil viscosity, but the calculated flow rate changes only a little because the flow path resistance increases in proportion to the viscosity. The calculation result indicates that at a flow rate of about 0.1 L/s, the entire volume of the oil in the bearing, which is about 3 L, takes about 30 s to make a complete circuit. C. Analysis of Oil Sealing Capability According to the analysis model in Fig. 4, the oil in the oil reservoir (the oil reservoir of the slanted gap plus that of the slanted narrow gap) flows into the gap in the radial bearing as a result of surface tension. In Figs. 4 and 6, the symbol \u201cA\u201d indicates the radial bearing gap (2 to 3 m); \u201cB\u201d denotes the narrow cylindrical gap (20 to 40 m) between the sleeve and the outer surface of the upper part of the shaft; \u201cC1\u201d represents the narrow gap (20 to 60 m) between the outer surface of the shaft and the oil cap hole; \u201cC2\u201d symbolizes the opening part; \u201cD1\u201d represents the narrow gap near the shaft of the slanted gap; \u201cD2\u201d corresponds to the gap wider than D1 and further radially from the shafts of the slanted gap; \u201cE\u201d indicates the oil reservoir provided in ring form around the slanted narrow gap; and \u201cF\u201d implies the part where the gap is wide enough and which has a vent hole for discharging air" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000681_j.triboint.2006.01.006-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000681_j.triboint.2006.01.006-Figure1-1.png", "caption": "Fig. 1. Geometry of the lubricating film for an infinitely long contact on Oy direction.", "texts": [ " Elastohydrodynamic Lubrication\u2014governing equations The formulation of a thermal and non-Newtonian elastohydrodynamic problem involves the coupled Reynolds equation, the film geometry equation, the load balance equation, the energy equation and the lubricant rheological model. 2.1. Reynolds equation Assuming a Newtonian behaviour of the lubricant and taking into account the usual simplifications of the Navier\u2013Stokes equations applied to the elastohydrodynamic lubrication problem, the Reynolds equation for an infinitely long contact (see Fig. 1) can be expressed as [7,8] q qx qp qx rF 12Z h 3 \u00bc \u00f0U 1 \u00feU 2\u00de 2 q qx \u00f0rFh\u00de, (1) where the lubricant viscosity Z will be replaced by an apparent viscosity Zap as referred ahead. 2.2. Film geometry equation The lubricant in the contact neighbourhood is drawn inside being squeezed and stretched. The generated hydrodynamic pressure is high enough to cause appreciable elastic deformation of the solids surfaces and modify the lubricant viscosity and density. In defining the lubricant film geometry it is necessary to include the elastic deformation in the free load surfaces geometry. According to Fig. 1, the film geometry can be defined by [9]: h\u00f0x\u00de \u00bc hc \u00fe \u00bdu\u0304z\u00f0x\u00de u\u0304c \u00fe h 0 \u00f0x\u00de, (2) where u\u0304z is the normal displacement of both surfaces. The normal displacement of any surface point (see Fig. 1) being submitted to the pressure distribution p(x) is obtained by [10,11]: u\u0304zi\u00f0x\u00de \u00bc \u00f01 u2j \u00de pEj Z xmax xmin p\u00f0s\u00de ln\u00f0x s\u00de2 ds\u00fe C. (3) ARTICLE IN PRESS A. Campos et al. / Tribology International 39 (2006) 1732\u201317441734 The lubricant film geometry, defined in Eq. (2), can be rewritten as h\u00f0x\u00de \u00bc hc \u00fe h0 \u00f0x\u00de 2 pE 0 Z xmax xmin p\u00f0s\u00de ln\u00f0x s\u00de2 ds; (4) with E 0 \u00bc 2 1 u21 E1 \u00fe 1 u22 E2 1 . 2.3. Force balance equation The normal force applied to the contact is supported by the hydrodynamic pressure generated in the contact" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001242_12.659594-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001242_12.659594-Figure2-1.png", "caption": "Figure 2: Coordinate definitions in the closed form AX=XB.", "texts": [ " Thus for the purpose of QC, it is sufficient to recalibrate the system for the remaining (rotational and scale) degrees-of-freedom. This paper presents the concept, mathematical framework, experimental implementation, and in-vitro evaluation of a phantomless real-time method that detects intra-operative failures of the tracked US while recovering the calibration matrix. 2. MATHEMATICAL FORMULATION The key enabler of our self-calibration method is a closed-form mathematical formulation of the problem. Figure 2 presents the coordinate systems for the mathematical formulation. A1, A2 are the transformations of US picture coordinate system (P) with respect to the fixed construction frame (C) at poses 1 and 2, respectively. Note that the actual selection of C is arbitrary and the only requirement is that it must be rigidly fixed during the calibration process. Using A1 and A2, we obtain the transformation between poses 1 and 2, as A=A2A1 -1. At the same time, the transformation between the two poses can be recovered using a calibration phantom or recovered directly by matching the 2D ultrasound images acquired in these poses to a prior 3D model of the phantom object", "org/ss/TermsOfUse.aspx We have extended this solution method to account for inhomogeneous scale in the three coordinate axes [8,20]. Prior to this work, we have exploited this closed form formulation to solve the calibration problem based on various mechanical phantoms including the double-wedge phantom [21], z-shape phantom [8], and thin-wall phantom [20]. In all these methods, phantoms were built in a way to assist estimation of A\u2019s, which is the relative motion between successive US frames as shown in Figure 2. Our present task reduces to recovering A\u2019s as we are scanning real tissue and collecting the corresponding B\u2019s from the tracker, and then obtaining the calibration by solving the homogenous linear system in equation (6). Obviously, recovering A\u2019s from real-time US sequences is a tracking problem instead of feature segmentation from static US phantom images. Tracking the 6 DOF of A\u2019s based on 3DUS data is considered a straight forward problem. The main challenge lies in the full recovery of the A\u2019s based on 2DUS data", " Finally, the AX=XB solver receives corresponding A and B data, and recovers the X calibration matrix. The Quality Control unit analyzes the new calibration and compares it with previous runs. In case of suspected discrepancy, an appropriate Action is initiated to deal with a hazard condition. The action could range from generating a warning message to demanding a halt of the procedure and full recalibration of the system. Real-time Tracker As mentioned above, the role of Real-time Tracker is to recover the A matrices, the motion of the US image in construction frame, as it was described in Figure 2. What is necessary is to compute the relative motion in pairs of ultrasound images for which the absolute (tracked) motion is known. We accomplish this using direct image registration methods similar to those described in [25]. Specifically, we introduce an intermediate \u201cwarped\u201d image representation W defined as: )*)((),;,( puRotIptuW += \u03b1\u03b1 where u=(x,y)T is an image location, P is a translation offset, and \u03b1 is an interframe rotation. Let W(t; P, \u03b1) denote the column vector constructed by stacking the value of W for all possible image locations u" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.125-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.125-1.png", "caption": "Fig. 14.125 Single-drum sander for parquet floors", "texts": [ " The following types of machines can be identified: \u2022 Drum (single- and double-drum) sanders\u2022 Disk sanders\u2022 Oscillating sanders Drum sanders are intended for rough and finishing sanding of wooden floors. Sanders with one and two working drums can be distinguished. In the latter case, one drum performs sanding while the other stretches the abrasive belt. Both drums have a horizontal axis of rotation. The sander\u2019s working element is an abrasive (sandpaper or abrasive cloth). The sanding drum\u2019s working width is up to 250 mm and is limited by the necessity for the drum to exert appropriate linear pressures (about 20 N/cm) on the surface. The design of a single-drum sander is shown in Fig. 14.125. A drum sander may weigh as much as 90 kg and the power of the driving motor can reach 3.5 kW at a drum rotational speed of about 2300 rpm. Electric motors are exclusively used to drive sanders because the latter are intended mainly for work indoors and the possibility of ignition of the wood dust collected in dust bags has to be eliminated. The described drum sander cannot sand the floor under heaters. For this purpose disk sanders (Fig. 14.126) are used. The working tool in this machine is an abrasive disk mounted at an angle of about 3\u25e6 relative to the base" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.155-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.155-1.png", "caption": "Fig. 14.155 Robot for spraying fireproofing rock wool mass in small rooms", "texts": [ "154 consists of two main fire-resistant mass preparation units (Fig. 14.154a) and the fire-resistance mass spraying robot proper (Fig. 14.154b). Once the beam\u2019s elevation and length are keyed in the robot starts moving, sensing its position by means of an ultrasonic sensor and spraying fireproofing mass on the section\u2019s bottom and sides. The spraying nozzle can be raised to a height of 2.8\u20134.4 m by means of a screw elevator. Special software makes the entering of robot operation data simple and easy. The robot shown in Fig. 14.155 is intended for work in narrow rooms that are inaccessible to equipment operators. It consists of a carriage, an articulated arm with a spraying nozzle, and a control unit. The carriage has two sets of wheels: one for moving the robot longitudinally and the other for moving it transversely relative to the beam being sprayed with fireproofing mass. The articulated arm can rotate in both the horizontal and vertical plane and it can be moved to a distance of up 1500 m along the carriage\u2019s platform" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000942_cimsa.2004.1397255-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000942_cimsa.2004.1397255-Figure4-1.png", "caption": "Fig. 4. Case I Comparison a) Output b) Output Error c) F- Output", "texts": [ " Further the output angular position for the for the desired time span is plotted for traditional MRAC and proposed scheme along with the reference models output. Correspondingly, the angular position error with the output of the reference model and the variation in the fuzzy output (\u2018 w, \u2019) in case of the MFRMAC is also shown. Please note that on all these cases the \u2018 w, \u2019for the single reference model is kept as five. A. Case1 In the Case I the tip load of this manipulator is changed at different time instant as in table 2. Figure 4 shows the position trajectory plot and the output of the robotic arm with the proposed scheme and the single reference model adaptive controller in which W , is kept as \u20185\u2019. The mentioned figure also shows the output error comparison for both the methods and the change in \u2018U, \u2019 at every time instant. It can be seen that the single reference model Adaptive Controller is clearly unstable and was unable to control. B. Case 2 In the Case 2 the tip load of this manipulator is changed at different time instant as in table 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002959_20090909-4-jp-2010.00076-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002959_20090909-4-jp-2010.00076-Figure4-1.png", "caption": "Fig. 4. Different Ground Contact Conditions", "texts": [ "1, the ground contact conditions are changed according to gait phases. The different ground contact conditions result in different constraints of the human model. For example, the pivot point which means a fixed point in the model is changed with the ground contact conditions. It is the heel in IC, but is the hallux in PSw. From this pivot point, the position and the velocity of each segment are calculated. In swing phases, the hip joint is used as a pivot point, and its accelerations in x and y directions of GRS are measured. Fig. 4 shows the four ground contact conditions of the swinging leg, the supporting leg with whole foot, the supporting leg with forefoot, and the supporting leg with heel. These different constraints of the model introduce multiple sets of equations for estimation of the joint torques. Thus, the joint torques of S (\u03c4\u0302S), MS, TS (\u03c4\u0302MSt , \u03c4\u0302T St), PS (\u03c4\u0302PSw), and IC, LR (\u03c4\u0302IC, \u03c4\u0302LR) are calculated by the dynamics equations obtained by the models in Fig. 4 (a), (b), (c), and (d). Due to the multiple set of joint torque equations, there may be discontinuity in the calculated joint torques when the ground contact conditions are changed. Such discontinuities are smoothed by the gait phases detected by Smart Shoes as follows. The fuzzy membership values of each gait phase, \u00b5i, are set as the weighting vector, i.e., (1). W = [\u00b5IC \u00b5LR \u00b5MSt \u00b5T St \u00b5PSw \u00b5S] T \u2208 \u211c6\u00d71 (1) where \u00b5IC + \u00b5LR + \u00b5MSt + \u00b5T St + \u00b5PSw + \u00b5S = 1. The joint torques are calculated based on the human model with corresponding constraint as follows", " The virtual work is the sum of the virtual work done by conservative forces (Wc) and by non-conservative forces (Wnc), i.e., W = Wc +Wnc (5) The conservative work can be expressed as the negative of the potential energy of the system (V), i.e., \u03b4Wc = \u2212\u03b4V . Thus, \u03b4 (W +T ) = \u2212\u03b4V +\u03b4Wnc +\u03b4T = \u03b4L+\u03b4Wnc (6) where L=T-V is called the Lagrangian function. Then (4) can be rewritten as follows. d dt \u03b4L \u03b4 q\u0307 \u2212 \u03b4L \u03b4q = \u03b4Wnc \u03b4q \u2212 d dt \u03b4Wnc \u03b4 q\u0307 (7) For the details of Lagrangian mechanics, refer Curry (2005). In this model, the distance from GRS to the pivot point (X in Fig. 4) and joint angles of ankle, knee and hip (\u03b8 \u2019s in Fig. 4) are used as the generalized coordinates. Using (7), the equations for the joint torques of the lower extremity are derived. For the calculation of the joint torque, GCF measured by Smart Shoes, anthropometric data and motion information are required. To calculate the joint torques using the derived dynamics equations, anthropometric data such as length, mass and moment of inertia of each segment are required. There are huge anthropometric database about human body segments with different height, weight, and age (CDC (2008))" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002284_20080706-5-kr-1001.00797-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002284_20080706-5-kr-1001.00797-Figure5-1.png", "caption": "Fig. 5. Nonlinear guidance logic setup.", "texts": [ " The inner loop has roll and yaw-rate feedbacks; the yaw-rate is computed from the attitude angles as discussed in \u00a76.1 below. The lateral guidance scheme used is based on the work by Park et al. [2004]. The basic scheme is modified by introducing an adaptive tuning of the reference length to enhance the performance for large cross-track errors. Integral action is also added to improve tracking in the presence of constant disturbances. Here we will mainly discuss the proposed modifications in the basic scheme (Fig. 5). In the figure C is the vehicle\u2019s instantaneous position, D is a reference point selected on the desired path AB, L1 is the length of the line CD, V is the velocity of the vehicle, and \u03b7 is the angle between the velocity vector and CD. The roll command \u03c6c generated by the basic scheme is given as: \u03c6c = tan\u22121 [ 2V 2 gL1 sin { sin\u22121 ( y L1 ) + \u03c8E }] , (3) where y is the instantaneous cross-track deviation, \u03c8E is the heading error, and g is the acceleration due to gravity. Guidelines for selection of the length L1 are given by Park et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002616_jsen.2009.2031347-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002616_jsen.2009.2031347-Figure2-1.png", "caption": "Fig. 2. (A) Schematic illustration of MDEA 5037 2 cell -disc electrode array. (B) Photomicrograph of the microfabricated MDEA 5037 chip. (C) Single MDEA electrochemical cell with silverized reference electrode.", "texts": [ " The was then etch-patterned via fluoro-plasma to expose the 37 microdisc electrodes on each of the working electrodes, the large area counter electrode, and a smaller electrode slated to serve as a reference electrode [Fig. 1(b) and (g)]. The remaining metallic patterns of the transducer therefore remained passivated, insulated and isolated from the analyte and matrix. These dimensions satisfy the fundamental requirement that the counter electrode is able to support sufficient current into the contained medium without requiring an excessive cell voltage or creating a nonuniform current distribution at the working electrode. The die of the final implantable biotransducer is shown in Fig. 2. Fig. 2 shows a schematic illustration (A), a photomicrograph of the die (B) and a high magnification image of a MDEA5037 with a silverized reference electrode (C). Microdisc electrode arrays (MDEAs) are formed by arraying single microdiscs in various formats (hexagonal closed packed (HCP), square, etc.) [20]. Microdiscs are formed from a single contiguous layer of electrode material (Au) by patterning an insulator film deposited on the metallic electrode. The geometric area of the exposed electrode is therefore divided amongst an array of micron-dimensioned electrodes (in this case 50 ) each connected one to the other but separated from the electrolyte by a passivating layer, in this case of 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002492_iemdc.2009.5075175-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002492_iemdc.2009.5075175-Figure8-1.png", "caption": "Fig. 8 Approximate finite-element solution: 1 block conducting", "texts": [ " (It is not the only example of the possibility of a false result obtained from the finite-element method by the unwary user). The simple finite-element formulation satisfies (43) only over the whole solution domain, and not individually in each magnet region. An improved result can be obtained by suppressing the conductivity of, say, magnets B, C, and D, and simply calculating the loss in magnet A. In a situation where the losses must be equal in all magnet blocks, the final result is obtained by multiplying the result for one block by 8. When this is done in Fig. 8, the total calculated loss is 0@97 W when scaled up to include all magnets. This is slightly less than half the incorrect first estimate. The method can be described as a crude form of residual current suppression without using external circuits. This method of single-region eddy-current calculation obviously relies on the assumption that the eddy-currents in any magnet do not affect the eddy-currents in any other magnet. One could assume that this is characteristic of \u201cresistance-limited\u201d eddy-currents, but it is more correct to describe it as the neglect of proximity effect, which in other situations is known to be dangerous" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001028_bfb0042540-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001028_bfb0042540-Figure2-1.png", "caption": "Figure 2: Denavit-Hartenberg coordinates, and tip vector b~.", "texts": [ " The combined variation in all the parameters is presumed to cause this endpoint variation. Specifically~ a variation of the contribution to the end effector D-H parameter s t along the local link z ~xis, zj_1, causes a linear velocity of Asjz~. I\" The parameter variation A% about the local link z axis, x~, causes a contribution to the endpoint's angular velocity of (Aa~)x~. = w~j, and a linear b * velocity contribution of w~j x bj+1, where j+x is a vector from the jtl~ coordinate system to the endpoint (Figure 2). The 0j and aj parameters are treated analogously. In total, the endpoint translation due to all of the parameter variations is given by: tt z i b~A0j + z~ aAs t + x$ x b$+lAa t + j=l and angular variation given by: r t . . E Z;-l~k0J \"~- K;Ao~j (15) $=1 Comparing these to (11) it is seen that the columns of each of the four Jacobians are and zj_l. x bj co15 - - , colt = = . j+ l (17) co15 , cols = The Jacobian columns for parameters of the alternate Hayati convention are found analogously to be: and Z I i ] a~ [ xxj_~ ] (18) , colj ~ = 0 [ , , ] xx~ x bj+~ (19) colj ~ ---- xx~ __ y~ \u2022 , COlj ~ = y j X hi+ 1 where xxj stands for the local x-axis just prior to the last rotation about the f lh y-axls by % (see Figure 3)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000462_978-1-4020-2249-4_38-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000462_978-1-4020-2249-4_38-Figure6-1.png", "caption": "Figure 6 Flat Ring Articulation for the Octahedral Zig-Zag Linkage", "texts": [ " = 0 (left side) and in a medium extension If! = Ji / 6 (right side). Figure4 Parallel Plate Articulations for the Hexahedral Zig-Zag Linkage Figure5 Hexahedral Zig-Zag Linkage with Parallel Plate Articulations The third kind of articulation of the Nuremberg scissors will be called the \"Flat Ring Articulation\" which will be demonstrated at the octa- hedron. It consists of as many gussets as there are edges meeting at the polyhedral apex (four gussets in the octahedron case). The gussets sur round the apex in the form of a ring (Fig.6). The orientation of the gus sets is independent of the position angle cpo The parameter set {a,s,d,t} together with f3 = 7r / 4 determines the length of the octahedron edge: L( cp) = 2[t + .fi(a coscp - ssincp)]+ 6(a sin cp + scoscp). Unlike the mechanisms in Fig. 2 and Fig.4 the mechanism shown in Fig.6 is not overconstrained but mobile with six degrees of freedom. This cannot however, be proved by carelessly applying the structure formula F = I./; - 61 because this mechanism is not purely spatial, as it contains planar part mechanisms. Nevertheless F = I./; - 61 can be applied if the planar Nuremberg scissors are taken as \"nonrigid bodies\" and the num ber of fundamental loops is determined by disregarding the loops of the planar part mechanisms and the sum of the degrees of freedom due to the defomations of the nonrigid bodies is finally added" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002971_j.mechmachtheory.2009.11.001-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002971_j.mechmachtheory.2009.11.001-Figure1-1.png", "caption": "Fig. 1. Schematic illustration of the gear-bearing system under nonlinear suspension.", "texts": [ " Section 2 derives dynamic models for the gear-bearing system with a nonlinear suspension effect, strongly nonlinear gear mesh force and strongly nonlinear oil-film force. Section 3 describes the techniques used in this study to analyze the dynamic response of the gear-bearing system. Section 4 presents the numerical analysis results obtained for the behavior of the gear-bearing system under various operational conditions. Finally, Section 5 presents some brief conclusions. The dynamic model to stimulate the gear-bearing system under the assumptions of nonlinear suspension effect and strongly nonlinear fluid film force effect is established in Fig. 1. Fig. 2 presents a schematic illustration of the dynamic model considered between gear and pinion. Applying the principles of force equilibrium, the forces acting at the center of journal 1, i.e. Oj1 (Xj1, Yj1) and center of journal 2, i.e. Oj2 (Xj2, Yj2) are given by Fx1 \u00bc fe1 cos u1 \u00fe fu1 sin u1 \u00bc Kp1\u00f0Xp Xj1\u00de=2; \u00f01\u00de Fy1 \u00bc fe1 sin u1 fu1 cos u1 \u00bc Kp1\u00f0Yp Yj1\u00de=2; \u00f02\u00de Fx2 \u00bc fe2 cos u2 \u00fe fu2 sin u2 \u00bc Kp2\u00f0Xg Xj2\u00de=2; \u00f03\u00de Fy2 \u00bc fe2 sin u2 fu2 cos u2 \u00bc Kp2\u00f0Yg Yj2\u00de=2; \u00f04\u00de in which fe1 and fu1 are the viscous damping forces in the radial and tangential directions for the center of journal 1, respectively, and fe2 and fu2 are the viscous damping forces in the radial and tangential directions for the center of journal 2, respectively", " (1), (2), (3), (4), (15), (16), (17), (18), (19), (20), (21), and (22) can be expressed as e01 \u00bc b1\u00bd\u00f0xp x1\u00de cos u1 \u00fe \u00f0yp y1\u00de sinu1 e1 \u00f01 e2 1\u00de 3=2 ; \u00f029\u00de u02 \u00bc 1 2e2 e2 b2\u00bd\u00f0xg x2\u00de sinu2 \u00f0yg y2\u00de cos u2 \u00f02\u00fe e2 2\u00de ffiffiffiffiffiffiffiffiffiffiffiffiffi 1 e2 2 q ; \u00f030\u00de u01 \u00bc 1 2e1 e1 b1\u00bd\u00f0xp x1\u00de sinu1 \u00f0yp y1\u00de cos u1 \u00f02\u00fe e2 1\u00de ffiffiffiffiffiffiffiffiffiffiffiffiffi 1 e2 1 q ; \u00f031\u00de e02 \u00bc b2\u00bd\u00f0xg x2\u00de cos u2 \u00fe \u00f0yg y2\u00de sin u2 e2 \u00f01 e2 2\u00de 3=2 ; \u00f032\u00de x00p \u00bc 2n2 s x0p 1 s2 \u00f0xp x1 e1 cos u1\u00de \u00fe b cos\u00f0/=4\u00de 2n3 s \u00f0x0p x0g Ep sin /\u00de K s2 \u00f0xp xg Ep cos /\u00de; \u00f033\u00de y00p \u00bc 2n2 s y0p 1 s2 \u00f0yp y1 e1 sinu1\u00de \u00fe b sin\u00f0/=4\u00de 2n3 s \u00f0y0p y0g Ep cos /\u00de K s2 \u00f0yp yg Ep sin /\u00de f s2 ; \u00f034\u00de x00g \u00bc 2n4 s x0g 1 s2 \u00f0xg x2 e2 cos u2\u00de \u00fe bg cos\u00f0/=8\u00de \u00fe 2n5 s \u00f0x0p x0g Ep sin /\u00de Kg s2 \u00f0xp xg Ep cos /\u00de; \u00f035\u00de y00g \u00bc 2n4 s y0g 1 s2 \u00f0yg y2 e2 sinu2\u00de \u00fe bg sin\u00f0/=8\u00de \u00fe 2n5 s \u00f0y0p y0g Ep cos /\u00de Kg s2 \u00f0yp yg Ep sin /\u00de fg s2 ; \u00f036\u00de x001 \u00fe 2n1 s1 x01 \u00fe 1 s2 1 x1 \u00fe a1 s2 x3 1 1 2C1ps2 \u00f0xp x1 e1 cos u1\u00de \u00bc 0; \u00f037\u00de y001 \u00fe 2n1 s1 y01 \u00fe 1 s2 1 y1 \u00fe a1 s2 y3 1 1 2C1ps2 \u00f0yp y1 e1 sinu1\u00de \u00fe f s2 \u00bc 0; \u00f038\u00de x002 \u00fe 2n6 s2 x02 \u00fe 1 s2 2 x2 \u00fe a2 s2 x3 2 1 2C2ps2 \u00f0xg x2 e2 cos u2\u00de \u00bc 0; \u00f039\u00de y002 \u00fe 2n6 s2 y02 \u00fe 1 s2 2 y2 \u00fe a2 s2 y3 2 1 2C2ps2 \u00f0yg y2 e2 sinu2\u00de \u00fe f s2 \u00bc 0; \u00f040\u00de Eqs. (29)\u2013(40) describe a nonlinear dynamic system. In the current study, the approximate solutions of these coupled nonlinear differential equations are obtained using the fourth order Runge\u2013Kutta numerical scheme. The nonlinear dynamics of the gear-bearing system shown in Fig. 1 are analyzed using Poincar\u00e9 maps, bifurcation diagrams and the Lyapunov exponent in this study. The basic principles of each analytical method are reviewed in the following sub-sections [17\u201318]. The dynamic trajectories of the gear-bearing system provide a basic indication as to whether the system behavior is periodic or non-periodic. However, they are unable to identify the onset of chaotic motion. Accordingly, some other form of analytical method is required. In the current study, the dynamics of the gear-bearing system are analyzed using Poincar\u00e9 maps derived from the Poincar\u00e9 section of the gear system" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000718_1.1829068-Figure13-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000718_1.1829068-Figure13-1.png", "caption": "Fig. 13 Kinematic similarity of the wobble gear and harmonic drive \u201eS1 \u2026", "texts": [ " This is the same input-output relation as the wobble gear, as pointed out near the beginning of this paper, so the wobble gear and the sun-ring-planet train are similar mechanisms, and Eq. ~4! represents the speed ratio of each. Eliminating link 1 of the sun-ring-planet train gives the more intuitive similarity relation illustrated in Fig. 12. The planar trains in Figs. 11 and 12 differ only in numbers of teeth and are topologically the same. It follows that the wobble gear is also related to the harmonic drive through their respective reduction ratio formulas, giving a similarity relation as shown in Fig. 13. Once again labeling the input and output as links 3 and 4, respectively, the two formulas for v4 /v3 are identical ~see Eq. ~3!!. Here link 2 functions as a planet link. For the wobble gear mechanism, the input is analogous to a wave generator in that it initiates contact between the two meshing links. Then instead of flexing, contact is achieved by a wobbling motion. So the wobble gear mechanism acts similarly to the harmonic drive in doing away with the problem of manufacturing a tiny carrier link, except MARCH 2005, Vol", " Link 1 can be dropped, since as an \u2018\u2018extra\u2019\u2019 link it can have arbitrary motion. This modified Humpage train then mimics the behavior of the wobble gear if N55N28 ~causing speed ratios for link 4 to match! and if the length of link 3 approaches zero in the limit ~causing the velocity vectors for link 2 to align!. The effect of these constraints is to force the planet to nutate about a point on the input/output axis, with no net angular displacement about this axis; in the wobble gear itself, this function is performed by the cam tabs protruding from link 2 as shown in Fig. 13. From Table 2 representing this constrained case of the Humpage set with links 3 and 4 being the input and output, respectively, the reduction ratio simplifies to v4 /v3512N2 /N4 (5) Recall that the wobble-gear speed relationship has been given in Eq. ~4!. This is identical to the analogous speed ratio given by the special case of the Humpage set in Eq. ~5!. In summary, for the physical construction of the train satisfying the tooth-number constraint mentioned earlier, the similarity index improves to S1, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000586_j.ijmachtools.2005.01.003-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000586_j.ijmachtools.2005.01.003-Figure10-1.png", "caption": "Fig. 10. Results of centered DBB test with setup eccentricity of (0.5, 0.5) m", "texts": [ " When the workpiece ball setup eccentricity is on the axis from which the trajectory starts and finishes, the problem does not appear. However, when the distance of workpiece ball is increased from that axis the problem appears as a radial gap between the beginning and the end of the trajectory. The poor overlap may be due to a rapid change in resulting volumetric errors in that region or a setup eccentricity. The missed or overlapped trajectory artificially increases or decreases the amount of eccentricity and ovalization. Some index adjustment was done in order to reduce this effect. In Fig. 10, a comparison between the coordinate transformation and the proposed method is shown using index adjustment to remove the aforementioned gap. As can be seen the coordinate transformation method causes false monitoring of the machine errors, whereas the proposed method is in a good agreement with the non eccentric trajectory. Setup eccentricity is an integral part of any DBB test. It was demonstrated that a simple numerical centering of the data using a coordinate transformation contaminates the data with an unwanted and well-known fictitious ovalization effect" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003251_s1068371211020088-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003251_s1068371211020088-Figure1-1.png", "caption": "Fig. 1. Function F Q\u0302( ).", "texts": [ " x i 1+( ) f x i( ) u i( ) w i( ), ,( ),= x\u0302\u2013 x\u0302 x\u0302\u2013 (x\u0302 i( ) u i( ) 0 ),, , x\u0302 i 1+( ) x\u0302\u2013 i 1+( ) K i 1+( )+= \u00d7 (z i 1+( ) h x\u0302\u2013 i 1+( ) 0,( )( )\u2013 T\u0302L x\u03026 x\u03025 2 72 RUSSIAN ELECTRICAL ENGINEERING Vol. 82 No. 2 2011 KISELICHNIK, BODSON If the efficiency of the pump is taken to be con stant, then after substituting the estimates instead of the real values, the system (7) is transformed into (16) where kn = \u03c1g/\u03b7; kp = H01 / (i + 1) = (i + 1) (i + 1) is the estimation of consumed power at the pump shaft; and is the desired flow rate estimate. The rate is selected as a positive root of Eq. (16). Let us consider the case of a non null flow rate. Then, the function F( ) has two extremum points (Fig. 1) and Eq. (16) has three roots, and One of the roots is negative and, therefore, it is not accepted. Follow up studies have shown that it is necessary to select the first positive root. It is possible for the first and second roots to coincide; however, there is always at least one positive or null solution because, if the pump operates normally, there is a positive or null (due to the check valve) flow rate. The right hand line 1 in Fig. 1 is a tangent to the point (0, (i + 1)/(knap)). The right hand line joins two points (0, (i + 1)/(knap)) and F = (i + 1) can easily be found iteratively in the range (17) The estimation of the pump head (i + 1) is obtained by substituting the corresponding rates of the velocity and performance into the equation of the head (7) instead of the real values. F Q\u0302( ) Q\u0302 3 kpx\u03025 2 i 1+( ) ap Q\u0302\u2013 P\u0302p i 1+( ) knap + 0,= = ig 2 \u03c9r 2 ; P\u0302p T\u0302L x\u03025 Q\u0302 Q\u0302 Q\u0302 Q\u03021, Q\u03022 Q\u03023. Q\u03023 P\u0302p P\u0302p (Q\u0302e1, Q\u0302e1( )). Q\u03021 Q\u0302 Q\u03021b P\u0302p i 1+( ) knkpx\u03025 2 i 1+( ) Q\u03021 Q\u03021f\u2264 \u2264 3P\u0302p i 1+( ) 2knkpx\u03025 2 i 1+( ) " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003987_j.foodchem.2012.11.076-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003987_j.foodchem.2012.11.076-Figure2-1.png", "caption": "Fig. 2. Modified Job\u2019s method for the study of stoichiometry between manganese(III) sulphate and diphenylamine: 1. Series 1 involving the solutions amounted to 5 ml and diluted to 10 ml with water. 2. Series 2 the solutions involving distribution same as that of series1 but diluted to 25 ml with water.", "texts": [ " The composition between manganese(III) sulphate and diphenylamine was determined by modified Jobs method of continuous variation (Shyla, VijayaBhaskar, Sathisha, & Nagendrappa, 2010) using equimolar solutions, 10 4 M of manganese(III) sulphate and Series A = 3 ml 1.92 10 4 M Mn3+ or 0.6 ml MnO 4 or 3 ml Ce4+ or 0.5 ml Cr2O2 7 or 3 ml Fe3+ + 0.5 ml 5 M sulphuric acid + 1.5 ml of 10 4 M diphenylamine, diluted to 10 ml with water. Series B = 3 ml 1.92 10 4 M Mn3+ or 0.6 ml MnO 4 or 3 ml Ce4+ or 0.5 ml Cr2O2 7 or 3 ml Fe3+, without diphenylamine diluted to 10 ml with water. diphenylamine and it was found to be 4:1 in the respective order, (Fig. 2) and the chromophore obtained may be N,N0-diphenyl-pdiphenoquinondiimine (Pankratov, 2001). The experiment for determining stoichiometry between manganese(III) sulphate and barium diphenylamine suphonate was conducted with equimolar, 2 10 4 M solutions of manganese(III) sulphate and barium diphenylamine sulphonate. A series of eleven flasks were arranged. To each flask 0.5, 1.0 1.5, 2.0, 2.5, 3.0, 3.5, 4.0, 4.5, 4.7 or 4.9 ml 2 10 4 M manganese(III) sulphate and 4.5, 4.0, 3.5, 3.0, 2.5, 2.0, 1.5, 1", " The results of ascorbic acid (mg/100 g) obtained from the fifteen kinds of fruits are indicating that amla (380) is the best sources of ascorbic acid, followed by guava (202), lemon (174), avocado (162) are good sources of ascorbic acid, orange (82), lime (69), grapes (55), chikoo (51), papaya (46) are the moderate sources of ascorbic acid whereas banana (27), watermelon (24), tomato (22) and apple (15) are the poor sources of ascorbic acid. But such a kind of classification cannot be made for fruit juices since the results obtained from the different kinds of juices are found to be containing less amount of ascorbic acid and there is no considerable variation of ascorbic acid content from one kind of fruit juice to another. The new method is based on established chemical principles like stoichiometry Fig. 2, reaction scheme and the nature of manganese(III) over other oxidants, Table 1. The method is considered to be having a good selectivity for the determination of ascorbic acid present in the samples of various fruits, commercial juices and sprouted food grains since some common ascorbic acid associated agents are not interfering in the determination of ascorbic acid, however, the interference of sulphurdioxide over and above 2.68 lg ml 1 could be prevented by its effective masking with formaldehyde, Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000402_icpp.2005.9-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000402_icpp.2005.9-Figure6-1.png", "caption": "Figure 6. (a) Section identification. (b) Edge Identification. (c) Edge construction. (d) Boundary construction.", "texts": [ " In phase three, it will route around that neighboring section to reach its corresponding start corner. Once this message reaches that corner, those same type corners in neighboring sections are identified as preceding and succeeding edge nodes and the path between them will be used for future edge construction. For example, the (+Y \u2212 X)-edge of an MCC is defined by the (+Y \u2212 X)-corners of all XY sections of this MCC. In phase one, from a (+Y \u2212 X)-corner c(xc, yc, zc) in the XY section z = zc in Figure 6 (a), a message will be sent to route around this section. When such a message passes through a node u(xu, yu, zc) with an unsafe neighbor in the \u2212Y dimension, the identified information of the Y Z section on the plane x = xu is used to find a neighboring section on plane z = zc + 1. In phase two, the message will go around the corresponding Y Z section to the neighboring XY section (seen in Figure 6 (a)). In phase three, once the message arrives at a node of the neighboring XY section, a twohead-on message propagation will be initiated to go around that section (one clockwise and one counter-clockwise) to reach its corresponding (+Y \u2212 X)-corner u\u2032 (seen in Figure 6 (b)). At node u\u2032, c is identified as its succeeding node along the edge and the information of the path to node c (see in Figure 6 (c)) is saved for future information propagation. Edge construction. If the neighboring section cannot be found in phase one in the above identification process, that start corner is identified as one end of this edge and the cor- responding section is identified as the surface of this MCC (see corner c\u2032 in Figure 6 (a)). In phase three of the above identification process, a concave region of the MCC containing the start corner can be identified if there is another 2-D section in the same plane. For example, in Figure 6 (b), at node u\u2032\u2032, the second section is found. Once such a concave region is found, the start corner (corner c\u2032\u2032 in the example in Figure 6 (b)) will be identified as one end of another edge (an inner edge of this MCC towards a concave area). Starting from an identified end node u , the entire edge will form by collecting all the links between preceding and succeeding nodes along this edge. From u, a message is sent along the path to its succeeding node and such propagation will continue until it reaches the other end (an edge node without any succeeding node). At each edge node v it passes through, the section information is collected and the previously saved information is used to form a part of this MCC, M(v). This M(v) includes the MCC area from the current node v to the end node u. With the information of M(v), the information of forbidden region Q(v) and the critical region Q\u2032(v) can also be formed at node v (see in the example in Figure 6 (c)). Boundary construction. At each edge node u, say along a (+Y \u2212 X)-edge, after M(u) is formed, the information of M(u), QY (u), and Q\u2032 Y (u) will be propagated along the boundary (also called (+Y \u2212 X)-boundary) to block the routing from entering the detour area QY (u) in the +X dimension if the destination is inside the critical region Q\u2032 Y (u). Initially, a message carrying the information is 6 Proceedings of the 2005 International Conference on Parallel Processing (ICPP\u201905) 0190-3918/05 $20", " Meanwhile, the forbidden region of node v (QY (v)) will be merged into QY (u) (QY (u) = QY (v) \u222a QY (u)). If node u is the (\u2212Z)-most edge node (the edge node without any succeeding node), at node u\u2032, a copy of the message will be sent to node v. From node v, it will go along that (+Y \u2212 X)-edge and reach the (\u2212Z)most edge node of the latter MCC. At each edge node v\u2032 it passes through, a boundary is constructed for node u with the information of M(u), Q\u2032 Y (u), and QY (u) \u222a QY (v\u2032). A sample of boundary construction is shown in Figure 6 (d). The whole procedure is shown in Algorithm 5. Algorithm 5: Identifi cation and boundary construction of an MCC in 3-D meshes 1. Identifi cation of each 2-D section (see in Algorithm 2). 2. Identifi cation of each edge: (a) a message is sent along each XY , Y Z, or XZ section from its start corner to fi nd the neighboring section; (b) the message reaches the neighboring section in one hop; (c) an identifi cation process in algorithm 2 is applied to reach the corresponding corner in this neighboring section, the preceding node of that start corner" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003032_j.wear.2009.01.047-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003032_j.wear.2009.01.047-Figure1-1.png", "caption": "Fig. 1. Schema", "texts": [ " The high frequency reciprocatng tribometer measures friction coefficient in both the forward and ackward movements and is typically used to measure the lubricity f fuels in IC engines which simulates high frequency reciprocating otion [6]. In order to simulate the low speed reciprocating slidng behaviour encountered in most of the real life applications in ndustries such as machine tool slide ways, the above-mentioned ribometers are not suitable. Thus there is a need for a system that imulates the actual conditions over a wide range of speed. The urrent investigation addresses the development of an improved ribometer for such an evaluation. . Experimental .1. Reciprocating tribometer A schematic of the tribometer is shown in Fig. 1. The test rig is riven by a motor coupled with a gearbox and interfaced with a PC. he motor drives the slide way with the help of a timing belt so as o give a positive drive. When the slide way reaches the extreme nds, the proximity switch reverses the direction of rotation of the otor, thus moving the slide way in the opposite direction. The lide way thus continuously reciprocates between the two sensors. he tribometer. The limit switches provided at the extreme ends act as a safety device and shuts down the machine in case the slide way overshoots the set limits" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002146_2007-01-1251-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002146_2007-01-1251-Figure3-1.png", "caption": "Figure 3. Coordinate system and notation.", "texts": [], "surrounding_texts": [ "At each crankangle, the piston position within the cylinder bore and the skirt deformation specify the oil film thickness distribution h used in Eq. (3). The instantaneous oil film squeeze term is denoted by h . The hydrodynamic pressure is primarily a function of the local film thickness and the film squeeze action; i.e. ),( hhPP (5) An inverse solution of the Reynolds equation is required, because the position of the piston within the cylinder bore is not known a-priori. Eq. (3) is solved iteratively at each crank angle. The skirt clearance profile is calculated using this piston eccentricity which is defined by the piston secondary translation and rotation. For a very small time step, or crank angle increment, any change in the oil film clearance field h and its time derivative h induces a change in the hydrodynamic pressure distribution. Considering the following small perturbation of h and h o o hh hh , (6) a first-order approximation for the hydrodynamic pressure around the equilibrium position oP can be used as follows, h P h PPP o (7) 5 Note that oh , oh and oP satisfy Eq. (3). The terms h PK and h PC constitute the oil film dynamic coefficients and comprise the stiffness and damping matrices respectively. Substitution of Eqs (6) and (7) in Eq. (3) results in a perturbed Reynolds\u2019 equation which provides the following three differential equations for the hydrodynamic pressure, stiffness and damping, respectively ohz oh ZV oPyhxoPzhz 126 33 (8) zZVK y h x K z h z 633 oPzohzoPzohz 2323 (9) 1233 C y h x C z h z . (10) A detailed derivation of these equations can be found in [15, 19, 20]. Because there is no closed form solution, a numerical method is utilized. In this work, we use a computationally efficient finite-difference methodology which is described in the next section. Ref [20] explains the computational advantage of this method in relation with the commonly used finite-element approach." ] }, { "image_filename": "designv11_20_0002449_j.mechatronics.2009.06.012-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002449_j.mechatronics.2009.06.012-Figure5-1.png", "caption": "Fig. 5. Spindle bearings with actuator.", "texts": [ " Besides, this section identifies the characteristic of Foucault Current and its application on the spindle growth measurement. 2.1. Bearings loading control Fig. 4 shows a motorized high speed spindle using spring force to preload the spindle bearings. Most high speed spindles are designed for high rotational speed in order to RPM (rotations per minute), and high simultaneous axes feeds for light cutting tool load applications [7,8]. A piezoelectric actuator has been placed at the rear of the spindle to maintain proper bearing loading shown in Fig. 5, but the front spindle bearings continue to be of the traditional design and configuration which ultimately still produces considerable heat with the increase of spindle RPM and spindle running time. X 0g bX X i oiD Spacer Tb Shaft iT \u00a3c t directions. Right: preload mechanism. Fig. 6 outlines the thermal growth of a motorized high speed spindle running at different RPM for a constant time period. All speed curves are nonlinear and the differences between all curves are not parallel. Each thermal raise curves are not identical and nonlinear" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001878_j.ijsolstr.2008.06.008-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001878_j.ijsolstr.2008.06.008-Figure1-1.png", "caption": "Fig. 1. Symbols and notation.", "texts": [ " Homogeneous solutions refer to situations in which the membrane may inflate freely, while inhomogeneous solutions consider practical situations when a large variety of constraints are imposed on the equilibrium surface. When considering the homogeneous case of axisymmetric membranes, the \u2018\u2018pseudo-deformed surfaces\u201d first introduced by Wu (1974b), using different means, are naturally recovered. Herein, the membrane in question is modeled as an ideal, perfectly flexible, two-dimensional body made of an inextensible material. In the initial configuration C0, assumed as reference, the membrane is unloaded and stress-free, and coincides with the boundary C0 of the closed region X0 2 R3 (Fig. 1). When a uniform pressure p acts internally, the membrane reaches the deformed configuration C, coinciding with the unknown boundary C of the closed region X 2 R3. Over the years, a number of equilibrium problems regarding partly or fully wrinkled elastic membranes have been solved either by means of a kinematical approach (Wu, 1974a; Liu et al., 2001; Roddeman et al., 1987a,b; Barsotti et al., 2001) or, more frequently, following a physical approach (Reissner, 1938; Stein and Hedgepeth, 1961; Pipkin, 1986; Steigmann, 1990; Haseganu and Steigmann, 1994)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001927_wst.2008.558-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001927_wst.2008.558-Figure1-1.png", "caption": "Figure 1 | Exploded view of the FM01-LC laboratory cell electrolyzer in the undivided mode (Griffiths et al. 2005).", "texts": [ " A potentiostat-galvanostat EG&G model PAR 273 and M270 software were used for all electrochemical experiments. Chemical oxygen demand (COD) analyses were performed using a dry-bath Lab Line Model 2008, and a spectrophotometer Genesys 20. The cell potential was determined through an AgilentTM high impedance multimeter. As previously mentioned, the FM01-LC reactor has been described in detail in the literature (Brown et al. 1992; Griffiths et al. 2005; Butro\u0301n et al. 2007; Nava et al. 2007). An exploded view of the cell that includes the turbulence promoter used within the cell channel is shown in Figure 1. In this work the spacer was 0.6 cm thick. The electrodes were 3D BDD and a platinum coated titanium flat sheet, respectively. Details on the FM01-LC characteristics are given in Table 1. BDD electrodes were provided by MetakemTM, with thickness of 2\u20137mm supported on Ti expanded metal. The undivided mode with a single electrolyte compartment and the electrolyte flow circuit for the FM01-LC cell is shown in Figure 2. The electrolyte was contained in a 1L polycarbonate reservoir. A magnetic coupled pump of 1/15 hp March MFG, model MDX-MT-3, with a flow rate capacity up to 300 cm3 s21 was used; the flow rates were measured by a variable area glass rotameter from Cole Palmer, model F44500" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003066_icinfa.2010.5512065-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003066_icinfa.2010.5512065-Figure2-1.png", "caption": "Fig. 2 Physical photo of 3-RPS parallel robot", "texts": [ " 1, the upper and lower platforms are two equilateral triangles with different lengths, which were connected with three retractable supporting poles. The connection between three supporting poles and the upper platform are three ball hinges, while the connection of the lower platform is three revolute pairs. according to type of analysis, in the movement of 3-RPS parallel robot, the upper platform can achieve six-dimension movement relative to lower platform, but only three of them are controllable. 3-RPS parallel robot physical photo is shown as Fig. 2. III. POSTURE INVERSE KINEMATICS OF 3-RPS There are three controllable DOF in 3-RPS parallel robot, which can control x, y, z axises(another three DOF were uncontrollable), or control the angle around the x-axis, the angle around the y-axis and z too. They are respectively called as Position Kinematics and Posture Kinematics. In posture inverse kinematics, the posture of the upper platform is known,which includes the angle around the x-axis, the angle around the y-axis and the displacement along the zaxis relative to the fixed platform, and then we solve the length 1s , 2s and 3s of every pole" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001111_soli.2006.328881-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001111_soli.2006.328881-Figure2-1.png", "caption": "Figure 2. Twin Rotor seen from the above", "texts": [ " azimuth angle in horizontal plane and elevation angle in vertical plane are influenced by the rotating propellers simultaneously. The helicopter body is equipped with the third motor which is used for changing the center ofthe gravity ofthe body. The helicopter model is a multivariable dynamical system with up to three manipulated inputs and two measured outputs. The system is essentially nonlinear and the mathematical model can be linearized around the set point. 3.1 Helicopter body dynamics in the horizontal plane (Azimuth subsystem) Figure 2 shows a sketch of the helicopter body seen from the above. f(x,u) By balancing the torques acting on the body in the horizontal plane, as shown in figure2: where, St: Azimuth angle (Rad) , I: Moment of Inertia (kg mi), T : Friction torque(Nm), tr: Main rotor reaction torque (Nm), Z2 Torque generated by the side rotor (Nm). 3.2 Helicopter body dynamics in the vertical plane (Elevator subsystem) Figure 3 shows a sketch of the helicopter body seen from the side. By balancing the torques acting on the body in the vertical plane, as shown in figure3: ITW = l +,Gc+ Gf m (8) VI: Elevation angle (Rad), TG: Gyroscopic torque (Nm), 17: Friction torque(Nm), Im Moment of Inertia (kg m), Z': Lift torque generated by the main rotor (Nm), TC Centrifugal torque (Nm) , ZTn: Torque generated by the mass of the body (Nm) Now we have a model of the helicopter dynamics in both vertical and horizontal planes [9]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001730_ac60310a059-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001730_ac60310a059-Figure1-1.png", "caption": "Figure 1. Ultraviolet absorbance scans of 0.1% NTA in water adjusted to the following pH values: (1) 8.4; (2) 8.8; (3) 9.2; (4)10.2", "texts": [ " To determine if closely related chelating agents interfered with the detection of NTA, the chelating agents listed in Table I were chromatographed., The possibility of metal interferences was checked by testing a series of solutions containing 0.001M NTA in 0.1M metallic salt. Similar mixtures of 0.001M NTA in 0.1M NaOH and 0.1M \u201c 0 8 were also tested. The separating column was 1 meter long by (11) J. J, Kirkland, J . Chromatogr. Sci., 8, 72 (1970). 418 ANALYTICAL CHEMISTRY, VOL. 44, NO. 2, FEBRUARY 1972 Monitoring the column effluent only at 254 nm was found to be somewhat restrictive due to the UV absorbance characteristics of NTA. Figure 1 shows ultraviolet absorbance curves for buffered solutions at various pH values. As can be seen, at 254 nm the absorbance does not become significant unless the environmental pH is quite high. The degree of ionization of NTA is dependent upon pH and ionic strength. The relationship in Figure 1 reflects the formation of the NTA3-ion, which uniquely absorbs UV light at 254 nm. If the pH of the buffer is raised to pH 10 or above, the coating of quaternary ammonia substituted methacrylate polymer will hydrolyze rapidly and serious damage to the column can occur (12). Fortunately, it was possible to find a buffer system at pH 9.0 that permitted the separation of NTA from the other chelates, and still gave sufficient NTA response to be useful. The response of the system to NTA was linear through the range of 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001879_tt.51-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001879_tt.51-Figure2-1.png", "caption": "Figure 2. Two identically sized hollow rollers in contact (M 1-60 Two)", "texts": [ " The behaviour of the main roller has been investigated when it is in contact with another identically sized roller, as is the case in Model 1, and when it is in contact with the inner race or the outer race, as is the case in the non-identically sized rollers model (Model 2). A consistent convention is used for naming the models based on whether the rollers are of the same size or not, the percentage of hollowness, and whether one roller is hollow or both rollers are hollow. The name of each model starts with M, which refers to Model. As mentioned earlier, identically sized models are called Model 1, while non-identically sized roller models are called Model 2. Figure 2 shows a model of two identically sized rollers, with 60% of hollowness. Thus, it is called M 1\u201360. Then the name is followed by the word (One) or (Two) to refer to the number of hollow rollers in the model. In the case of a solid model, there is no need to use any of these two words. The letter T was added to the name of the models to indicate that those models are subjected to a combined normal and tangential loading. A new model for fatigue life prediction of roller bearings was suggested by Ioannides and Harris15 in 1985" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002773_12.878749-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002773_12.878749-Figure1-1.png", "caption": "Figure 1. LMD station with pyrometry temperature monitoring system.", "texts": [ "es XVIII International Symposium on Gas Flow, Chemical Lasers, and High-Power Lasers, edited by Tanja Dreischuh, Petar A. Atanasov, Nikola V. Sabotinov, Proc. of SPIE Vol. 7751, 775123 \u00b7 \u00a9 2010 SPIE \u00b7 CCC code: 0277-786X/10/$18 \u00b7 doi: 10.1117/12.878749 Proc. of SPIE Vol. 7751 775123-1 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 07/17/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx 2. EXPERIMENTAL PROCEDURE The experiments were carried out with the setup shown in Fig. 1. LMD process has been performed by means of a Nd:YAG laser (wavelength 1064 nm) TRUMPH model HL1006D, with maximum output power of 1kW at the work piece. The beam is conveyed to a lens assembly of focal length 200 mm within a coaxial deposition head via an optical fibre with internal diameter 0.6 mm. This focused the beam to a circular spot with diameter 1,95 mm. and nominally Gaussian beam profile at the deposition point, which was positioned 12 mm beyond the end of the nozzle. The powder is delivered by a Sulzer Metco Twin 10C accurate powder feeder" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003327_s11249-010-9602-8-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003327_s11249-010-9602-8-Figure9-1.png", "caption": "Fig. 9 The schematic diagram of the finite element model", "texts": [ " The shape of the entrapped grease abode by the cosine function, and the entrapped grease was just at the center of the contact area. 3. The contact stress that can deform the disk and ball in the contact area is close to the pressure in the entrapped grease. In this way, the shape of the entrapped grease can be expressed by: h \u00bc h0 cos h \u00bc h0 cos px 2r0 \u00f05\u00de where h is the entrapped grease film thickness, h0 is the maximum entrapped grease film thickness (h0 = 200 nm), x is the distance from one point to the center point of the contact area, and r0 is the radius of the entrapped grease (r0 = 48 lm). Figure 9 shows the schematic diagram of finite element model. The pressure distribution in the contact area is shown in Fig. 10. It can be seen that the pressure increased rapidly in the entrapped grease region as compared with other regions in the Hertzian contact area. The mean pressure of the entrapped grease region is about 0.85 GPa in the sliding tests. As shown in Fig. 11, the critical yield shear stress of the grease at atmospheric pressure is 244 Pa. According to Eq. 4, the critical yield shear stress of the entrapped grease is 32" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001822_1.2825392-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001822_1.2825392-Figure8-1.png", "caption": "Fig. 8. Geometry of the collision between a cue and ball. The deflection of the cue is exaggerated in this drawing, at least when the cue is supported by hand near the impact point. Without this support, the cue would deflect approximately as shown. The positive directions of and are indicated.", "texts": [ " 7 arriving at the tip after the collision with the ball. Similar behavior of the tip was observed at all impact points on the ball and on the glass block provided there was no miscue. As described in Sec. 208 Am. J. Phys., Vol. 76, No. 3, March 2008 Downloaded 30 Sep 2012 to 136.159.235.223. Redistribution subject to AAPT VI, the sudden change in the sign of the acceleration can be attributed to a release of the elastic energy stored in the cue tip. The impact of a cue stick with an initially stationary billiard ball is shown in Fig. 8. The cue stick is modeled as a uniform, rigid rod of mass m and length L and is incident at speed v1 and angle with impact parameter b. A real cue stick is moderately flexible and contains a flexible tip, but we ular block of glass at three angles of incidence. The time average value of ctang struck 8 mm off center, in the manner shown by the inset. 208Rod Cross license or copyright; see http://ajp.aapt.org/authors/copyright_permission assume here that the stick is sufficiently rigid that it has a well-defined center of mass speed and angular velocity both before and after the collision" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000413_robot.2003.1242208-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000413_robot.2003.1242208-Figure6-1.png", "caption": "Fig. 6 : Force components a t toller", "texts": [ " The position and motor current of each axis, force at the roller, and range sensor signal were recorded during forming. Figure 4 shows the forming experiment. Figure 5 shows an example of ttie blank and the finished product. The product was formed to the shape of the mandrel. Next, t,he forming force applied to the material by the roller was measured. The force component in line with the movement of the roller is defined as F x , the normal force component against the surface of the mandrel is defined as F y , and t,he tangential component to the mandrel rotation is defined as FZ (Fig. 6 ) . F x , FY and FZ are plotted for the displacement, X of the roller along the mandrel (Fig. 7). The velocity of the roller in the X direction was 0.1 mm/S, and the angular velocity of the 0 axis was 12Orpm (411 rad/s). The movement of the roller for one turn of the mandrel was 0.05 mm/rev. F y , which forces the material onto the mandrel, was about 200N constantly. The feeding force of the roller, F x , was about 80N at first, and later decreased gradually as the forming proceeded. As the width of the flange, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001012_ias.2003.1257791-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001012_ias.2003.1257791-Figure1-1.png", "caption": "Fig. 1. Basic Model for analysing zero-sequence current response to voltage test vectors Vk, k=1\u20266.", "texts": [ " These are superimposed on the normal PWM waveform. The test signal method is described in [1] and involves equal and opposite test vectors (V1,V4), (V3,V6) and (V5,V2) of short duration applied on successive PWM cycles. The voltage test vectors excite a 3-phase zero sequence current response in the machine, each phase corresponding to a particular test vector pair. The method exploits the fact that, over successive PWM cycles, both the flux and rotor position may be considered constant for low speed operation. Fig. 1 shows a simplified model for analysing the response of the zero-sequence currents to the voltage test vectors. The machine phase is modelled as a (small) back-emf in series with a leakage inductance assumed to vary sinusoidally under the effects of main-flux/tooth saturation and rotor slotting. In practice, the model should also contain the winding capacitances to earth; these are considered in Section IV. Considering only the fundamental space distribution of the leakage inductance around the machine, then the variation of the leakage inductance may be represented by: _ _ 4 4 _ _3 3 2 2 _ _3 3 - cos(2 ) - cos( ) - cos(2 ) - cos( ) - cos(2 ) - cos( ) a sat e slot r b sat e slot r c sat e slot r l l l t l n t l l l t l n t l l l t l n t \u03c3 \u03c3 \u03c3 \u03c3 \u03c0 \u03c0 \u03c3 \u03c3 \u03c3 \u03c3 \u03c0 \u03c0 \u03c3 \u03c3 \u03c3 \u03c3 = \u2206 \u03c9 \u2206 \u03c9 + \u03c6 = \u2206 \u03c9 \u2212 \u2206 \u03c9 \u2212 + \u03c6 = \u2206 \u03c9 \u2212 \u2206 \u03c9 \u2212 + \u03c6 (1) 0-7803-7883-0/03/$17", " In practice, higher harmonics 2 j em \u03c9 and k rn\u00b1 \u03c9 of the _ satl\u03c3\u2206 and _ slotl\u03c3\u2206 variations will also exist. In addition, some induction machines also exhibit intermodulation harmonics of the form 2k r j en m\u03c9 \u00b1 \u03c9 . The test vectors Vk, k=1\u20266 correspond to the switching states defined in Fig. 2. Switching state (100) requires switches (S1,S6,S2) ON while state (011) requires switches (S4,S3,S5) ON. The technique applies a test pulse Vk of 15\u00b5s in the centre of the PWM cycle. With the assumptions of Fig. 1, the resulting phase and zero sequence currents across the pulse duration will have a constant derivative. If we denote the j th phase current (derivative) response to Vk as ( )k jdi dt , then applying V1 (100), from Fig. 1 we have: (1) (1) (1) 0 ab a a ca c c bc b b di E l e dt di E l e dt di l e dt \u03c3 \u03c3 \u03c3 = \u22c5 + = \u2212 \u2212 = + (2) from which the zero-sequence current derivative during the V1 pulse is derived as: (1) (1) (1) (1) 0 (1) 0 ab bc ca a b c a c a b c di di di di dt dt dt dt di e e eE E dt l l l l l\u03c3 \u03c3 \u03c3 \u03c3 \u03c3 = + + = \u2212 \u2212 + + (3) From applying V4 (011) we get: (4) 0 a b c a c a b c di e e eE Edt l l l l l\u03c3 \u03c3 \u03c3 \u03c3 \u03c3 = \u2212 \u2212 + + (4) At low speed, the back-emf is small and the magnitude of the second term on the right hand side of (3) may be neglected, being \u2248 0", " This measures a voltage proportional to the rate of change of flux in an NT turn (per unit length) coil tightly wrapped around a cylindrical former of cross sectional area A and relative permeability \u00b5r. The former is flexible and itself forms a circuit C about the current to be measured. If NC is the number of circuits that the former makes about the currents, and the three phase currents are passed NI times through C, the from Ampere\u2019s Law one has: 0 0. T C I m r r k k N N N N d dV NA H dl NA I dt dt = = \u00b5 \u00b5 = \u00b5 \u00b5 \u2211\u222b (8) where N and \u00b5 r determine the sensitivity. IV. ZERO\u2013SEQUENCE CURRENTS In Fig. 1, only the zero-sequence current arising from machine leakage imbalance was considered. In practice, further zero-sequence components arise from both inter-winding capacitance and the machine winding capacitance to the earthed machine frame [13][14][15]. The response of these components to a sudden voltage step takes the form of lightly damped high frequency oscillations that can interfere with the current derivative measurements. The transient oscillations arising from the inter-winding capacitance is in the megahertz region [13][14] and is found to disappear in a few \u00b5s" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001920_1.2988480-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001920_1.2988480-Figure6-1.png", "caption": "Fig. 6 Accuracy of fit t", "texts": [ " Then, we estimate each value f , r, and l in the same manner as for . After these estimates, e calculate t according to Eq. 9 for each , , r, and l. t orresponds to standard deviation in statistics and means the acuracy of fit, which expresses whether the theoretical tooth surace fits the measured data or not. t is calculated using the folowing equation: t = S F n 9 here S denotes the mean radius of the gear. When the theoretical ooth surface fits the measured data Mi i=1,2 , . . . ,n exactly, t s zero. In other words, as shown in Fig. 6, when the value of t s small, we consider that the theoretical tooth surface fits the easured data Mi well, and in reverse, when the value of t is arge, the theoretical tooth surface does not fit the measured data i well. Hence, the factor indicating the smallest t is considered s a first factor of the occurrence of the error. However, it is uncertain whether the errors of the other factors re small or not. Therefore, we estimate the values of the other actors separately and calculate each t using the estimated values f the first factor and " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003573_s12239-011-0008-x-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003573_s12239-011-0008-x-Figure3-1.png", "caption": "Figure 3. Hypoid gear rotation direction and torque direction.", "texts": [ " pattern was measured by applying an appropriate volume of a gear-marking compound to the gears before the two sets of gear teeth were engaged with each other to turn and by copying the imprint on the gears with tape after turning. The tooth contact pattern was then checked. Figure 2 shows the measured tooth contact pattern. 2.2. Positions of the Deflection Measurement and the Axle Noise Section of Input Torque Conditions on the Vehicle The Hypoid gear rotation direction and torque direction are shown in Figure 3. The measurement position in the direction test is shown in Figure 4, and the axes (E, P, G, and \u03b1) of the hypoid gears are shown in Figure13. Except for the upper/lower and right/left torsions of the pinion and ring gears and the displacement, the real deflections were measured with a trigonometric function (Coleman, 1975). This test result also included tooth deflection. Moreover, each displacement sensor was fixed and measured based on the inner bearing assembly of the pinion gear. Because the input revolution in the direction test was 5~10 rpm and because oil was not used, the test was performed with the low revolution to prevent the teeth from being damaged by the heat generated in the surface of the gear under the torque condition" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000585_j.intermet.2006.03.012-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000585_j.intermet.2006.03.012-Figure6-1.png", "caption": "Fig. 6. Definitions of bead height, diameter and contact angle.", "texts": [ " The TieNi and TieFe intermetallic alloy beads were formed under Are4%H2 gas by 3DMW process. Fig. 4 shows morphological changes of TieNi and TieFe beads depending on the arc peak current. As the arc peak current increased from 5 A to 30 A, the bead diameter increased from 0.5 mm to 2.0 mm. An oxidized area on the substrate was observed at the peripheral region of a bead from 1.1 mm to 3.3 mm in diameter. Fig. 5 shows the diameter, height and contact angle of these alloy beads. The contact angle with the substrate was measured as shown in Fig. 6. As the arc peak current increased, the height and the contact angle decreased, while the bead diameter increased from about 0.8 mm to 1.7 mm in the TieNi system and from 0.8 mm to 1.9 mm in the TieFe system. The reactivity between a bead and substrate is thought to increase at peak arc currents over 15 A in the TieNi, and 20 A in the TieFe systems because the contact angle decreased sharply at these currents. The concentration of oxygen and nitrogen in a bead was measured by an oxygen/nitrogen gas analyzer (EMGA-520)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000423_ip-b.1987.0033-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000423_ip-b.1987.0033-Figure2-1.png", "caption": "Fig. 2 Phasor diagram", "texts": [ " This voltage, neglecting the losses and the overlapping interval of the rectifier, is equal to the rotor rectified voltage. Then, the slip s is given by nEA s = i\\-df (1) Hence, the motor speed can be controlled by the duty factor dF of the chopper and the smoothing capacitor voltage Ed. Let the line reactor current be ir = r sin (2) The relationship between the input and output power of the PWM inverter is given by EdIa = 3\u00a3s/ rcos where -Es = Ir(Rr+jcosLr) (3) (4) and Ea is a voltage generated by the inverter. Fig. 2 shows the phasor diagram of the inverter voltage Ea and source voltage, Es. Hence, the line current Ir and power factor cos (f> can be adjusted by varying the PWM voltage Ea and its phase angle a which are generated by the inverter (see Fig. 2). In general, a conventional inverter supplies a recovery power with a lagging power factor. The proposed system can obtain a power factor of 100% because the power factor of the recovery power can be adjusted to arbitrary values as seen in Fig. 2. 3 Analysis of the characteristics When the smoothing reactor is sufficiently large and Id is regarded as a direct current, the 3-phase power loss in an induction motor and the voltage drop during the current overlap can be expressed by the equivalent resistances on the DC side [8, 9]. Then the total DC side resistance is Rm = {2R\\ + 3(X\\ + X2)/n}s + 2R2 + Rd (5) where the stator impedance is referred to the rotor side. Fig. 3 shows the DC equivalent circuit of the system, and the equivalent circuits of the chopper in operation are illustrated in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000978_2004-01-2911-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000978_2004-01-2911-Figure1-1.png", "caption": "Figure 1 - Forces and moments acting on the piston", "texts": [ " The numerical improvements greatly enhanced the computational efficiency of the simulation. Then, we discuss the effects of skirt surface roughness and waviness, skirt profile, skirt flexibility, oil film thickness and piston liner clearance on skirt-friction and impact forces, and skirt-friction power loss. Finally, we summarize there results in the form of a scaling law to illustrate the basic phenomena. The equations of motion for piston are the traditional Newton\u2019s law of vertical and lateral motion and rotation applied to the piston in the standard manner. Figure 1 shows the forces and moments acting on the piston. These forces and moments come from interactions of the piston with the liner, rings, wrist-pin, connecting rod, cylinder pressure, and inertia. The nomenclature section as well as References [1,2,22] explain in detail the definition of parameters and the forces and moments on the piston. The primary motion - the piston position, speed and acceleration along the axis of the cylinder - can be calculated as a function of crank speed [1]. Figure 2 shows the piston geometry and definition of piston transverse movements, te and be , i" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001207_tie.2006.874266-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001207_tie.2006.874266-Figure7-1.png", "caption": "Fig. 7. Schematic diagram of the articulated robot arm with two degrees of freedom.", "texts": [ " Thus, the controller of the industrial robot arms does not directly control the position and the velocity of the arm itself. It controls the angle and the angular velocity of the motors. Usually, the position loops and the velocity loops of the servo controllers are composed of the constant feedback gains. The parameters of the servo controllers are tuned to obtain high performance of the servo motors. Thus, changing the structure of the closed-loop control system is rather difficult. The proposed feedforward compensator is, therefore, designed to improve the performance of the contour control of the robot arm. Fig. 7 shows a schematic diagram of the articulated robot arm with two degrees of freedom. The kinematics of the articulated robot arm with two degrees of freedom are represented by x = l1 sin \u03b81 + l2 sin(\u03b81 + \u03b82) y = l1 cos \u03b81 + l2 cos(\u03b81 + \u03b82) (15) and the inverse kinematics are represented by \u03b81 = sin\u22121 ( y\u221a x2 + y2 ) \u2212 sin\u22121 ( l2 sin \u03b82\u221a x2 + y2 ) \u03b82 = \u00b1 cos\u22121 ( x2 + y2 \u2212 (l1)2 \u2212 (l2)2 2l1l2 ) (16) where (x, y) is the tip position of the robot arm in the working coordinates, (\u03b81, \u03b82) is the angle of the links in the joint coordinates, l1 is the length of the first link, and l2 is that of the second link" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003658_09544062jmes1973-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003658_09544062jmes1973-Figure1-1.png", "caption": "Fig. 1 Axis system, velocities, and forces", "texts": [ " T is the Tachikawa number [7] given by T = \u03c1Aq2 2 mg (6) where m is the mass of the ball. This represents the ratio of the initial flow inertia force on the ball to the Proc. IMechE Vol. 224 Part C: J. Mechanical Engineering Science JMES1973 at Univ of Connecticut / Health Center / Library on May 30, 2015pic.sagepub.comDownloaded from weight of the ball. It can thus be seen from equations (1) to (3) that the governing parameters of the problem are the product side and drag force coefficients and the Tachikawa number. The axis system and the aerodynamic forces are illustrated in Fig. 1. To solve the trajectory equations (1) to (3), a knowledge of the drag and side force coefficients is required. These parameters have been measured by a number of researchers in the past using a variety of experimental techniques. Experiments have been carried out on real cricket balls and model cricket balls of various scales, for different seam angles, with and without spin around the seams. These are fully discussed in the papers by Mehta in references [1] to [3]. In this section a compilation of these results is made, although it must be acknowledged that doubts exist concerning the adequacy of some of the experimental techniques that have been used", " This is most clearly seen in the experiments in reference [14], although the fall in side force coefficient occurs at an anomalously high Reynolds number in this case, presumably because the large-scale model that was used in these experiments was unrepresentatively smooth. At higher Reynolds numbers the force coefficients are close to zero. It would also be expected that the drag coefficient would also decrease in this Reynolds number range (the drag crisis referred to above), although the experimental Reynolds numbers in Fig. 1 are not high enough to reflect this. It is further to be expected that the values of side force coefficient with Reynolds number will depend on the angle of inclination of the seam to the flow and also on the spin rate of the balls. However, specific trends are hard to discern from the experimental results, although the results tend to indicate an optimum seam angle of around 20\u25e6 (i.e. for maximum side force coefficient) and the results of reference [4] suggest an increase in the side force coefficient with spin rate" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002873_05698197108983236-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002873_05698197108983236-Figure1-1.png", "caption": "Fig. 1 -Porous Gas-Bearing Spindle.", "texts": [ " This is not serious however, since the high specified unit loads result from unbalance and will be experienced only momentarily when passing through a critical speed. The stiffness and damping qualities of the bearing near the concentric position is of more importance than minimum clearance at the high load condition. c) Elastically mounting journal bearings could prove to be an effective method of stabilization. Three porous bearing designs are described (3, 4, 5,). The segmented porous journal bearing (5) for an air bearing supported grinder spindle is of particular interest. Figure 1 depicts the spindle and bearing. As shown, each bearing consists of three sectors individually mounted on spherical seats. The shoe radial positions are set for the desired shaft clearance by adjusting screws supporting the spherical seats. Once the desired clearance is established, the shoes are interconnected through the clamps. Spherical motions of the shoes are limited by the degree of flexibility in the clamps; however, the radial position of the shoes, once set, is fixed. The porous material is used in self-lubricating graphite" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000718_1.1829068-Figure16-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000718_1.1829068-Figure16-1.png", "caption": "Fig. 16 Side view of wobbling member \u201ewith mating gear\u2026 in \u201ea\u2026 position f\u00c40, h\u00c4hmax and \u201eb\u2026 position f\u00c4p\u00d52, h\u00c40", "texts": [ " The theoretical efficiency of a gear pair is based on the amount of slipping that occurs in the mesh and the friction losses which result. In order to understand the correlation of a cam pair to a gear pair, it is thus necessary to find the friction losses incurred in the cam pair. First, a complete understanding of the motion of the cam is necessary. As the input link rotates, the linear position of the cam in the slot varies sinusoidally. Let f represent the position of the input link and h represent the position of the cam in the slot, as shown in Fig. 16. Then, h5D sin b cos f (13) where D is the distance from the center of the wobbling link to the contact point on the cam. A differential change in h is thus dh52D sin b sin fdf (14) and the approximate total distance traveled through one revolution of the input is MARCH 2005, Vol. 127 \u00d5 275 shx?url=/data/journals/jmdedb/27802/ on 03/07/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F h total54D sin b (15) From quasi-static equilibrium of the mechanism, the magnitude of the torque applied to the mechanism at the cam is equal to the sum of the input and output torques, but has opposite sign", " Since the input torque is a unit torque, and the ideal output torque is equal to the negative of the reduction ratio as given in Eq. ~9! Tcam5 21 sec b21 215 21 12cos b (16) and the normal force is Fn5 2Tcam D cos b (17) So the total work lost in friction is approximately W loss5mFnh total5 4m tan b 12cos b (18) Now in order to find the efficiency of cam pair ~1,28!, the input work must be known. In particular, the force moving the cam in the h direction must be found. To do this, consider the relative motion of the wobbling member on its mating gear as shown in Fig. 16. The motion is characterized by rolling about the pitch-line axis. One significant difference between the wobble gear and its kinematically equivalent bevel-gear train is that instead of a carrier link applying the force necessary to produce motion, the \u2018\u2018wave-generator\u2019\u2019-like input link of the wobble gear actually applies a moment about the instantaneous axis of rotation ~or screw axis!. According to the Kennedy\u2013Aronhold theorem @9#, this axis must pass through both the pivot point of the wobbling member with respect to ground and the contact ~pitch" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000948_detc2005-84109-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000948_detc2005-84109-Figure1-1.png", "caption": "Figure 1. Some critical configurations of the slider-crank mechanism", "texts": [ " those based only on kinematic structure analysis; 2. those based on analytical criteria. Due to their simplicity, the first type is always discussed and included under various algebraic forms in textbooks. The second type is less frequently reported. Although the approaches for computing the d.o.f. are based on different simplifying hypotheses, it should be acknowledged that there are significant differences in the definitions of d.o.f. herein reviewed. These affect the d.o.f. estimate. For example, in the slider-crank shown in Figure 1, only one variable is required to specify the relative positions of all links or to determine their positions w.r.t. the ground link. Thus, according to Definitions 2, 3 and 4, this linkage has F = 1 d.o.f. However, since the slider-crank is in critical configurations (see Figure 1), from Definitions 1 or 5, one would conclude that the mechanism has instantaneously F = 2 two d.o.f. In fact, for a given infinitesimal displacement of the input link, the infinitesimal displacements of the remaining links are not uniquely defined. This ambiguity follows directly from the different definitions of d.o.f. and not from simplifying hypotheses. [1] Shigley, J.E., Uicker, J.J., 1980, Theory of Machines and Mechanisms, McGraw-Hill Book Company. Copyright c\u00a9 2005 by ASME url=/data/conferences/idetc/cie2005/72591/ on 03/06/2017 Terms of Use: http://www" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003011_gt2010-22632-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003011_gt2010-22632-Figure3-1.png", "caption": "Figure 3. SINGLE TOOTH REPRESENTATION OF WHOLE GEAR", "texts": [ " The mesh uses a single tooth with rotationally periodic boundaries, meaning the computational time taken for each simulation is cut down considerably, and also allows for greater mesh resolution for the single tooth. There is an assumption made by doing this that the flow is periodic across a single tooth, and this is explored in greater detail in the section on verification of the results. The periodic boundary for the entire domain follows the spiral angle of the tooth when viewed from the front of the gear, and passes through the toplands (the upper flat surface) of the gear teeth, and extends axially through the whole domain. This is shown in Figure 3, which shows in grey, the single tooth domain when viewed from the front of the gear. Equations 1 & 2 are the governing equations used by FLUENT for mass and momentum respectively. In Equation 2 , S represents body momentum sources, such as gravity. Further information about FLUENT\u2019s numerical treatment of the governing equations for fluid flow can be found within the user man- ual [12]. \u2202\u03c1 \u2202t +\u2207 \u00b7 (\u03c1ui) = 0, (1) \u03c1 ( \u2202ui \u2202t +ui \u00b7\u2207ui ) =\u2212\u2207p+\u00b5\u2207 2ui +S, (2) All simulations reported in this paper have second-order discretisation, temporally and spatially, which reduces numerical diffusion", " This boundary condition fixes the mass flux across the boundary and allows the total pressure to change. Due to computational time constraints, this was not possible to do for all gears. The outlet boundary for all simulations is a pressure outlet condi- tion, in which static pressure is prescribed as zero Pascals, once again replicating the conditions in which the experiment was run. Four surfaces are used for post-processing of static pressure and mass-flow-rates for analysis, and these are all shown in Figure 4 . The perspective for this view is marked on Figure 3, and shows a section of the domain through a periodic boundary. Nose Restriction is axially located half way along the restriction at the \u201cnose\u201d to the shroud. Shroud Outlet is at the exit to the shroud, where the flow passing between gear and shroud passes into the back chamber. These surfaces are annular in shape for a whole gear. Gear Inlet is effectively a conical surface which intersects the domain at a point just after the entrance to the gear teeth and is inclined at the back cone angle (90\u25e6 minus the pitch cone angle)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001806_10402000801918056-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001806_10402000801918056-Figure1-1.png", "caption": "Fig. 1\u2014A schematic diagram of the cam and tappet contact geometry.", "texts": [ " The objective of this investigation was to evaluate the friction reduction potential of surface textures in a direct acting mechanical valvetrain application with special attention to the ability to manufacture the textures relatively easily and less expensively. In this investigation, textures were created by two different techniques: (a) regular patterns using a diamond tool because it was thought to be less expensive than laser technology, and (b) random pattern by shot peening, which is a proven inexpensive technology in the industry. The patterns were created on tappet shims from a production 3.0 L engine. Figure 1 shows a schematic diagram of the contact geometry between the cam lobe and the tappet shim. The tappet shim is free to rotate on a groove on top of the tappet. The centerline of the cam lobe is offset from the centerline of the tappet and the tappet shim to facilitate the rotation of both the tappet and the tappet shim, which reduces friction. The shims are made out of AISI 52100 steel having a hardness of 62 Rc and have 0.1\u2013 0.2 \u03bcm (4\u20138 \u03bcin) centerline average roughness. The camshafts were made out of induction-hardened chilled cast iron having a surface hardness of Rc 50" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000057_s0069-8040(08)70029-3-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000057_s0069-8040(08)70029-3-Figure1-1.png", "caption": "Fig. 1. Coordinate systems for common electrode geometries. (a) Cylindrica symmetry: (i) ring-disc electrodes, (ii) tubular electrodes; (b ) Cartesian symmetry channel electrodes; (c) spherical symmetry: dropping mercury electrode.", "texts": [ " A summary of limiting currents calculated for various electrode geometries will be found in Table 3 (p. 384). Double electrodes are particularly useful in kinetic studies. Intermediates produced on the generator (upstream) electrode are transported to the downstream electrode where they react further. This is useful for the study of short-lived species, the quantity reaching the second References p p . 434-441 TABLE 1 The convective-diffusion equationa - = D v 2 c - vvc in various coordinate systems; ac a t w (5, 0 Diffusion Convection Cartesian Cylindrical polar Spherical polar aSee Fig. 1 for explanation of coordinates. 361 electrode depending on any homogeneous reactions or decomposition occurring in solution. The most widely used of these is the RRDE. It is a prerequisite to know what fraction of species which reacts at the upstream electrode reaches the downstream electrode without kinetic complications. This, for &/at = 0, is the steady-state collection efficiency, N o , and will be calculated for common electrode geometries. It will be seen that the form of the expression for N o is quite general for double electrodes. 2.3.2 Rotating electrodes In this section, we consider mass transport-controlled currents to disc and concentric ring electrodes on a planar spinning disc surface. For other less common rotating electrodes, e.g. rotating hemisphere, see Table 3. (a ) Rotating disc and rotating ring electrodes The problem of laminar fluid flow to a rotating disc has been amply discussed in the literature [7, 101. We may describe the velocity components as [17,18] (see Fig. 1) u6 = r w G ( y ) U, = - ( c J u ) \u201d ~ H(y) u, = r w F ( y ) where y is a dimensionless distance from the electrode surface with CJ the rotation speed in rad s-l . The variation of the flow functions F, G , and H with distance and resultant streamlines are shown in Fig. 2. The particular form of the velocity components satisfies the Navier - Stokes equation and the equation of continuity. Note that u, is independent of the radial coordinate. We are interested primarily in the convection pattern close to the electrode surface in order to calculate the flux of electrons", " 434-441 362 the steadv state is where we neglect radial diffusion as being negligible compared with radia convection. We are now in a position to calculate the diffusion-limitec current at a rotating disc or rotating ring electrode. The appropriatl boundary conditions are z + m C\u2019C, bulk concentration r = O c = o z = o c = o 363 If we define the dimensionless variables c - c, y = - C, then it can be easily shown that eqn. (17) is transformed into ay a2r with boundary conditions x = o y = l where p = 1 for a disc electrode, p = 3 for a ring electrode and r l , r2 and r 3 are defined as in Fig. 1. The limiting current at a disc or ring electrode is given by rP iL = 2nnFD r(g),, dr rp-1 The integral (22), after undergoing Laplace transformation [ 201 with respect to E p , becomes References p p . 434-441 364 which is the Airy equation [21] whose solution with the boundary conditions (21) is, after the inversion A i ' ( 0 ) A i ( 0 ) r ( 3 ) ti l I3 The Nernst diffusion layer thickness as defined in eqn. (6) is given by c , /( ac/az), and is When p = 1 (disc electrode), 6, is independent of r ", " It turns out, as is to be expected, that for high Schmidt numbers (thin diffusion layer) the mass transport in the appropriate dimensionless variables is virtually iden tical for both electrodes. (a ) Tubular electrodes Laminar fluid flow in tubes has been described by Levich [ 3 ] . An entry length, 1, , is necessary to establish Poiseuille flow, given approximately by 1, - 0.1R - Re ( 5 2 ) where R is the radius of the tube: this is derived from the boundary layer thickness a t a flat plate. The velocity profile after this point (see Fig. 3 ) is there being no radial or angular convection. uo is the fluid velocity in the centre of the tube and coordinates are defined in Fig. 1. The timeindependent convective diffusion equation is Making the assumption of a thin diffusion layer compared with the tube radius, we approximate (55) 2 ( R - r ) R = 2p u, = uo 3 7 1 where. as before c - c , CCC Y = -_ and Analogously to rotating electrodes, we take p = 1 for the upstream of two electrodes (generator) and p = 3 for the downstream (detector). Since, in electrochemical experiments, radial diffusion will be much less than axial convection, we can say that and eqn. (54) becomes a7 a27 X g = g This equation is of exactly the same form as the dimensionless convectivediffusion equation at the RDE (p", " However, for total kinetic control (irreversible reaction at the foot of the wave), the flux is uniform as radial convection is uniformly zero along the tube. Digital simulation [37] has shown that the approximations made by Levich are valid under most conditions: limits to his assumptions are given, particularly with regard to potential scan rates. ( b ) Channel electrodes The velocity profile at a channel electrode is u, = uo (l+) where v o is the maximum fluid velocity in the x direction and h is the half-height of the channel as shown in Fig. 1. Poiseuille flow will be established after an inlet length I, which, according to Schlichting [4] is given by 1, = 0.11 (uo h / v ) . Neglecting the contribution of axial diffusion as being small compared with axial convection, we have ac a z C ax ay2 u , - - - - Assuming a thin diffusion layer Then, by putting we can follow exactly the same argument as at the tubular electrode. The reason for the similarity is that the approximations we have made 373 regarding diffusion and convection have removed any curvature effects", " In the particular case where p = 1 (no adsorption) For the rotating disc electrode, we plot i-\u2019 vs. w-\u2019\u2019\u2019 : we can determine D from the slope and ik from the intercept. Application of this relation has been thoroughly discussed [144] (Fig. 11). An alternative method is to eliminatej (or i) from eqn. (115) and write k m l : = G W I , - [OI*) (123) If we define then eqn. (123) reduces to 1-7, TP = K Like eqn. (121), this is only valid for the whole electrode when the surface is uniformly accessible. I t is most commonly evaluated graphically by using a nomogram such as that shown in Fig. 1 2 [145]. In general, because of non-uniform accessibility of the electrode, equations for current-voltage curves have to be solved either by using approximations or numerically. For historical reasons, the dropping mercury electrode was the first t o be treated theoretically [ 146- 1491 and the equations obtained depend on whether the steady-state, expanding-plane, or expanding-sphere models are utilised [ 1501 . Matsuda and co-workers have provided generic approximate expressions for current-voltage curves for first-order reactions at a variety of hydrodynamic electrodes" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000262_tmag.2004.824897-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000262_tmag.2004.824897-Figure3-1.png", "caption": "Fig. 3. 3-D finite element mesh (except the air).", "texts": [ " , , and are the numbers of the maximum or minimum values of the flux density of the radial direction, the rotation direction and the direction of the th element, respectively. , , and are the amplitude of the flux density of major and minor hysteresis loops of the radial direction, the rotation direction, and the direction, respectively. Fig. 2 shows the analyzed model of a squirrel-cage induction motor, the rotor of which is skewed with one rotor slot pitch. It is 1/2 of the whole region because of the periodicity. Fig. 3 TABLE I ANALYZED CONDITIONS Fig. 4. Distributions of flux density vectors (0 rpm). (a) Upper section (z = 41 mm). (b) Lower section (z = 1 mm). shows the 3-D finite element mesh. Table I shows the analyzed conditions. Fig. 4 shows the distributions of the flux density vectors. It is found that the flux density vectors in the upper section are larger than those in the lower section due to the skew. This tendency becomes small as the rotation speed is high, and disappears at the synchronous speed (1500 rpm)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.136-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.136-1.png", "caption": "Fig. 14.136 Diagram of automatic system for transporting concrete mix on dam construction site", "texts": [ "8 Types of Concrete Works Covered by Automation The automation of concrete works covers primarily the transport and distribution of concrete mix, and then the removal of the layer of corroded concrete from reinforced concrete structures, applying a new layer of concrete, the fabrication of reinforcement, the removal of irregularities in the surface of freshly set concrete, and the vertical shifting of formwork in the sliding erection process. Several applications of related techniques have been described [14.54]. The automation of concrete mix production in concrete batching and mixing plants is discussed in Sect. 14.3. Examples of the automation of the particular types of concrete construction tasks are provided below. Transport of Concrete Mix Most concrete mix transport automation solutions are found in the construction of dams [14.53]. The automation solution depends on the location and size of a dam. Figure 14.136 shows a diagram of an automatic concrete mix transport system used in the construction of a dam with a concreting work volume of 510 000 m3. The system covers the delivery of concrete mix from a concrete mixing plant to tipper trucks transporting it Rotary encoder Depth detection wire Measuring tower Trace wire Tower inclinometer Differentialtransformer transducer Adjustable guide Angular meter Inclinometer Hydraulic unit Base machine Load cell Measuring equipment Excavator Engine Brake Input/ output unit CRT monitor TV monitor Control modul Personal Computer Computer Control unit Control module Excavation wall irregularity measurement (fuzzy controller) Control panel Adjusting module Monitor Excavator positioning system Excavator load control system Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003068_j.jbiomech.2009.04.048-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003068_j.jbiomech.2009.04.048-Figure3-1.png", "caption": "Fig. 3. Coordinate systems for shoulder bones (Hogfors et al., 1991).", "texts": [ ", 2005; Herda et al., 2003; Hogfors et al., 1991; Six Dijkstra et al., 1997). Among these, Hogfors et al. (1991) obtained ideal shoulder rhythm solutions. In the experiment, the orientation angles of the scapula bone and the clavicle bone were measured when the arm was elevated at an angle y in the abduction plane, which has 451 angle with respect to the coronal plane in Fig. 2, where y varied from 01 to 1201 with intervals of 151. Three bones\u2019 body-fixed frames are used to define their orientations in Fig. 3. The global frame (x, y, z), attached on the sternum, and the clavicle local frame (x1, x2, x3) have the same origin O. The origin of the scapula local frame (x1, x2, x3) is located at point Os. The origin of the humerus frame (k1, k2, k3) is at point Oh. The Euler angles for the clavicle bone are defined as ac, bc, and gc, and obtained from the clavicle local frame (x1, x2, x3) with respect to the global frame (x, y, z). The Euler angles for the scapula are as, bs, and gs from (x1, x2, x3) with respect to the global frame (x, y, z)", " The principle of transferring coupling equations from the Euler to the DH system is that the same posture can be represented by different orientation representation systems. The procedure is shown in Fig. 4. Both Euler and DH system have the same global coordinate system. However, DH local frames are different from Euler\u2019s frames. Figs. 5 and 6 show the postures corresponding to zero Euler angles and DH angles, respectively. The posture of zero Euler angles does not exist and it is used for calibration purpose. In Fig. 5, the local frames are same as those in Fig. 3. This section will illustrate the analytical inverse kinematics method by one example. By choosing ah \u00bc 451 and bh \u00bc 451, through Eqs. (1)\u2013(3), we obtain all Euler angles in Table 1 and the posture is in Fig. 7. Considering AC joint center for both systems, one has the following equations: R\u00f0ac;bc ; gc\u00de L2 0 0 0 B@ 1 CA 1 2 6664 3 7775 \u00bc T0 1\u00f0q1\u00de T 1 2\u00f0q2\u00de 0 0 0 1 2 6664 3 7775 (4) Bringing in all necessary terms in Eq. (4) yields sin\u00f0q2\u00de \u00bc cos\u00f0bc\u00de sin\u00f0ac\u00de (5) sin\u00f0bc\u00de \u00bc cos\u00f0q2\u00de sin\u00f0q1\u00de (6) Solving Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003451_jmes_jour_1967_009_028_02-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003451_jmes_jour_1967_009_028_02-Figure1-1.png", "caption": "Fig. 1", "texts": [ " By integrating the resulting velocity distributions across the film, Constantinescu found that the pressure gradients are related to the mean velocity components by - aP h2 urn = -%p(12+0.53&?0'725) - -:: 1 (6) wrn = -- The components urn and w,, which are solutions of the equations of motion, have to satisfy as yet the equation of continuity. If urn, w, are the mean pressure velocities, u,, the mean shear velocity and if the surface velocities of the bearing and the journal are given by V2 and V, respectively, the principle of continuity may be expressed as h2 ap - aP h2 a ~ p ( 1 2 + 0 * 2 9 6 R e ~ ' ~ ~ ) - -2z J - - - f - - f (Fig. 1) a a -\"h(urn+umc)l+~ (hwm)+(Vzy+ v 2 x tan 8- Vly) = 0 ax \"c For the case of 'almost' parallel surfaces tan 6 = 8. After substituting for u, and w,, the 'macroscopic' equation of continuity assumes the form Employing the notation of Fig. 2 one has cos TJ 2 1, sin TJ E 0 and h z c ( ~ + E cos 0). Since the tangential velocity of the bearing VZx = 0, JOURNAL MECHANICAL E N G I N E E R I N G SCIENCE Neglecting quantities of the order of C/R, one obtains -(hum,) = -+CW sin 0 , . . (8) a ax If the bearing is stationary, the second term on the right of equation (7) reduces to - Vly, and When equations (8) and (9) are substituted into the equation of continuity, the differential equation (10) in pressure is obtained: Vly = ICOS 8+e$sin 6 " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000182_1.2114888-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000182_1.2114888-Figure1-1.png", "caption": "Fig. 1 3\u201eUPS\u2026-S fully parallel spherical wrist", "texts": [ " Thus, redundancy of the information measured positions of more than two platform points decreases the sensitivity of the resolving scheme. Indeed, measurement errors can be smoothed, thus gaining accuracy of the DPA estimate. In order to make the presentation of the algorithm that is proposed in this paper more compact and clearer, and without losing generality, we refer especially for the case study in Secs. 4 and 5 to a spherical wrist of type 3 UPS -S here U, P, and S are for universal, prismatic, and spherical joint, respectively . The 3 UPS -S fully parallel wrist architecture Fig. 1 comprises a fixed base B1B2B3 connected through a spherical joint centered at point G to a mobile platform P1P2P3 . The latter is driven by means of three legs PiBi, i=1,2 ,3, of type UPS, with variable length li= Pi\u2212Bi , controlled by actuated prismatic pairs, which are connected to the base and to the platform by universal joints centered at point Bi and spherical pairs centered at point Pi, respectively. Each leg is instrumented by a sensor that measures its length. In this context, the leg lengths are taken as joint space variables", " In fact, it can be seen that the reciprocal locations of the vectors wi directly influence pW , while the manipulator configuration and the sensor accuracy have a great effect on b W / bW , since they are both responsible for the discrepancy wi= w\u0303i\u2212wi. However, besides these considerations, the choice of the locations of the vectors wi is also influenced by the overall hardware cost, i.e., the evaluation of the component arrays bwi and pwi have to require the use of fewer sensors than possible. In the context of 3 UPS -S fully parallel wrists such as the one depicted in Fig. 1, a satisfying compromise consists in taking wi such that wi = Pi \u2212 G i = 1,2,3 30 Thus, for every platform location vectors pwi, which in the literature are often referred to as pi= pwi, will be constant and a priori known. Conversely, vectors bwi, which in the literature are often referred to as Pi= bwi, will vary and need to be determined through simple kinematic computation from data acquired by a proper sensory system. As for the latter, since due to feedback control of the manipulator the legs li= Pi\u2212Bi are just instrumented with linear sensors, then, in order to uniquely and rapidly estimate the Pi\u2019s, we have to add just three rotary sensors, each of them to be placed so as to measure the angle i which describes the relative rotation of the triangle BiGPi about the axis BiG, i =1,2 ,3", "asmedigitalcollection.asme.org/ on 08/12/201 Note that the calculation of the generalized inverse p+ which requires a consistent computational burden is to be performed off-line, i.e., at 0.iv . Thus, when comparing Method 1 versus Method 2, not only is the real-time performance not affected negatively by such a calculation, but it enhances the computation by speeding up of the convergence of the PD algorithm, while, at the same time, increasing the accuracy of the DPA estimate refer to Sec. 3 . With reference to Fig. 1, the manipulator we have chosen for the simulations can be described as follows. Resorting to cylindrical coordinates, the platform vertices, i.e., the centers Pi of the spherical joints, are on three coplanar directions, equally distributed at 120\u00b0, all intersecting at point C, center of the mobile system Sp. The three points have distances pRpi from C given by pRp1 =600 mm, pRp2 =500 mm, and pRp3 =400 mm, respectively. The spherical joint G coincides with the same center C, i.e., G C. As for the base, the centers Bi of the universal joints are on three coplanar directions, equally distributed at 120\u00b0, all intersecting at point O, center of the fixed system Sb" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002940_ast.72.331-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002940_ast.72.331-Figure3-1.png", "caption": "Fig. 3. A schematic cross section of a supercapacitor manufactured on paperboard substrate.", "texts": [ " Because the same activated carbon ink was used in all experiments, the specific capacitance value remained constant at 30 F/g (for net activated carbon mass). Fig.2. A schematic cross section of a supercapacitor with graphite foil current collectors. In order to make the manufacturing process simpler, it is necessary to fabricate also the current collectors by printing. To achieve this polymer coated paperboard was used as substrate and the current collectors were made of Timrex graphite ink, figure 3. Using this ink about 20 \u2126-square resistivity for 80 \u00b5m thick layer was obtained. With the normal layout (current collector width 14 mm and length 50 mm) the ESR values measure were between of the order of 100 \u2013 150 \u2126. This resulted in low efficiency values: even with 1 mA current the efficiency was about 40 %. By changing the layout by widening the current collectors and printing two graphite layers on each other, the ESR was decreased to about 40 \u2126, which improved the efficiency with 1 mA to 71 %", " When a voltage of 1.2 V was applied, the current going through the electrolysis system with the typical supercapacitor dimensions was in the order of 0.5 mA indicating electrochemical reactions on the silver surface. Although the use of silver on the cathode side might have been possible, it was decided that both silver current collectors were to be coated with graphite ink to prevent the silver contacting with the electrolyte. The capacitor structure is otherwise similar to the one presented in figure 3. In this configuration the resistance of graphite ink causes a negligible increase of the ESR since the length of the current path is only the thickness of the ink layer and the cross-sectional area is 14 mm x 14 mm. Figure 4 shows a photograph of the supercapacitor. When preparing the supercapacitors with graphite coated silver current collectors also the thickness of activated carbon electrodes was reduced compared with the previous components. This resulted in a capacitance of 0.3 - 0.5 F In these cases typical ESR values were of the order of 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002336_iecon.2007.4460324-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002336_iecon.2007.4460324-Figure1-1.png", "caption": "Fig. 1. Dimensions and frequencies related to bearing.", "texts": [ " Usually bearing faults manifest themselves as rotor asymmetry faults [2], which are usually covered under the category of eccentricity-related faults. Mechanical failures (flaking or spalling of bearings) occur when fatigue causes small pieces to break and loose from the bearing. Even, under normal operating conditions, with balanced load and good alignment, fatigue failures may take place and for this reason it is very important to detect the failure in an early stage. Mechanical bearing failures result in a vibration movement. The frequencies of these vibrations are related to the mechanical parameters shown in Fig. 1 Apart from mechanical reasons (internal operating stresses, inherent eccentricity...), bearing currents [2] due to solid state drives can also damage the bearings. Coupling between the non-linear magnetic effects and nonlinear electric circuits should be token into account in order to determine the behavior of the electrical motor under fault conditions [3]. In this paper Finite Element Analysis (FEA) is proposed for the simulation of electrical machines under bearing fault. FEA analysis reveals itself like an accurate an easy method to determine the interaction between non-linear effects" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001980_sisy.2008.4664900-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001980_sisy.2008.4664900-Figure4-1.png", "caption": "Figure 4. Two-link foot and support area where ZMP can exist", "texts": [ " In this case, functional requirements that are realized with respect to the Cartesian (outer) coordinates (e.g. walking) are not fulfilled despite the fact that the change of joints coordinates may be perfectly realized while the system is falling down. Hence, in the case the humanoid's overturning, the point at which is formally fulfilled 0=\u03a3 xM and 0=\u03a3 yM , does not represent the ZMP, because dynamic balance has not been preserved. Let us consider the case of regular gait of a humanoid robot with two-link foot (Fig. 4). In that case the link representing toes is fixed with respect to the ground, whereas the link of the rear part of the foot moves. In view of the fact that the gait is regular the ZMP is inside the area covered by the toes link, and the system retains its dynamic balance. We should especially emphasize the difference between the situations illustrated in Figs. 3 and 4. In contrast to that shown in Fig. 3, the humanoid robot whose foot is sketched in Fig. 4, preserved its desired position with respect to the environment. In view of the fact that in the case shown in Fig. 4 dynamic balance has not been lost, the point inside the support area (and this is the area covered by the toes link) for which 0=\u03a3 xM and 0=\u03a3 yM , represents the ZMP. In all above examples, only rigid foot was considered. In the literature one can find examples of imprecise definition of ZMP as the point on the ground for which the ZMP conditions are fulfilled, i.e. 0xM\u03a3 = and 0yM\u03a3 = , without taking into account the size of support area. Let us suppose that, under the assumption that the foot by its entire area is in contact with the ground (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001111_soli.2006.328881-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001111_soli.2006.328881-Figure3-1.png", "caption": "Figure 3 Sketch ofthe helicopter model seen from the side", "texts": [ " The system is essentially nonlinear and the mathematical model can be linearized around the set point. 3.1 Helicopter body dynamics in the horizontal plane (Azimuth subsystem) Figure 2 shows a sketch of the helicopter body seen from the above. f(x,u) By balancing the torques acting on the body in the horizontal plane, as shown in figure2: where, St: Azimuth angle (Rad) , I: Moment of Inertia (kg mi), T : Friction torque(Nm), tr: Main rotor reaction torque (Nm), Z2 Torque generated by the side rotor (Nm). 3.2 Helicopter body dynamics in the vertical plane (Elevator subsystem) Figure 3 shows a sketch of the helicopter body seen from the side. By balancing the torques acting on the body in the vertical plane, as shown in figure3: ITW = l +,Gc+ Gf m (8) VI: Elevation angle (Rad), TG: Gyroscopic torque (Nm), 17: Friction torque(Nm), Im Moment of Inertia (kg m), Z': Lift torque generated by the main rotor (Nm), TC Centrifugal torque (Nm) , ZTn: Torque generated by the mass of the body (Nm) Now we have a model of the helicopter dynamics in both vertical and horizontal planes [9]. Combining the system denoted by (7), (8), motors, propeller and sensors models yields the complete helicopter body dynamics. After redefining the state variables the resulting model in formx = (x,u) , y = g (x) that completely describes the helicopter body dynamics is therefore given by, X2 (-z- sin(xl) + kGy,,(ulx6 cos(xl)) - BvX2 + a,x3 + b1x3 x4 ,(ul-x3 -2TIx4) X6 Ip -Bx6 - rK u1 + xg] + a2x2 + b2x ) X8 2 (U2 - X7 - 2T2x8) T LKX(1 - T\u00b0')ul - X9 g(x) = rXikv + yv) where, (9) yvG, Elevator angle read by sensor k Elevator constant YvoG Azimuth angle offset Y3Z Azimuth angle read by sensor k Azimuth constant Y3(0, Azimuth angle offset Ua Control voltage applied to rotors k Amplifier gain Xi is elevator angle and X5 is azimuth angle" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000868_tmag.2005.862764-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000868_tmag.2005.862764-Figure1-1.png", "caption": "Fig. 1. Phase portrait of PTOS.", "texts": [ " The PTOS uses a smooth nonlinear velocity profile by introducing the acceleration discount factor, . However, the discounted deceleration makes the rising time longer than that of the original time-optimal controller. The PTOS has analogy to the sliding mode control using a boundary layer [11]. Outside the boundary layer, the sliding mode controller uses a maximum output and inside the boundary layer, it uses a nonsaturated output. If we define the sliding surface (7) and the thickness of the boundary layer is set as (8) Fig. 1 shows the sliding surface of the PTOS in the phase plane. The sliding surface is continuous at , and the surface becomes linear within . When we adopt the boundary layer, the PTOS is more robust to model uncertainty. A larger value for produces a thinner boundary layer and a larger acceleration discount factor. That is, simultaneously changes the sliding surface and the boundary layer. Gain scheduling means changing the gain according to a certain rule. Generally, gains are scheduled by time or states of system" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure7-1.png", "caption": "Fig. 7. Datum surface is pitch cone.", "texts": [ " As the depictions in the introduction, the flank of transverse circular-arc brings the difficulty to manufacture and is not easy to be adopted in the engineering. In fact, W\u2013N gear in industry is also normal circular-arc gear. It can be seen that only when we change transverse circulararc into normal circular-arc, the theory of W\u2013N gear drive can be united with the theory of manufacture. The basic principle of this type of drive is similar to the previous, only the pitch cylinder in paralleled-axis drive is replaced by the pitch cone in intersected-axis drive. In Fig. 7, R\u00f01\u00dep ;R\u00f02\u00dep and IA are a pair of pitch cones and the instantaneous axis (line of action), respectively, d1, d2 are pitch angles. Additionally, u1, u2 are the rotational angles of two gears. Under the condition of constant speed ratio, the following relationship can be obtained: d1 \u00fe d2 \u00bc p R I \u00bc u2 u1 \u00bc sin d1 sin d2 When conjugate point P moves along IA with some motion rule and the pitch cones rotate about respective rotational axes at constant speed ratio, it forms a pair of conjugate directrixes on two pitch cones", " In the engineering, the determination of the directrix is very important, which not only affects the bending strength of gear tooth but also relates closely to the technics. During the design, we usually choose nominal spiral angle b firstly, once confirmed, the influence of the shape of the directrix to the strength is inapparent. So the main base of determining the directrix is convenient to manufacture. Here two types of valuable schemes are proposed. (1) Loxodromic-type normal circular-arc spiral bevel gear drive. Now the directrix is a loxodrome on the pitch cone, which is also named as the equal spiral angle curve. Taking C\u00f01\u00dep in Fig. 7 as the example, in the coordinate system of gear 1, the equation of pitch cone (datum surface) R\u00f01\u00dep and its unit surface normal vector are as follows: P\u00f01\u00de \u00bc m\u00bdsin d1e\u00f0k1\u00de \u00fe cos d1k1 n \u00bc cos d1e\u00f0k1\u00de \u00fe sin d1k1 ) \u00f026\u00de Where m and k1 are the parameters of the length and the rotational angle of straight generatrix of pitch cone R\u00f01\u00dep , respectively, e(k1) is circle vector function defined in Appendix A. According to differential geometry, let m \u00bc m0 exp\u00f0bk1\u00de b \u00bc sin d1 tan b \u00f027\u00de Here, m0 is the constant, and the constant b is the spiral angle" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001048_s11527-006-9199-4-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001048_s11527-006-9199-4-Figure2-1.png", "caption": "Fig. 2 Specimen and loading arrangement acc. to EN 408", "texts": [ " 1) and these three specimen types were evenly distributed among the five density sub-groups (Table 2). The growth ring orientation was defined as radial for shear planes and crack propagation perpendicular to the growth rings. It was defined as tangential if the shear plane was parallel to the growth rings (Fig. 1). Additionally 20 pieces containing pith and 102 pieces containing knots were tested to evaluate the influences of pith and knots on shear strength. The small cross section of the specimen (Fig. 2) did not allow to cover the full range of knot types and knot sizes. Nevertheless an increase in shear strength was expected for knots in tangential growth ring direction due to a dowel-effect. 2.2 Methods Figure 2 shows the loading arrangement according to EN 408 and the small cross section of the specimen (32 \u00b7 55 mm2). It is difficult to achieve a sufficient adhesion between the specimen and the steel plates [5]. Table 1 Comparison of target and actual density values of the five sub-groups (u = 12%, n = 260) Density sub-group q12, mean [kg/m3] Std. dev. [kg/m3] q12, k [kg/m3] I C 16 Target 370 30 310 n = 52 Actual 371 32 320 II C 24 Target 420 35 350 n = 52 Actual 421 42 345 III C 30 Target 460 40 380 n = 52 Actual 460 51 381 IV C 35 Target 480 40 400 n = 52 Actual 480 55 401 V C 40 Target 500 40 420 n = 52 Actual 500 57 426 A failure in the glued area of the test piece/steel plate interface results in invalid shear strength values" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002791_fskd.2010.5569106-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002791_fskd.2010.5569106-Figure2-1.png", "caption": "Fig. 2. Guide structure and sensor installation diagram", "texts": [ " In testing, CINCINNATI V5-3000, a five-axis highprecision CNC machine tool, the Y- axis feeding rolling guide of it was chosen for studying. The linear rolling guide, THK\u2019s SNR55RH, was level used, had two tracks and each owed two LM sliders spaced 500mm, and the travel length was 762mm. BK4321, a acceleration-vibration sensor, and BK2365, a vibration signal-amplifier, both of them were the products of the B&K Company in Denmark, were selected for testing. The linear rolling guide and sensor installation showed in Fig.2. III. SIGNALS PROCESSING BASED ON WAVELET PACKET AND FEATURE EXTRACTION It was necessary to denoise the vibration signals for approaching to the real signals, which inevitable associated with noises when in the data collecting. Wavelet packet analysis technology maintained excellent properties of the wavelet orthogonal basis and provided a more refined analysis method for the vibration signals [4]. The adaptive ability to different signal characteristics facilitated it selecting the best fundamental function of the signal depending on the frequency range" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002090_9780470264003-Figure12.8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002090_9780470264003-Figure12.8-1.png", "caption": "Figure 12.8 Exploded view of pressure recorder assembly.", "texts": [ " \u2022 Adjustment: Reduce customer intervention through tuning and adjustment. Use robustness techniques. \u2022 Poka-yoke: Use color codes and very clear instructions. and Dewhurst, 1990; Huang, 1996) In this approach we study the service disassembly and reassembly processes by identifying all individual steps, including part removal, tool acquisition, pickup and orientation, and insertion. The time standard in this procedure is the result of work by Abbatiello (1995) at the University of Rhode Island. An exploded view is shown in Figure 12.8. The worksheets shown below in the tables were developed to utilize the serviceability time database. The first step of the DFS approach is to complete the disassembly worksheet in Table 12.5. The DFSS team may disassemble the pressure recorder to reach the printed circuit board (PCB), the item requiring service. In the disassembly process, the team will access, row by row, several disassembly locations and record all operations in the disassembly worksheet. Subassemblies are treated as parts when disassembly is not required for service; otherwise, the disassembly operation recording will continue when removing them" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002401_j.apergo.2007.06.002-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002401_j.apergo.2007.06.002-Figure1-1.png", "caption": "Fig. 1. Experimental apparatus, pulling task and the three-dimensional coordinates.", "texts": [ " This pallet truck had a weight of 169 kg and a load capability of 2273 kg. The width (distance between the outer rims of the forks) and the length of the forks were 68.5 and 121 cm, respectively. Three plastic pallets, each with a cardboard carton, were used to load the weights. The weight of the plastic pallet and the cardboard carton was approximately 15 kg. To help the subject stop the truck at the end of the trial, two steel wings were attached to the sides of the two forks of the truck, each with a shock absorber on it (see Fig. 1). These shock absorbers were blocked by two steel stoppers near the target areas on the sides of the walkway to end a trial after the subject passed the target area. The truck was prevented from moving any farther when the shock absorbers were arrested by the stoppers, so as not to hit the subject. A piece of rubber foam was attached on the lower front of the truck to prevent the truck from running over the subject\u2019s feet. A passive cinematographic motion-tracking system (10 Eagle Digitals cameras with EVaRTs 4.01 Beta 12 software; Motion Analysiss Corp.) was used to collect the kinematics data. The motion-tracking system, calibrated to a residual error of less than 0.04 cm, was used to collect the three-dimensional coordinates (see Fig. 1) of the reflective markers (1 cm in diameter) on the toes and heels of the subjects\u2019 boots. The heel and toe markers were placed approximately 2 cm above the sole of the boots. The sampling rate of the motion-tracking system was 240Hz. At the end of the walkway, a target area (80 cm wide and 60 cm long) was delineated (see Fig. 1). On the target area, one of the three surface conditions including dry, wet, and glycerol-contaminated conditions was manipulated. For the dry conditions, a clean, dry floor was tested. For the wet conditions, a uniform film of water was applied using a sprayer. Under the glycerol-contaminated conditions, a mixture of 90% (by volume) glycerol (98+) and 10% water was applied using a paint-brush roller to generate a lowfriction condition. The reason for using the glycerol mixture was that it is an aqueous solution and is easy to clean", " The average walking speed was calculated using the horizontal distance between the locations of the first and the last landing of the left foot on the floor divided by the time. The spatial coordinates of the left heel marker were used for this calculation. The stride length immediately ARTICLE IN PRESS K.W. Li et al. / Applied Ergonomics 39 (2008) 812\u2013819 815 prior to the landing on the target area was calculated using the X coordinates of the heel markers of the landing foot. The cycle of the foot\u2013floor interaction for backward walking includes the landing of the toe, foot-flat, and taking-off of the heel. The toe slipped backward (in the negative X direction in Fig. 1) upon landing. This slip was termed slip I and the slip distance was termed slip distance I accordingly. Only the slips of the first foot landing on the target area were analyzed. Slip distance I was calculated using Eq. (1): Slip distance I \u00bc \u00f0\u00f0X 2 X 1\u00de 2 \u00fe \u00f0Y 2 Y 1\u00de 2 \u00de 1=2, (1) where (X1,Y1) and (X2,Y2) are the coordinates of the toe at slip-start and slip-stop. A negative Z velocity indicated a descending of the foot, while a positive Z velocity showed a rebound of the foot on the floor. A slip started when the first positive Z velocity of the toe marker occurred (see Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002815_1.4002527-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002815_1.4002527-Figure1-1.png", "caption": "Fig. 1 Frames of reference of the disk", "texts": [], "surrounding_texts": [ "i t\nr t i d m s i s c p d a s o r I b c f t b d u t p f a t r t p t f o T a L r t t A d w t\nb A p q s fl b t t a o b N i t T fl l l\n0\nDownloaded Fr\nmpacts should be evaluated, another approach utilizing the Hertz heory 9 must be applied.\nMost of the publications from the field of rotor dynamics are eferred only to the case when the rotor turns at constant revoluions. Moreover, if the disks or rotors turning at variable speed are nvestigated, the forms of the equations of motion derived by ifferent authors slightly differ and the differences depend on the ethod used for their derivation. In mathematical models, the haft is often represented by a beamlike body that is discretized nto finite elements. Derivation of the mass, stiffness, and gyrocopic matrices of the disk and of the shaft elements rotating at onstant angular speed were performed by Nelson and McVaugh 11 but nowadays, it can be found also in a number of further ublications 7 . The motion equations of a rotating symmetric isk with four degrees of freedom is derived in Ref. 3 . Using the ssumption of small displacements and rotations, the authors have implified Euler\u2019s dynamical equations by neglecting small terms f the second and higher orders. But the resulting relations are eferred again to the disk rotating only at constant angular speed. n Ref. 5 , the equations of motion for a Stodola rotor are derived y means of the impulse theorems. Even if they are related only to onstant speed of the rotor rotation, they can be easily extended or the case when the rotor turns with variable revolutions. Then, he terms of inertia moment for rotor spin motion would occur in oth moment equations. Such equations of a flexibly supported isk performing a spherical movement with variable speed are sed in Refs. 12,13 . The system has two degrees of freedom and herefore its vibration is governed by two equations. Attention is aid especially to nonlinear coupling between the disk and the rame. In Ref. 14 , there are derived motion equations of a disk nd of a shaft element. The authors used Lagrange\u2019s equations of he second kind. The resulting equations of motion are asymmetic because only one moment equation contains the inertia term of he spin motion of the disk or of the shaft element. Moreover, the rinciple of superposition, valid for small deformations and rotaions, is not satisfied here. If the rotations about the axes of the rame of reference used by the authors were performed in reverse rder, the inertia term would appear in the other moment equation. he equations of motion referred to a rigid disk rotating at varible speed can also be found in Ref. 15 . The author used agrange equations of the second kind for their derivation. Both esulting moment equations of motion do not contain any inertia erms corresponding to the spin motion of the disk. This form of he equations of motion was used also by some other authors 16 . s evident from this short survey, it is desirable to deal with the erivation of the governing equations of flexible rotors turning ith variable revolutions in more details and to discuss the menioned differences. Impacts of the rotors against the stationary part were studied oth numerically and experimentally by many investigators. zeez and Vakakis 17 showed that even if the excitation is a eriodic function of time, the produced vibration can also have a uasiperiodic or even chaotic character. Chang-Jian and Chen 18 et up an algorithm for analyzing a Jeffcott rotor supported by uid film bearings and performing disk-housing impacts. The earings are modeled by force couplings and the components of he bearing forces are calculated by approximate analytical relaions. The same model and the same approach was chosen by Chu nd Zhang 19 to investigate periodic, quasiperiodic, and the chatic vibration of a rub impact rotor system supported by oil film earings. Two different computational approaches based on the ewton and Hertz theories for investigating the nonlinear behavor of rotors supported by hydrodynamic bearings and performing he disk casing impacts were developed by Zapome\u030cl et al. 20 . he shaft is modeled by a beamlike body and the properties of the uid film bearings are linearized in the vicinity of the rotor equiibrium position. The same mathematical model is used in Refs. 21,22 , only the hydrodynamic bearings are represented by noninear force couplings. The impact forces are calculated by apply-\n21001-2 / Vol. 78, MARCH 2011\nom: http://appliedmechanics.asmedigitalcollection.asme.org/ on 01/28/201\ning the Hertz theory. The goal of the investigation is to analyze the propagation of the mechanical waves induced by the impacts and their transmission into the stationary part. In all these mentioned publications, it is assumed that the rotors rotate at a constant angular speed and that the impacts occur only between one disk and the casing.\nThe principal aim of this article is to contribute to the development of numerical procedures for the efficient analysis of the lateral vibration of rotors supported by hydrodynamic bearings and rotating with variable velocity. This requires deriving the governing equations, setting the corresponding computer algorithm, and performing the computational simulations. Particular attention is paid to the operating conditions when the rotor having several disks performs disk-housing impacts because of its angular acceleration. The disks are assumed to be absolutely rigid, except in a small vicinity of the contact area. The flexibility of the shaft and the dry friction between colliding bodies are taken into account. The hydrodynamic bearings are implemented into the mathematical model by means of nonlinear force couplings. To analyze this problem, a motion equation of the rotor turning at variable angular speed was derived. For its solution, the Runge\u2013Kutta method of the fifth order with an adaptive time step is applied.\n2 The Equation of Motion of an Absolutely Rigid Disk In the mathematical model, it is assumed that i the disk is a thin, absolutely rigid, axisymmetric body, ii the force transmission between the disk and the shaft is completed at the point where the middle plane of the disk intersects the center line of the shaft, iii the displacements and rotations of the disk are small, iv the center of gravity of the disk finds itself in the middle plane of the disk and is slightly shifted from the shaft center line, and v the middle plane of the disk is perpendicular to the center line of the shaft.\nFrom the point of view of the lateral vibration, the disk has four degrees of freedom. The current position of the disk is defined by the position of point H situated in the intersection of the middle plane of the disk and the center line of the shaft and by its rotations of about two axes perpendicular to the shaft center line. To describe the movement of the disk, three frames of reference are adopted Figs. 1 and 2 .\nThe fixed coordinate system Oxyz is introduced so that axis x could be identical with the shaft\u2019s undeformed center line. Origin O lies in the middle plane of the disk and in the undeformed state, it is identical with point H. Axes y and z go in origin O and are perpendicular to axis x.\nThe coordinate system Tx y z is chosen in such a way that its origin T is situated in the disk center of mass gravity . It moves together with the disk but its axes x , y , and z remain parallel with axes x, y, and z of the fixed frame of reference Tx y z performs only a sliding motion .\nThe third coordinate system T has its origin in the center of gravity of the disk and axes , , and are the disk principal axes of inertia. This system slides together with the center of gravity and rotates about axes y and z so that and still remain to lying in the middle plane of the disk. These angles of rotation are equal to the rotation angles of the disk middle plane about coordinate axes y and z. The direction of axis is close to axis x and as the disk is axisymmetric, and remain the principal axes of inertia at each rotation of the disk.\nAssuming small deformations, the disk performs a motion composed of two components. The first one is the sliding motion given by the movement of the disk center of gravity T. The second component is the spherical motion about the center of gravity.\nThe position of the disk is then defined by displacements of the disk center of gravity yT and zT, by two Euler\u2019s angles 1 and 2 Fig. 3 , and by angle T Fig. 2 . The disk turns first about axis y by a small angle 1 and then about axis the turned axis z by angle 2. T is the angle of rotation of the disk about its axis of\nTransactions of the ASME\n6 Terms of Use: http://www.asme.org/about-asme/terms-of-use", "s r\nm F h\nw d m h\nr\nF a\nJ\nDownloaded Fr\nymmetry . The geometric parameters yT, zT, 1, 2, and T epresent the generalized coordinates of the disk position.\nBecause of the shaft deflection, the center of gravity of the disk oves in the direction perpendicular to the center line of the shaft. or its displacements, yT and zT in the fixed coordinate system it olds Fig. 2\nyT = y + eT cos T, zT = z + eT sin T 1\nhere y and z are displacements of the disk center H in the coorinate system Oxyz, eT is the eccentricity of the disk center of ass center of gravity , and T is the directional angle of the alf-line connecting the disk center H with the center of mass T. Let the angular velocity and angular acceleration of the rotor otation be noted as\nig. 2 Frames of reference of the disk and definition of the ngular position of the rotor\nournal of Applied Mechanics\nom: http://appliedmechanics.asmedigitalcollection.asme.org/ on 01/28/201\n= \u0307T, = \u0308T 2\nDots \u02d9 , \u00a8 denote the first and second differentiation with respect to time in relation 2 and in all the following ones.\nUsing Lagrange equations of the second kind, the equations of motion of the disk have the form\nd dt WK y\u0307T \u2212 WK yT = QTy 3\nd dt WK z\u0307T \u2212 WK zT = QTz 4\nd dt WK\n\u03071 \u2212\nWK\n1 = Q 1 5\nd dt WK\n\u03072 \u2212\nWK\n2 = Q 2 6\nd dt WK \u0307T \u2212 WK T = Q T 7\nwhere WK is the kinetic energy of the disk, QTy, QTz, Q 1, Q 2, and Q T are the generalized forces corresponding to generalized coordinates yT, zT, 1, 2, and T, respectively, and t is time.\nNegatives of expressions on the left-hand side of Eqs. 3 \u2013 7 represent generalized disk inertia forces and moments FSyT, FSzT, MS 1, MS 2, and MS T\nFSyT = \u2212 d dt WK y\u0307T + WK yT 8\nFSzT = \u2212 d dt WK z\u0307T + WK zT 9\nMS 1 = \u2212 d dt WK\n\u03071 +\nWK\n1 10\nMS 2 = \u2212 d dt WK \u02d9 + WK 2 11\n2\nMARCH 2011, Vol. 78 / 021001-3\n6 Terms of Use: http://www.asme.org/about-asme/terms-of-use", "M r\nw m d r\nb s t t\nT c t o r\nF x\n0\nDownloaded Fr\nMS T = \u2212 d dt WK \u0307T + WK T 12\naking use of Fig. 3 and because of small deformations, the elation for kinetic energy of the disk can be expressed as 14\nWK = 1 2mD y\u0307T 2 + z\u0307T 2 + 1 2JP + \u03071 2 2 + 1 2JD\u03071 2 + 1 2JD\u03072 2 13\nhere mD is the disk mass, JP and JD are the disk principal moents of inertia related to its axis of symmetry axis and to its\niameter, respectively, and is the angular velocity of the disk otation.\nAfter performing differentiations of the kinetic energy required y Lagrange equations of the second kind and after neglecting the mall terms of higher orders, the relationships for components of he inertia forces and moments Eqs. 8 \u2013 12 can be rewritten into he following forms:\nFSyT = \u2212 mDy\u0308T 14\nFSzT = \u2212 mDz\u0308T 15\nMS T = \u2212 JP\u0307 16\nMS 1 = \u2212 JD\u03081 \u2212 JP\u03072 \u2212 \u0307JP 2 17\nMS 2 = \u2212 JD\u03082 + JP\u03071 18\nhe generalized coordinates 1, 2, and T and corresponding omponents of the inertia moment acting on the disk are referred o axes y , , and , which are not orthogonal Fig. 4 . Projections f the component of inertia moment MS 1 T acting in plane x y or into axes y , , and x are denoted MS 1, MS T, and MSx\n, espectively Fig. 5 . Using Fig. 6, it holds\nig. 5 Decomposition of the inertia moment into the y , , and\ndirections\n21001-4 / Vol. 78, MARCH 2011\nom: http://appliedmechanics.asmedigitalcollection.asme.org/ on 01/28/201\nMSx = MS T \u2212 MS 1 \u2212 MS T 2 2 19\nAs evident from Fig. 4, y and z components of the disk inertia moment can be expressed as\nMSy = MS 1 20\nMSz = MS 2 \u2212 MSx 1 21\nAfter substitution of Eq. 19 into Eq. 21 and taking into account small rotations, one obtains\nMSz = MS 2 \u2212 MS T 1 22\nFrom the physical point of view QTy, QTz, Q 1, and Q 2 represent the applied and constraint forces and moments acting on the liberated disk. Substitution of Eqs. 17 and 18 into Eqs. 20 and 22 , Eqs. 14 , 15 , and utilization of Eq. 1 enable setting up the motion equations of the disk. Taking into account that 1 is equal to the rotation of the disk about axis y angle y and that because of small deformations angle 2 is equal to the rotation of the disk about axis z angle z , the equations of motion of the disk take the forms\nmDy\u0308 = mDeT 2 cos T + mDeT sin T + Fcy 23\nmDz\u0308 = mDeT 2 sin T \u2212 mDeT cos T + Fcz 24\nJD\u0308y + JP \u0307z + JP z = MOCy 25\nJD\u0308z \u2212 JP \u0307y \u2212 JP y = MOCz 26\nwhere Fcy, Fcz, MOCy and MOCz are the sum of all applied forces and moments acting on the disk and of all constraint ones by which the shaft affects the disk. Because of small deformations, these moments of forces can be related to point O that is situated on the undeformed center line of the shaft.\nAssuming that damping of the disk is of a viscous kind, the equations of motion of the disk Eqs. 23 \u2013 26 can be rewritten into a matrix form\nMDEx\u0308DE + BDE + GDE x\u0307DE + GDExDE = fDE 27\nwhere MDE, GDE, and BDE are the mass, gyroscopic, and external damping matrices of the disk, fDE is the vector of general forces constraint and applied acting on the liberated disk, and x\u0307DE and x\u0308DE are vectors of the disk general velocities and accelerations.\n3 The Motion Equation of a Shaft Element The shaft finite Fig. 7 element has two nodes and eight degrees of freedom two displacements in y and z directions, two rotations about axes y and z at each node . It can be deformed by deflection in two mutually perpendicular planes and in addition, it\nturns about its center line. In the following text, the geometric and\nTransactions of the ASME\n6 Terms of Use: http://www.asme.org/about-asme/terms-of-use" ] }, { "image_filename": "designv11_20_0002288_j.mechmachtheory.2008.03.007-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002288_j.mechmachtheory.2008.03.007-Figure1-1.png", "caption": "Fig. 1. Schematic of the Gleason UMC cradle-style machine in the basic configuration.", "texts": [ " The layout chosen here is just an interesting paradigm for the proposed approach; however, the methodology presented can be easily adapted to describe any process and machine configuration. 3.1. Motion of the tool with respect to the fixed frame In the Euclidean space E3, let us introduce the reference frame S \u00bc \u00f0Os; \u00bds1 s2 s3 \u00de, with si 2 U \u00f0i \u00bc 1;2;3\u00de, fixed with the machine frame. Moreover, let T \u00bc \u00f0Ot ; \u00bdt1 t2 t3 \u00de be the reference frame attached to the moving tool surface Rt . The kinematic structure of the upper body of the cutting machine, which accounts for the motion of the tool surface Rt with respect to the machine frame S, is a 4-dof RPRR serial chain, as in Fig. 1. We identify the configuration represented in Fig. 1 as the basic configuration, which is characterized by the coplanarity of the second, third and fourth axes of the machine, and unit vector t3 of the tool frame. Due to the conventions adopted by the Gleason model, this does not coincide with the reference configuration, because it differs from it for the orientation of the third and fourth axes. Nonetheless, it is introduced because it allows a more systematic definition of the reference configuration itself and an easier derivation of the components of the twist coordinates, which could be quite complex to express by inspection in the reference configuration. Let n0 ti be the components of twists as shown in Fig. 1 in frame S in the basic configuration, and let hti \u00f0u\u00de \u00f0i \u00bc 1; . . . ;4\u00de be the joint motion functions, casted as ht\u00f0u\u00de \u00bc \u00f0ht1 \u00f0u\u00de; . . . ; ht4 \u00f0u\u00de\u00de, all depending on a common parameter of motion, denoted as u. Therefore ht : R! Qt , where Q t is the configuration space of the joints controlling the motion of the tool. The common feature of the joint motion functions employed in the product of exponentials formula (23), is that they must be zero for a zero value of u, that is hti \u00f00\u00de \u00bc 0. Since the reference configuration, depicted in Fig. 2, differs from the basic configuration (Fig. 1) by a constant posture, we can take it into account by introducing the (constant) joint offset values h0 i . Due to the conventions adopted by Gleason, and the basic configuration chosen, the non-zero joint offset values are only h0 t3 and h0 t4 : this are the constant components of the swivel and tilt motions, respectively. From this, it follows imme- diately that only the coordinates of the last twist in the basic configuration, n0 t4 , differ from the correspondent ones in the reference configuration, nt4 ", " Motion of the gear blank with respect to the fixed frame In the Euclidean space E3, besides the fixed frame S \u00bc \u00f0Os; \u00bds1 s2 s3 \u00de and the (moving) tool frame T \u00bc \u00f0Ot; \u00bdt1 t2 t3 \u00de, let us introduce a (moving) frame G \u00bc \u00f0Og ; \u00bdg1 g2 g3 \u00de, fixed with the gear blank G, the workpiece where the teeth are being cut. The kinematic structure of the lower body of the cutting machine, which is responsible for the motion of the gear blank G with respect to the machine frame S, is 5-dof PPRPR serial chain. The configuration depicted in Fig. 1 is the chosen reference configuration for the lower machine body, in agreement, this time, with the convention adopted by Gleason. Therefore, the joint motion functions, denoted as hg\u00f0u\u00de \u00bc \u00f0hg1 \u00f0u\u00de; . . . ; hg5 \u00f0u\u00de\u00de, are zero (for u \u00bc 0) in the above configuration and, now, no intermediate constant transformation has to be introduced. The needed components in S of the twists ngi \u2019s \u00f0i \u00bc 1; . . . ;5\u00de, as well as the initial (constant) posture gsg\u00f00\u00de, are those in the reference configuration of Fig. 1: due to our convenient choice, these can be written down easily \u2018\u2018by inspection\u201d, as we shall see later in Sections 6 and 7. The forward kinematic map gsg : R! SE\u00f03\u00de of the gear blank with respect to the fixed frame is then gsg\u00f0hg\u00de \u00bc en\u0302g1 hg1 en\u0302g5 hg5 gsg\u00f00\u00de: \u00f064\u00de Then, the proximal twist bV s sg of the motion of the gear blank G w.r.t. S is given by bV s sg\u00f0hg\u00de \u00bc \u00f0Js sg\u00f0hg\u00dehg;u\u00de^; \u00f065\u00de where, as in (41), the explicit formula for the proximal Jacobian Js sg , is found to be Js sg\u00f0hg\u00de \u00bc \u00bd ng1 n0g5 ; with n0gi \u00bc Ad en\u0302g1 hg1 en\u0302gi 1 hgi 1 ngi : \u00f066\u00de 4" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000636_ejoc.200500656-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000636_ejoc.200500656-Figure3-1.png", "caption": "Figure 3. Immobilization of cobalt porphyrin on a gold surface via sulfur tails.", "texts": [ " the growth stops after an initial layer. The molecular coverage has been cal- J.-E. B\u00e4ckvall et al.FULL PAPER Scheme 6. Reagents and conditions: l. HCl(aq.), DMF, room temp. culated using the XPS data. The result corresponds to the formation of a complete monolayer. These findings suggest that a complete monolayer of porphyrins is formed in analogy to the results obtained for self-assembled monolayers (SAM) on Au surfaces.[12a] www.eurjoc.org \u00a9 2006 Wiley-VCH Verlag GmbH & Co. KGaA, Weinheim Eur. J. Org. Chem. 2006, 1193\u201311991196 Figure 3 illustrate the immobilization of the cobalt porphyrin 13-Co (deprotected Co-2) to the gold surface. We have successfully prepared 1-Co and 2-Co in relatively good overall yields, by acid-catalyzed condensation of the corresponding S-acetyl-functionalized benzaldehydes and pyrrole followed by direct metallation of the porphyrins. Moreover, an alternative route is presented for 2-Co via the derivatization of the corresponding bromo-functionalized porphyrin 12. After deprotection of the cobalt() porphyrin 2-Co a complete porphyrin monolayer on Au surface was formed, as indicated by the XPS results" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001427_3.60495-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001427_3.60495-Figure3-1.png", "caption": "Fig. 3 The F transfer.", "texts": [ " The optimal transfers are of five different types: 1) Type F: nongrazing transfer with one impulse at a finite distance; 2) Type Foo: nongrazing transfer with a first impulse at a finite distance and a second impulse at an infinite distance from the planet; 3) Type RF: grazing transfer corresponding to type F; 4) Type RFoo: grazing transfer corresponding to type Foo; 5) Type PNP: transfer through the parabolic level of energy, par le niveauparabolique. In the case V2> Vl9 the types Foo, RF, and RFoo become ooF, FR, and ooFR. Transfers of Type F This case is easy to calculate using simple max-min theory; Fig. 3. To simplify the results, the following variables will be introduced: A = arc tan[(Vl-V2/VI+V2)tan E] (4) F=\u00a3-A, G = E + A (5) where F and G determine the direction to the point 7 where the optimal impulse is applied. Using these variables, the distance from the center of attraction is given by 01= [R L2sin2F/V2cos2A[2 cos2\u00a3-cos2A] (6) and the transfer is completely determined. The characteristic velocity of the transfer is C=(V1+V2) sin A (7) = 45\u00b0 -A/4 (3) The figure made at / by the velocities of arrival (K'7) and departure (V2) for the optimal F transfer has many geometrical properties which will be very useful for the other types of transfers; Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003951_j.apenergy.2012.03.026-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003951_j.apenergy.2012.03.026-Figure4-1.png", "caption": "Fig. 4. Experiment framework of mixer connected to MFC.", "texts": [ " In order to study the effect of biometric mixer and biometric flow channel, two experiments were set up, as described as follows: (1) Biometric flow channel experiment. Based on the common continuous MFC, the biometric flow channel was added into the anode chamber, and the schematic diagram was shown in Fig. 3. When the nutrient and bacteria flow through the MFC, the existing block could change the movement of fluid and cause a disturbance. Then the disturbance could improve the mixing of nutrient and bacteria. (2) Biometric mixer experiment. As shown in Fig. 4, the biometric mixer was added ahead of the MFC. Before the nutrient and bacteria affluxes to the MFC, they would mix in the biometric mixer first. It can make the mixing more uniformly. The related experiment operating parameters were as following: In the anode, the E. coli were adopted as the bacteria. The bacteria culture was cultivated by quantitative method. After the cultivation, the OD value would be measured and ensured between 1.8 and 1.9. In the cathode, potassium ferricyanide was used as reducer, with a concentration of 0", " The simulation was conducted with the analysis settings below: (1) Flow channel simulation Total grid amount is N = 7.3 106, Solver is set as AMG (Algebraic Multi-Grid), and the convergent value as 10 4. The fluid flows into the MFC through the upper inlet, with the Reynolds number as Re = 3.73, and flow out from the lower part (as shown in Fig. 3). (2) Flow channel combining biometric mixer simulation The total grid number is N = 4.7 105, Solver is set as AMG, and the convergent value as 10 4. As shown in Fig. 4, the fluid flows into MFC through the upper three inlets with the Reynolds number ratio Rer = 1, and out from the lower outlet. In the fluid mixing simulation, water \u2018\u2018A\u2019\u2019 is supposed to flow through the major inlet in the middle, while water \u2018\u2018B\u2019\u2019 is supposed to flow through the other two inlets. In addition, in the experiment, gray scale technique was employed to quantize the fluid mixing effect through the color changes caused by different fluid density. In the study, in order to analyze the biometric channel\u2019s mixing efficiency, a mixing indicator emixing is defined as shown in Eq", " Compared with them, it could be obviously found that the fluid of (b) had more opportunities to mix together than that of (a). Adding channel in the anode can make full use of the fluid flow, the fluid can be mixed well when they run to the blocks existing in the channel, and then the homogeneity of the fluid can be highly enhanced. As for MFC, the concentration polarization can be effectively improved; as a result, the power performance can be highly increased. (2) Biometric mixer experiment Based on the above model, the biometric mixer was added as shown in Fig. 4, and the role of mixer was discussed in the study. The results of the experiment were shown in Fig. 8, from which the active part of biometric mixer played in MFC can be seen. The maximum power density of MFC with biometric mixer reached 118.34 mW/m3, which increased by 28.9% compared with that of MFC without mixer. Simulations on model combining biometric channel and biometric mixer have been carried on, and the result was as displayed in Fig. 8. After external biometric mixer connected to MFC, the fluid mixing efficiency greatly increased from mixer inlet of emixing = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003397_s12289-010-0686-3-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003397_s12289-010-0686-3-Figure2-1.png", "caption": "Fig. 2 a Geometry and b meshed axisymmetric model of the tube\u2013tubesheet joint with roller", "texts": [ " The equivalent sleeve is a single hole model that will produce contact pressure, stress distribution or deflection, depending on the objective of the study, around the hole equivalent to the average of those around the test hole on the real tubesheet configuration. Since this study uses the same configuration of the stabilizer feed/bottom exchanger reported in an earlier work by Merah et al. [10], Shuaib et al. [16], and Al-Aboodi et al. [17], the sleeve dimensions will remain unchanged. Using the geometry in Fig. 2(a), the tube inner and outer radii are 7.425 mm, 9.525 mm respectively, and the tubesheet inner and outer radii are 7.425 + c mm and 36 mm respectively, where c is the radial clearance, which will be varied here from 0 to 0.5 mm. It should be mentioned that for this tube dimension the TEMA (Tubular Exchanger Manufacturing Association) allowable radial clearance is about 0.16 mm [18]. The tube and tubesheet areas are meshed using the ANSYS 2-D VISCO108 element as shown in Fig. 2(b) [19]. This is a quadratic element defined by eight nodes having two degrees of freedom at each node: translations in the nodal x and y directions. From its special features; it has a rate-dependent plasticity, stress stiffening, large deflection, large strain, and adaptive descent. This element could be used with axisymmetric or plane strain problems. Because of expected large deflections due to large over tolerances used in this study NLGEOM, which accounts for geometric nonlinearities is activated in order to update the geometry at each sub step", " An elastic-perfectlyplastic material is that having zero tangent modulus. The approximate value of the tangential modulus of plasticity, Ett, for the tube material used in the current stabilizer feed/ bottom exchanger was 733 MPa (106.4 ksi). However, to investigate the effect of tube and tubesheet material strain hardening on contact stresses, Ett values ranging from 0 to 1.2 GPa will be considered. This range covers most of steel materials in heat exchanger tubes and tubesheets. The roller profile in Fig. 2(a) is represented as a rigid body line. The tube and tubesheet were constrained from translation in the axial direction on the primary side as shown in Fig. 2(b). Loading is performed in three steps. The first one consists of displacing the roller radially in small increments until contact is established between the tube and tubesheet. The targeted percent wall reduction (5% in this study) is reached by performing 50 sub steps in the second loading step. The third step simulates the retraction of the roller after achieving the required expansion. The expanded length, which is equal to the roller length, shown in Fig. 2(a), was 47.25 mm (1.872 in.); this represents about 75% of the tubesheet thickness as practiced in the industry. The total displacement load was specified with the knowledge of the required percentage wall reduction (5%WR), tube thickness, t, and initial clearance, c, using the following equation: ur \u00bc 5:0:t 100 \u00fe c \u00f02\u00de Figure 3 shows the variation of the residual contact stress with tube\u2013tubesheet initial clearance for different values of friction factors. The percent wall reduction is taken as 5% and the tube and tubesheet materials are assumed to be elastic-perfectly plastic (Ett=0) for this case" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002815_1.4002527-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002815_1.4002527-Figure9-1.png", "caption": "Fig. 9 Scheme of the investigated rotor system", "texts": [ " 68 where DB,E and DK,E are the Jacobi matrices of partial derivatives, DB,E = fB x\u0307 x=xE,x\u0307=o , DK,E = fB x x=xE,x\u0307=o 69 The calculation of the system natural frequencies then arrives at solving the eigenvalue problem det O M M B + VKSH + G \u2212 DB,E + \u2212 M O O K + KC \u2212 DK,E = 0 70 Equations 67 and 70 hold only for the operating conditions when the rotor rotates at a constant angular speed and no contact between the disks and the stationary part takes place. 7 Example Rotor of the investigated rotor system Fig. 9 consists of a shaft SH and of three disks D1, D2, and D3 attached to it. Disks D1 and D3 are included in rectangular holes in the station- MARCH 2011, Vol. 78 / 021001-7 6 Terms of Use: http://www.asme.org/about-asme/terms-of-use a i d m i a s a r l d i a i s b c a f f r 4 a t b s t t r O t o c r i f n r t d m t m l s 0 Downloaded Fr ry part and the gaps between them and the housing are small. The nterior surface of the holes is planar; the outer surface of each isk has a spherical form. Both the disks and the housing are ade of steel" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001460_1.2713788-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001460_1.2713788-Figure7-1.png", "caption": "Fig. 7. A ball impacting on a baseball bat can result in bat rotation about two different axes, as shown in a and b .", "texts": [ " Of greater significance in the second experiment is the fact that the gear effect has a strong influence on the ball\u2019s rebound spin, speed, and angle for impacts resulting in block rotation, consistent with the models described in Sec. IV and in the Appendix. Although the gear effect is known to be important in the design of golf clubs, it has not previously been examined in relation to baseball collisions. There are two circumstances where the gear effect might play a role. For the situation shown in Fig. 7 a we can apply the formalism of the Appendix. We consider the collision of a baseball m=5.1 oz with two bats, one wood and one aluminum, each having the shape of an R161 wood bat, a length of 34 in., and a weight 662 662Am. J. Phys., Vol. 75, No. 7, July 2007 R. Cross and A. M. Nathan This article is copyrighted as indicated in the article. Reuse of AAPT content is subject to the terms at: http://scitation.aip.org/termsconditions. Downloaded to IP: 202.28.191.34 On: Thu, 26 Feb 2015 20:48:59 of 31", " If we transform to the bat rest frame, vy1=160 mph. For the wood bat we find vy2=37 mph, vx2=0.78 mph, and =226 rpm. The ball leaves the bat at an angle of 1.2\u00b0. If we transform back to the original frame, the ball-bat scattering angle is only 0.4\u00b0. For the aluminum bat we find vy2=30 mph, vx2=0.95 mph, =276 rpm, and a ball-bat scattering angle of 0.5\u00b0. We conclude that the gear effect results in a negligible scattering angle and a modest spin that is larger for aluminum than for wood. The situation shown in Fig. 7 b was considered previously,5 but can be regarded as being equivalent to an impact with b=0 in the second experiment. We find from Eq. A15 that the ball spin depends on the moment of inertia of the block or the bat, even when b=0. The spin is enhanced for a bat with a large moment of inertia, but the effect is not due to the gear effect because there is no tangential acceleration of the surface during the grip phase either by the friction force which is zero if the ball rolls with ex=0 or by the normal reaction force which acts through the bat center of mass ", " We further note that our expressions are equivalent to those derived by Penner1 if the rolling condition is assumed ex=0 , once transformed to the frame in which the ball is initially at rest. a Electronic mail: cross@physics.usyd.edu.au b Electronic mail: a-nathan@uiuc.edu 1A. R. Penner, \u201cThe physics of golf: The convex face of a driver,\u201d Am. J. Phys. 69, 1073\u20131081 2001 . 2R. Cross, \u201cGrip-slip behavior of a bouncing ball,\u201d Am. J. Phys. 70, 1093\u20131102 2002 . 3Baseball bats, golf clubs, and tennis racquets all differ in this respect. For an off-axis impact, N can act through the center of mass of a bat as in Fig. 7 b , but it does not normally act through the center of mass of a club or racquet. For a tennis racquet, the center of mass lies in the plane of the impact surface, and the center of mass of a bat or a club lies behind the plane of the impact surface. 4 www.vpfundamentals.com . 5R. Cross and A. Nathan, \u201cScattering of a baseball by a bat,\u201d Am. J. Phys. 74, 896\u2013904 2006 . 6A ball impacting on a soft surface such as rubber would cause the surface to accelerate locally in the contact region due to the finite tangential stiffness of the surface" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000262_tmag.2004.824897-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000262_tmag.2004.824897-Figure2-1.png", "caption": "Fig. 2. Analyzed model (1/2 region).", "texts": [ " The hysteresis loss taking into account the major and minor loops of the hysteresis loop can be estimated as follows [5]: (5) where is the coefficient of the hysteresis loss, is the density of the steel sheet, is the period of analysis time, is the number of the elements in the steel sheet, and is the volume of the th element. , , and are the numbers of the maximum or minimum values of the flux density of the radial direction, the rotation direction and the direction of the th element, respectively. , , and are the amplitude of the flux density of major and minor hysteresis loops of the radial direction, the rotation direction, and the direction, respectively. Fig. 2 shows the analyzed model of a squirrel-cage induction motor, the rotor of which is skewed with one rotor slot pitch. It is 1/2 of the whole region because of the periodicity. Fig. 3 TABLE I ANALYZED CONDITIONS Fig. 4. Distributions of flux density vectors (0 rpm). (a) Upper section (z = 41 mm). (b) Lower section (z = 1 mm). shows the 3-D finite element mesh. Table I shows the analyzed conditions. Fig. 4 shows the distributions of the flux density vectors. It is found that the flux density vectors in the upper section are larger than those in the lower section due to the skew" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003863_b978-0-12-417049-0.00002-x-Figure2.11-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003863_b978-0-12-417049-0.00002-x-Figure2.11-1.png", "caption": "Figure 2.11 Geometrical structure of the N-trailer WMR.", "texts": [ "66) is converted to the (2,4)-chain form: _x1 5 u1 _x2 5 u2 _x3 5 x2u1 _x4 5 x3u1 \u00f02:69\u00de This is an extension of the car-like WMR, where N one-axis trailers are attached to a car-like robot with rear-wheel drive. This type of trailer is used, for example, at airports for transporting luggage. The form of equations depend crucially on the exact point at which the trailer is attached and on the choice of body frames. Here, for simplicity each trailer will be assumed to be connected to the axle midpoint of the previous trailer (zero hooking) as shown in Figure 2.11 [20]. The new parameter introduced here is the distance from the center of the back axle of trailer i to the point at which is hitched to the next body. This is called the hitch (or hinge-to-hinge) length denoted by Li. The car length is D. Let \u03c6i be the orientation of the ith trailer, expressed with respect to the world coordinate frame. Then from the geometry of Figure 2.11 we get the following equations: xi 5 xQ 2 Xi j51 Lj cos \u03c6j i5 1; 2; . . .;N yi 5 yQ 2 Xi j51 Lj sin \u03c6j which give the following nonholonomic constraints: _xQ sin \u03c60 2 _yQ cos \u03c60 5 0 _xQ sin\u00f0\u03c60 1\u03c8\u00de2 _yQ cos\u00f0\u03c60 1\u03c8\u00de2 _\u03c60D cos \u03c85 0 _xQ sin \u03c6i 2 _yQ cos \u03c6i 1 Xi j51 _\u03c6jLj cos\u00f0\u03c6i 2\u03c6j\u00de5 0 for i5 1; 2; . . .;N. In analogy to Eq. (2.52) the kinematic equations of the N-trailer are found to be: _xQ 5 v1 cos \u03c60 _yQ 5 v1 sin \u03c60 _\u03c60 5 \u00f01=D\u00dev1tg\u03c8 _\u03c85 v2 _\u03c61 5 1 L1 sin\u00f0\u03c60 2\u03c61\u00de _\u03c62 5 1 L2 cos\u00f0\u03c60 2\u03c61\u00desin\u00f0\u03c61 2\u03c62\u00de ^ _\u03c6i 5 1 Li L i21 j51 cos\u00f0\u03c6j21 2\u03c6j\u00desin\u00f0\u03c6i21 2\u03c6i\u00de ^ _\u03c6N 5 1 LN L N21 j51 cos\u00f0\u03c6j21 2\u03c6j\u00desin\u00f0\u03c6N21 2\u03c6N\u00de \u00f02:70\u00de which, obviously, represent a driftless affine system with two inputs u1 5 v1 and u2 5 v2 and N1 4, states: _x5 g1\u00f0x\u00deu1 1 g2\u00f0x\u00deu2 We observe that the first four lines of the fields g1 and g2 represent the (pow- ered) car-like WMR itself" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002271_s11434-008-0370-x-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002271_s11434-008-0370-x-Figure1-1.png", "caption": "Figure 1 The concept of a new PZT on/off valve.", "texts": [ " Therefore the displacement amplification of the piezoelectric actuator should be figured out firstly. At the moment some methods such as a lever, flexible hinge and fixed hydraulic volume are employed to amplify the displacement of the PZT actuator[6], which do realize the function of amplifying the displacement, but at the same time also reduce its output force, response and control precision, which implies that these methods are not effective. An innovative concept of a PZT digital valve based on fast-response and high output force of the PZT actuator is outlined (Figure 1). The high pressure oil flows from port 1 to port 2. The movement resistance (including flow force and pressure force) of the poppet is maximal at the start moment of the open or close of the poppet, and reduces quickly with the poppet moving (Figure 2). Three PZT actuators are used to open and close the poppet in the new digital valve. When the two PZT1 ac- Figure 2 Resistant force/displacement of the poppet. tuators are powered on, they output instant force of above 3 kN to kick the poppet to accelerate running, once the poppet opens, the pressure p2 increases rapidly, the resistant force of the poppet drops quickly, and the poppet keeps moving at the spring force of Fs1 until its end", " On the contrary, when the PZT2 actuator is powered on, it produces high instant force of 3 kN to impact the poppet to move, until its another stroke ends. Here it is the high instant force of the PZT actuator that contributes to the movement of the poppet, and therefore the problem of micro displacement amplification of a PZT actuator is successfully avoided. In this case, the stroke of the poppet can reach up to 1 mm, accordingly the high flow rate can be gained and the conflict among high flowrate, pressure and quick response is also solved. The following mathematical model of the PZT valve is obtained from Figure 1. Assuming the mechanical and viscous friction force is zero, when the PZT1 actuators are powered on, the force equation is expressed as follows: PZT1 1 2 s1 s2 ,f rF F F F F F F mx\u2212 + \u2212 + \u2212 + = (1) where F1 force is produced by pressure p1: 2 2 1 1 1 \u03c0 ( ) ; 4 F d d p= \u2212 F2 force is produced by pressure p2: 2 2 2 \u03c0 ; 4 F d p= Ff flow force is ,2 cos ;f vF q v\u03c1 \u03b1= Fsi spring force is siF = s 0( )i i iK x x+ \u0394 , i = 1, 2; Fr damper force of rubber is 1 2 :r r rF F F= \u2212 + 1 1(rF k h= \u2212 )hx + rx 1(0 ),hx h\u2264 \u2264 2 2(rF k h= \u2212 )x rx+ (x\u2264 2 )h " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002539_icma.2007.4303740-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002539_icma.2007.4303740-Figure7-1.png", "caption": "Fig. 7. Model for Cruise-in-turning", "texts": [ " The 1st phase of first joint, \u03d51,1, is set to zero without losing generality. So the form of the joint angle to be controlled can be written as follows: qj(t) = aj sin(\u03c9t + \u03d5j), j = 1...K (5) When a fish is performing Cruise-in-turning, its body motion function can be modelled by the sum of the body motion function of cruise straight and a deflected centre curve, denoted as d(x). So the undulation centre of the tail movement becomes a curve rather than a straight line as in cruise straight, as shown in Fig. 7. However, nobody has shown the form of d(x). Based on our observation of fish swimming, we assume: d(x) = d1x d2 , d2 > 1, x > 0 d1 { > 0 cruise-in-right-turn < 0 cruise-in-left-turn (6) where d2 is the curvature factor to control the curvature of the deflected centre and d1 is the direction factor to control the turning direction, which is also proportional to the turning angle when d2 is fixed. Therefore, the body motion function of cruise-in-turning can be written as follows: y = fB(x, t) = (c1x + c2x 2) sin(kx + \u03c9t) + d1x d2 (7) where all the parameters have a similar definition to Equation (1)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000139_0020-7403(86)90032-9-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000139_0020-7403(86)90032-9-Figure4-1.png", "caption": "FIG. 4. Magnus' apparatus (a) for demonstrating effect of pressure on circulation of vanes (or flags) and (b) torsion balance.", "texts": [ " 3(c); at A, velocities u and v reinforce and oppositely at B, they oppose. Magnus next remarks that where the velocities reinforce a stronger pressure towards the centre occurs than where they detract, so that it may be expected that any deviation would be towards this opposite side. However, experience proves to be contrariwise. He thus felt it necessary to investigate experimentally actual existing pressures. To do this and in order to facilitate 'greater certainty of observation' he utilized a cylinder instead of a sphere; Fig. 4(a) * von Heim is reported [4, p. 309] as having conducted experiments with excentric shells, shows the actual figure from Magnus' paper. A small centrifugal fan was usually employed and in the 6 in. drum of the fan, F, a shaft rotated to which six fixed paddle-shaped fans were attached. The drum of air was uniformly impelled and forced out through orifice ran. Small movable vanes were made to serve as detectors of pressure changes in the air current during the rotation of the cylinder. Vanes such as a and b on each side of the cylinder turned on pivots equally distant from mn at equal distances from the plane passing through the axis of rotation of the cylinder and the middle of the current", " 3(b), according as rotational and progressive motions are combined to create appropriately enhanced or reduced pressures, in conformity with his developed earlier understandings or explanations. Finally, Magnus endeavours to show experimentally that the difference of pressure across a sphere is indeed great enough to deflect a projectile from its course. A light hollow brass cylinder 3 in. long and 2 in. in diameter, ab, turning on two pivots was fixed in a metal ring; this latter was itself attached to the end of a light wooden beam yz, 4 ft long, suspended in its centre by a fine wire vw and thus forming a kind of torsion balance. The beam had a counterweight system; see Fig. 4(b). The pulley, at E, could be rotated by a string wrapped around it, ~o that an air current from the small centrifugal fan, F, issued through mn against ab. The fan, on the board AB which could turn on a vertical axis C, was always able to tollow the current and act parallel to it. Fig. 4(b) indicates one resulting motion 5, for a given air current direction ~ and a cylinder rotation b. Magnus' closing statement after further discussion of this last experimental set-up leads him to write that considering the 'great velocity of the periphery of a rotating projectile in comparison to that of the cylinder, it cannot be doubted that the difference of the atmospheric pressures against the opposite sides of such a projectile is sufficient to cause the observed lateral deviation'. To all the experiments described above, no quantitative measurements are attached: qualitative description only obtains as with reported work by earlier researchers" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001196_1.2346690-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001196_1.2346690-Figure5-1.png", "caption": "Fig. 5 Benckert and Wachter\u2019s t", "texts": [ " This modification to the straight through labyrinth seal created five rows of four cavities. Their configuration was similar to a pocket damper seal described below , but differed in that there were no notches in any of the blades. Their seal contained 4 circumferential pockets, whereas the typical pocket damper seal PDS designs have eight circumferential pockets. In both cases the pocket partition walls block the swirl of the gas and so would be expected to reduce the de-stabilizing CCS. The test rig they used and the seal tested are shown in Fig. 5. Static pressure probes were installed at 30 deg angle locations around the periphery of the test seal chambers and were used to determine the resultant force on the seal housing. The test rotor was overhung and was supported by two back to back mounted angular contact ball bearings and one cylindrical roller bearing. Also, Benckert and Wachter 4 used pre-swirl rings to impose gas swirl into the inlet chamber of the seal. The rotor was driven by a variable speed drive and the test rig had the ability to offset displace the rotating assembly relative to the housing in one direction. This is illustrated in Fig. 5, where the relative eccentricity ing seals: \u201ea\u2026 centered rotor and \u201eb\u2026 offset erg FEBRUARY 2007, Vol. 129 / 29 6 Terms of Use: http://www.asme.org/about-asme/terms-of-use i r fi a r s d s o u o o t C n c b r l b o r p l d p i e A c m p t f p p l c s S s fi m T t 1 T m 4 a s c F P t v p c i n 3 Downloaded Fr s plotted versus the lateral force Q. The right side graph shows esults for a labyrinth seal without the swirl webs and their modied labyrinth with the swirl webs. Their results show that the ddition of the swirl webs increased the lateral force Q, which esolved in the direction of rotor whirl" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001569_2007-01-0136-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001569_2007-01-0136-Figure2-1.png", "caption": "Fig. 2. Schematic diagram of slider crank mechanism identifying the various forces developed", "texts": [ " During transient operation, the thermal inertia of the cylinder wall is taken into account, using a detailed heat transfer scheme that models the temperature distribution from the gas to the cylinder wall up to the coolant (convection from gas to internal wall surface and from external wall surface to coolant, and conduction across the cylinder wall). At each instant of time, the piston displacement from the top dead center (TDC) position is determined by [8] 2 2 rodx( ) r(1 cos ) L 1 1 sin (3) where r is the crank radius, =r/Lrod with Lrod the connecting rod length (see also Fig. 2), and the crank angle is measured from the TDC position. Differentiating the above equation with respect to time, we get the instantaneous piston velocity pist 2 2 cosu ( ) r sin 1 1 sin (4a) while differentiating once again with respect to time, we get the instantaneous piston acceleration 2 4 pist2 2 2 3 / 2 2 ucos2 sin 1b( ) r cos r(1 sin ) (4b) The last term on the right hand side of Eq. (4b) takes into account the crank\u2019s angular acceleration \u2018 \u2019 in the piston acceleration. The loading of the bearings can then be computed from the following equations, with reference to Fig", "r F( )B sin( ) cos F( )B cos( ) m cos r (7) for the crank pin, 3x 2 3y rod.r crank F( )B sin( ) cos F( )B cos( ) (m m ) cos r (8) for the crank journal1, and 2 4x rod.r crank 2 4y rod.r crank B F( ) tan (m m ) r sin B F( ) (m m ) r cos (9) for the main crankshaft bearing. The corresponding total bearing force is then 2 2 i ix iB B B y (10a) and the angle , shown in Fig. 3, is given by 1 ix i iy B tan B (10b) with i = 0\u20264 according to the bearing studied. In Eqs (5-9), corresponds to the connecting rod angle (see also Fig. 2), i.e. 1 2 2cos [ (1 sin )] . The total force acting on the piston is composed of the gas and the inertia force, i.e. , which then propagates into the thrust force and the force in the direction of the connecting rod . The gas force is determined by while the reciprocating masses (inertia) force by , with the reciprocating mass and , i.e. the connecting rod is assumed equivalent to two masses, one reciprocating with the piston assembly and the other rotating with the crank. g lF( ) F ( ) F( ) thrF ( ) F( ) tan rodF ( ) F( ) / cos g g atmF ( ) (p ( ) p ) Apist l lF( ) m b( ) l rod", " However, it induces errors in the slider-crank mechanism dynamics by miscalculating the actual rod\u2019s moment of inertia and the various forces of the kinematic mechanism. For a more accurate computation of engine torque, we have developed a detailed model of the connecting rod based on rigid body dynamics. Here, we analyze the complex, elliptical movement of the rod\u2019s center of gravity that is produced by its reciprocating and rotating motion. A system of 3X3 equations provides the respective inertia force FTin \u2018acting\u2019 perpendicularly to the crank axis (see Fig. 2) that is needed in Eq. (28). Details about this analysis are given in Appendix A. Figure 6 demonstrates the difference in the inertia force results, obtained by using the rigid body and the lumped mass models for an early cycle of a load increase transient. As can be observed, the simplified model generally overestimates the inertia force, particularly at the crank angles where the local maxima or minima occur. A sensitivity analysis carried out using both approaches showed that the detailed connecting rod model resulted in only modest differentiations (of the order of 1%) in the engine speed response predictions, compared to the lumped mass approach" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000759_imece2004-60714-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000759_imece2004-60714-Figure4-1.png", "caption": "Figure 4. Hypothetical actuator strip", "texts": [ "25 J/m2 Concentration of ADP: [ADP ] = 8 mM Concentration of Phosphate: [PO2\u2212 4 ] = 1 mM Concentration of ATP: [ATP ] = 10 mM to 100 mM Diffusion coefficient: D = 5 x 10\u22129 m2/s Partition coefficient: h = 4.2 x 10\u22123 Results indicate that capsule presures from 5 to 15 MPa can be developed by the capsules for 10 mM to 100 mM ATP at pracitally possibly fluid flow rates. Apropos, we are also working on fluid delivery and scavenging scheme in the material matrix. The actuator strip may be visualized as shown in Figure 4. The pressure developed as Copyright c\u00a9 2004 by ASME rl=/data/conferences/imece2004/71571/ on 05/09/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use D a function of ATP flow rate is plotted in Figure 5. Pressure developed varies directly as the concentration and flow rate of ATP fuel. The capsules will develop 10 MPa turgor pressure consuming slightly over 1 g/s of 10 mM ATP. Since hydrolysis of ATP and osmotic regulation in a sheet 200 \u00b5m thin sheet are instantaneous, the frequency and bandwidth of operation mainly depends upon fuel delivery constraint through the material" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003711_pime_conf_1967_182_341_02-Figure29.1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003711_pime_conf_1967_182_341_02-Figure29.1-1.png", "caption": "Fig. 29.1. Generation of vibration by local indentation", "texts": [ " I t must be mentioned that we are not interested in micro-geometrical properties of the surface; the imperfections to be considered should have an extension of at least the same order as the surface of contact between the rings and the rolling bodies. The latter means a length of 50 pm or greater if we do not consider miniature bearings or bearings with an extremely light load. To get an idea of how vibrations are generated, consider the case where the outer ring of a radially loaded deep groove ball bearing has a local indentation in the raceway (Fig. 29.1). Let the depth of the indentation be a few micrometres and let the length be of the same order of magnitude as the ball radius. Moreover, let the indentation be situated in the loaded zone of the bearing, the inner ring of which is in rotation while the outer ring is stationary. When a ball passes over the indentation it gets a displacement which depends on time, and the inner ring will also be displaced as observed from the outer ring, i.e. a motion is set up. The motion of the inner ring is constrained by all of the balls in the loaded zone and also by a set of forces from the axis on which it is mounted" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003619_s11012-010-9331-y-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003619_s11012-010-9331-y-Figure3-1.png", "caption": "Fig. 3 Example of DP minimization and variation", "texts": [ " Transformation matrix [A] established by Suh [1] for the relation in Fig. 2 is inverse matrix [A]\u22121. Numeric values of matrix members define the relations in design component, which are constrained by numerous limitations, such as safety or reliability, stiffness, standards, rules, etc. These limitations and constraints are the result of service conditions, which is deducted by LAHP module to the level of design component. For the purpose to present this relation more clearly, the following example is processed. In Fig. 3 the example is presented. The assembly of the gear, shaft and bearing is defined by a great collection of design parameters, especially dimensions. The calculation of dimensions is reduced to the three dimensions, gear diameter d , gear width b and shaft diameter dhs. In this way, the axiom of information minimum is fulfilled. Other dimensions are in relation with those calculated. Matrix [G] (Fig. 3) is the shape vector which defines transformation of parameters in the all shape dimensions. This is the shape parameterization where varying of the shape parameters varies the complete shape and dimensions. In Fig. 3 are presented the two shapes of the same assembly obtained in this way. Similar approach is incorporated in CAD tools for the shape modeling. The structure of the matrix [A] according to Suh can be uncoupled, coupled and decoupled. The ideal situation is with uncoupled matrix where one DP is responsible for one FR. Real situation is more complex. In order to obtain the decoupled matrix of transformation, the matrix [A]\u22121 is presented in the form of matrix [C] in the following form. \u23a7 \u23aa\u23aa\u23a8 \u23aa\u23aa\u23a9 d dhs C SE \u23ab \u23aa\u23aa\u23ac \u23aa\u23aa\u23ad = \u23a1 \u23a2 \u23a2 \u23a3 c11 0 0 0 0 c22 0 0 0 0 c33 0 0 0 0 c44 \u23a4 \u23a5 \u23a5 \u23a6 \u23a7 \u23aa\u23aa\u23a8 \u23aa\u23aa\u23a9 T 1/3 T T 1 \u23ab \u23aa\u23aa\u23ac \u23aa\u23aa\u23ad " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003863_b978-0-12-417049-0.00002-x-Figure2.14-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003863_b978-0-12-417049-0.00002-x-Figure2.14-1.png", "caption": "Figure 2.14 Four-mecanum-wheel WMR (A) Kinematic geometry (B) A real 4-mecanumwheel WMR. Source: http://www.automotto.com/entry/airtrax-wheels-go-in-any-direction.", "texts": [ "17), that is: _xQ _yQ 5 cos \u03c6 2sin \u03c6 sin \u03c6 cos \u03c6 _xr _yr 5R\u00f0\u03c6\u00de _xr _yr or _xr _yr 5 cos \u03c6 sin \u03c6 2sin \u03c6 cos \u03c6 _xQ _yQ 5R21\u00f0\u03c6\u00de _xQ _yQ Now, we have: v1 5 r _\u03b8 1 5 uT1 _xr _yr 1D _\u03c65 uT1R 21\u00f0\u03c6\u00de _xQ _yQ \" # 1D _\u03c6 v2 5 r _\u03b8 2 5 uT2 _xr _yr 1D _\u03c65 uT2R 21\u00f0\u03c6\u00de _xQ _yQ \" # 1D _\u03c6 v3 5 r _\u03b8 3 5 uT3 _xr _yr 1D _\u03c65 uT3R 21\u00f0\u03c6\u00de _xQ _yQ \" # 1D _\u03c6 v4 5 r _\u03b8 4 5 uT4 _xr _yr 1D _\u03c65 uT4R 21\u00f0\u03c6\u00de _xQ _yQ \" # 1D _\u03c6 or, in compact, form: _q5 J21 _pQ \u00f02:75a\u00de where _q5 _\u03b8 1 _\u03b8 2 _\u03b8 3 _\u03b8 4 2 664 3 775; _pQ 5 _xQ _yQ _\u03c6 2 4 3 5 \u00f02:75b\u00de J21 5 1 r \u00f0UTR21\u00f0\u03c6\u00de1D\u00de \u00f02:75c\u00de with: U5 u1 ^ u2 ^ u3 ^ u4 ; D5 D; D; D; D Th \u00f02:75d\u00de As usual, this inverse Jacobian equation gives the required angular wheel speeds _\u03b8 i\u00f0i5 1; 2; 3; 4\u00de that lead to the desired linear velocity \u00bd _xQ; _yQ , and angular velocity _\u03c6 of the robot. A discussion of the modeling and control problem of a WMR with this structure is provided in Ref. [17]. Consider the 4-wheel WMR of Figure 2.14, where the mecanum wheels have roller angle 6 45 [2,4]. Here, we have four-wheel coordinate frames Oci\u00f0i5 1; 2; 3; 4\u00de. The angular velocity _qi of the wheel i has three components: 1. _\u03b8 ix: rotation speed around the hub 2. _\u03b8 ir: rotation speed of the roller i 3. _\u03b8 iz: rotation speed of the wheel around the contact point. The wheel velocity vector vci 5 _xci; _yci; _\u03c6ci T in Oci coordinates is given by: _xci _yci _\u03c6ci 2 4 3 55 0 ri sin\u03b1i 0 Ri 2ri cos\u03b1i 0 0 0 1 2 4 3 5 _\u03b8 ix _\u03b8 ir _\u03b8 iz 2 4 3 5 \u00f02:76\u00de for i5 1; 2; 3; 4, where Ri is the wheel radius, ri is the roller radius, and ai the roller angle", "13) is: _pQ 5 _xQ _yQ _\u03c6Q 2 4 3 55 cos \u03c6Q ci 2sin \u03c6Q ci dQciy sin \u03c6Q ci cos \u03c6Q ci 2d Q cix 0 0 1 2 4 3 5 _xci _yci _\u03c6ci 2 4 3 5 \u00f02:77\u00de where \u03c6Q ci denotes the rotation angle (orientation) of the frame Oci with respect to QxQyQ, and dQcix, d Q ciy are the translations of Oci with respect to QxQyQ. Introducing Eq. (2.76) into Eq. (2.77) we get: _pQ 5 Ji _qi \u00f0i5 1; 2; 3; 4\u00de \u00f02:78\u00de where _qi 5 _\u03b8 ix; _\u03b8 ir; _\u03b8 iz T , and Ji 5 2Risin \u03c6Q ci risin\u00f0\u03c6Q ci 1\u03b1i\u00de dQciy Ricos \u03c6 Q ci 2ricos\u00f0\u03c6Q ci 1\u03b1i\u00de 2dQcix 0 0 1 2 4 3 5 \u00f02:79\u00de is the Jacobian matrix of wheel i, which is square and invertible. If all wheels are identical (except for the orientation of the rollers), the kinematic parameters of the robot in the configuration shown in Figure 2.14 are: Ri 5R; ri 5 r; \u03c6Q ci 5 0 jdQcixj5 d1; jdQciyj5 d2 \u00f02:80\u00de \u03b11 5 a3 52 45 ; \u03b12 5\u03b14 5 45 Thus, the Jacobian matrices (2.79) are: J1 5 0 2r ffiffiffi 2 p =2 d2 R 2r ffiffiffi 2 p =2 d1 0 0 1 2 64 3 75; J2 5 0 r ffiffiffi 2 p =2 d2 R 2r ffiffiffi 2 p =2 2d1 0 0 1 2 64 3 75 J3 5 0 2r ffiffiffi 2 p =2 2d2 R 2r ffiffiffi 2 p =2 2d1 0 0 1 2 64 3 75; J4 5 0 r ffiffiffi 2 p =2 2d2 R 2r ffiffiffi 2 p =2 d1 0 0 1 2 64 3 75 \u00f02:81\u00de The robot motion is produced by the simultaneous motion of all wheels. In terms of _\u03b8 ix (i" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003320_gt2010-22058-Figure13-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003320_gt2010-22058-Figure13-1.png", "caption": "FIG. 13 SEAL MODELS WITH DIFFERENT INFLOW CAVITY CONFIGURATIONS", "texts": [ " However, applying quadratic rotordynamic model for the rotational speed case of 12000 rpm still results in very small addedmass coefficients and insignificant influence on the stiffness and damping coefficients. Additional studies are needed to analyze in detail possible non-linear effects at high frequencies. Effect of Inflow Cavity Configuration Additionally to the basic inflow cavity configuration shown in Figure 6, two other inflow geometries are also considered to study their influence on rotordynamic coefficients. In the first configuration (Figure 13, left) the flow enters the upstream cavity horizontally through forty inlet nozzles. The second configuration (Figure 13, right) consists of twenty passages with flow entering vertically. Table 3 summarizes predicted results of the effect of inlet cavity geometry on the stiffness and damping coefficients for the reference case (7). The inflow cavity has a great influence on seal rotordynamic coefficients calculated by both static and dynamic methods. The configuration with forty horizontal nozzles shows maximum deviation from the basic configuration with higher absolute direct stiffness and cross-coupled damping, and smaller cross-coupled stiffness" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002203_b615417d-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002203_b615417d-Figure1-1.png", "caption": "Fig. 1 Schematic representations of (a) the tubular flow electrode including the parabolic flow profile, and (b) the microdisc electrode.", "texts": [ " In this paper, steady state concentration profiles at the tubular flow and microdisc electrodes are obtained by numerical simulation, from which the current\u2013potential behaviour may be found. The efficiency of the conformal space developed by OAS in simulating steady state currents is discussed. An account of the non-uniform current response at the tubular flow and microdisc electrodes is given, and their mass transport-corrected Tafel plots are considered, with comparisons drawn between the two geometries. A schematic representation of the tubular flow electrode is shown in Fig. 1a. The numerical methods used to calculate the steady state concentration profiles at a tubular flow cell by solving the mass transport equations have been explained previously.6 For the simple electron transfer reaction in eqn (1), normalised concentration profiles are obtained by solving eqn (11) subject to appropriate boundary conditions: @a @T \u00bc gt0gt \u00bc @ 2a @x2 2\u00f0Ps 1 3 x\u00de @a @w ; \u00f011\u00de where the following dimensionless variables are used: Ps \u00bc 2Vf pDAxe \u00f012\u00de This journal is c the Owner Societies 2007 Phys", " Under these conditions, it has been shown that varying the dimensionless volume flow rate, Ps, has no effect on the value of j or the shape of the simulated wave.6 Hence, a complete set of wave shape data may be obtained by performing the simulations at only one value of Ps. As in the previous work, the mesh over which the differential equations are solved is expanding in x and described by a parameter o. In this work, a value of 100 is chosen for Ps, a value of 15 is used for o, and the required grid size is found to be Nx = 5000 and Nx = 200. A schematic representation of the microdisc electrode is shown in Fig. 1b. The quasi-conformal mapping developed by OAS17 is described by the following set of equations: R \u00bc sinG cosZ \u00f019\u00de Z \u00bc cosG tanZ \u00f020\u00de which represent the transformation from the conformal space in the G, Z-space onto the first quadrant of the normalised real space with R = r/rd and Z = z/rd where r and z represent the radial and normal coordinates. The conformal space is defined in the range 0 G p 2 and 0 Z p 2 . The electrode is mapped onto the locus defined by Z = 0, the insulating boundary onto G \u00bc p 2 and the symmetry axis onto G = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003952_dscc2011-6191-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003952_dscc2011-6191-Figure3-1.png", "caption": "Figure 3. PARAMETERIZATION OF THE CART.", "texts": [ " We examine motion generation in the sense of [7]; the theme of [9] is steering when propulsive momentum is already available. Before proceeding, we note the qualitative similarity of our approach, in which a closed-loop system is constructed around a nonlinear plant so that the former behaves like a damped oscillator, to the formulation of PD control on Riemannian manifolds in [10] and related papers. A geometric analysis of our closed-loop system is forthcoming. We also note that our system complements the class of mobile wheeled inverted pendulums considered in [11] and related papers. Fig. 3 depicts the symbols we employ in modeling the system from Fig. 1. The wheel is assumed to align with the cart and to make contact with the ground along the cart\u2019s rear edge. The point (x,y) represents the cart\u2019s center of mass, about which the cart has rotational moment of inertia J. The longitudinal displacements a and b are constant, while the lateral displacement c(t) of the mass relative to the cart is manipulated for control. The angle \u03b8 is measured between the laboratory-fixed x axis and the longitudinal axis of the cart", " Given the system\u2019s initial state, its subsequent behavior can be reconstructed completely from the time history of the displacement c(t) together with the first-order dynamics of the system\u2019s two-dimensional nonholonomic momentum, a concept developed in detail in [6]. A complete mathematical presentation of this concept is beyond the scope of the present paper, but the essential idea is straightforward to convey. 3 Copyright \u00a9 2011 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/01/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use Consider the unconstrained motion of a rigid body in the plane \u2014 say, the cart in Fig. 3 with neither the wheel nor the additional mass present. The invariance of the cart\u2019s kinetic energy under rigid translations and rotations of the laboratory frame of reference gives rise, via Noether\u2019s theorem [12], to a threedimensional conservation law. We may equate this with the independent conservation of linear momentum in the x direction, linear momentum in the y direction, and angular momentum about the origin of the xy plane, but this decomposition is partly arbitrary. If the cart\u2019s momentum is decomposed relative to a bodyfixed frame \u2014 for instance, into longitudinal linear momentum, lateral linear momentum, and angular momentum about the cart\u2019s center of mass \u2014 then the components of momentum will evolve over time as the body frame rotates relative to the laboratory frame", " The equations governing the dynamics of these scalar momenta exemplify Lie-Poisson equations, analogous to Euler\u2019s equations for the body-fixed angular momenta of a rigid body rotating about its center of mass in three dimensions [12]. The principle of Hamiltonian reduction is manifest in the fact that if the time histories of the momenta defined in the body frame are known, then the evolution of the body frame relative to the laboratory frame can be reconstructed after the fact \u2014 meaning that the dynamics of the momenta in the body frame encode the complete dynamics of the system. The addition of the wheel to the cart in Fig. 3 breaks the symmetry in the system that engenders the conservation of threedimensional momentum. Neither the linear momentum of the cart in the x direction, nor the linear momentum in the y direction, nor the angular momentum about the origin or about the cart\u2019s center of mass is conserved with the wheel\u2019s no-slip constraint in place. If the body frame is attached to the cart where the wheel contacts the ground, however, it\u2019s apparent that the constraint opposes the evolution of only one of three components of momentum" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003361_j.robot.2010.10.007-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003361_j.robot.2010.10.007-Figure5-1.png", "caption": "Fig. 5. Free space representation.", "texts": [ "2, is compatible with a map M if for every point T \u2208 V (c) with coordinates r = [pT , z]T it holds that z \u2265 zM(p) and for each point of the base of V (c) it holds that z = zM(p). This relation allows us to formulate a more formal definition of the feasible configuration than the one given in the Section 3. Definition. A vector c represents a feasible configuration if c \u2208 C and V (c) is compatible withM . In [8], an efficient free space detection approach is proposed that facilitates SSP based on a 2.5D map. This approach consists in segmenting the xy-plane of the world coordinate system S0 into 3 subsets, as illustrated in Fig. 5: 1. free regions Cfree\u2014regions in which the robot can walk on a flat horizontal surface at a safe distance from the closest obstacles and change its walking direction without constraints; 2. C\u0304free\u2014regions in the vicinity of obstacles, in which the movement of the robot is constrained in a particular way explained later in this section; 3. obstacle regions Cob\u2014regions occupied by obstacles. Let A1, A2, . . . , An be a set of all connected regions in R2 representing flat horizontal surfaces in the robot\u2019s environment, i", " Since rrot is determined in such a way that the robot can turn in place inside the cylinder of radius rrot, the robot can freely change its orientation if its reference frame SR is centered in a point p \u2208 Cfree. The motion planning in the regions included in Cfree can be performed by one of the standard motion planning methods for walking or wheeled robots. Motion planning in C\u0304free is not so simple as in Cfree since this set includes obstacle regions which constrain the robot motion.Walking in C\u0304free may include stepping over obstacles or climbing stairs. The approach to motion planning in C\u0304free proposed in [7] is to detect a set P = {\u03a81, \u03a82, . . . , \u03a8N} of paths crossing C\u0304free, as shown in Fig. 5. Each path\u03a8i \u2208 P represents a straight line segment in the xy-plane with 4 critical points, the endpoints of the path ai and di and two points bi and ei positioned at particular locations on the path. Segment biei of path \u03a8i can be traversed by stepping over an obstacle and ai, di \u2208 Cfree. The orientation of a path \u03a8i \u2208 P relative to S0 is represented in this paper by angle \u0338 \u03a8i. Set Cfree, togetherwith set P , represents a free space representationwhich can be used for efficient SSP. In order to state the properties of the sets Cfree and P important for SSP, we define the set C\u2217free as C\u2217free = Cfree \u222a N i=1 aibi \u222a diei ", "5), derived from the properties of the sets Cfree and P given in Section 4 actually describe a set of feasible steps the robot can perform in different situations. Condition (D.1) describes the robot motion in Cfree. Condition (D.2) describes situations where the robot steps from Cfree into C\u0304free, where its motion is constrained by obstacles and it is allowed to walk by following path \u03a8i. Condition (D.3) describes a step along a path \u03a8i. The step can be extended to Cfree. If the robot is positioned in a switching point of two paths \u03a8i and \u03a8j and its orientation \u03b1 is equal to or somewhere between the orientations of these paths, as indicated in Fig. 5 by a dashed line, then, according to condition (D.4), there is enough space for the robot to change its walking direction and perform a step along one of these paths. Furthermore, the robot can also change its orientation to any orientation between \u0338 \u03a8i and \u0338 \u03a8j by turning in place, which is stated by condition (D.5). Note that conditions (D.2)\u2013(D.4) allow stepping over obstacles, sincep andp\u2032 can be at opposite sides of an obstacle, i.e. path segment pp\u2032 can include segment biei. However, the stepmust always finish in a positionp\u2032 \u2208 aibi\u222adiei orp\u2032 \u2208 Cfree" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003459_tac.1966.1098419-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003459_tac.1966.1098419-Figure7-1.png", "caption": "Fig. 7. Steering coordinate systems.", "texts": [], "surrounding_texts": [ "520 IEEE TRANSACTIONS ON 4UTOMATIC CONTROL JULY\nThe block diagram of the ship dynamics with the steering loop closed around it which corresponds to (29) is shown in Fig. 4. Using values for the hydrodynamic coefficients which correspond to the mines\\\\-eeper, the transient responses of the ship moving a t 13 knots are shown in Fig. 6 for several different conditions. Response X shon-s the lateral displacement which Lvould occur if the steering loop \\\\-ere to be suddenly closed \\vith the deviation y having an initial value of 100 feet. A disturbance impulse \\vhich imparts a heading error of 6 degrees results in the transient behavior of the ship depicted as Response B. The sudden application of a constant transverse force, say due to water current, \\vi11 cause the minesweeper to initial117 move away from the desired track. Response C shou-s the effect of a step input of transverse force of magnitude such that i t eventually results in a stead>--state yaw angle of 4 degrees. Kote the effect of the current bias integrator in Response C as the deviation settles to a zero value in the presence of a constant transverse force.\n1,'. COXCLVSIOSS\nThe motion of a ship in general cannot be readily expressed in mathematical terms \\\\-hich lend themselves to linear analysis, except for the case \\\\-here the motion can be considered as perturbations to the state of the ship occurring in the vicinity of some constant operat-\ning point. Such is the case for the ship moving along a straight-line track at constant speed. -4 linear mathematical model is derived for this mode of operation which has constant coefficients.\nThe values used for the h\\;drod?-namic coefficients xvere those obtained from model tests performed at the Experimental Towing Tank at Stevens Institute during the period 1945 to 1949 [4], [SI. Having then obtained a frequency plane representation of the ship's d>-namic characteristics, stable performance of the steering loop is readily achieved n-ith linear compensation consisting of a combination of rate feedback and a lead-lag network.\nAn automatic ship steering loop has been developed for the practical application of keeping a constant track. The frequency plane analysis with a linear model can easily be extended to include such effects as limits on the rudder rate and the rudder angle since these nonlinearities are made tractable by using describing functions. The scope of the pioneering work of Davidson and Schiff [4], [ 5 ] in the use of linear h>-drodynamic coefficients for the mathematical modeling of ship yawing dynamics has been recently broadened bJ- the addition of nonlinear hydrodynamic terms to the system equations [i]. The use of this nonlinear model should make i t possible to close a steering loop around a maneuvering ship, particularl>- one which has light coupling of roll, pitch, and heave motions into yaw motion.\nAPPENDIX\nDERIPATIOK OF EQUATIONS OF 110~101; FOR s TRACK-KEEPIXG SHIP'\nSince i t is desired to steer the ship on a track (straightline path, relative to the earth) between two given points, the appropriate reference coordinate system for this derivation has its center a t the starting point with one horizontal axis (x) directed along the desired track. The other horizontal axis ( y ) , shown in Fig. 5 , is to the right for a ship directed to\\\\-ard the destination. The z-axis pointing don-n along the vertical completes the right-handed x-p-z coordinate s>-stem. The ship's coordinate system, which is convenient for expressing rudder forces and ship resistance, has its center a t the ship's center of gravitl; (CG). The u-axis is directed forxl-ard along the ship's longitudinal axis, \\\\-hile the v-axis is positive to the right. The w-axis points down along the vertical axis through the CG. The tu-o coordinate frames are displaced by the position vector i? and the angle 8.\nThe yaw angle is the angle between the total velocity of the ship with respect to the reference frame 1,' and\nwhich is fixed with respect to the earth. Course-keeping is maintainTrack-keeping is defined as moving along a straight line path ing a fixed heading with respect to the earth, where lateral motions of the ship are accounted for b?- onl>- infrequent changes in the course to be maintained. Station-keeping is maintaining a fixed position relative to the position and velocity vector of another ship.", "1966 GOCLOWSKI AND GELB: AUTOMATIC SHIP STEERING 52 1\nthe ship\u2019s that $, 8,\nlongitudinal axis, so that sin $ = - f/ I,?. S o t e and the rudder angle 6 are all positi\\,e for a\nsteady-state turn to the righL The track-keeping motions of a ship underway in still water are sufficiently decoupled from the pitching and rolling motions to permit the consideration of only three degrees of freedom : two-dimensional motion of the CG in the horizontal plane and rotation about a vertical axis through the CG.\nSince the frequency band of interest is much higher than the earth\u2019s inertial angular ate, the reference frame fixed to the earth is taken to be the same as the inertial frame, and the basic equations of translation and rotation are\nwhere\n.nz = mass of ship - R = T , x + i , ~ ~ = i , u + l B ~ ~ -\n=position vector of ship\u2019s CG relative to refer-\np r = time derivative operator relative to reference\np,?R =acceleration of ship\u2019s CG relative to reference\nence frame\nframe -\nframe - F = total external force vector I=moment of inertia of ship about its vertical\naxis\nG, = l,w, = l \u201dWS - -\n=angular rate of ship\u2019s coordinate frame relative to reference frame\np,W, =rate of change of W relative to reference frame - T = torque about the ship\u2019s vertical axis.\nThe ship\u2019s acceleration @i?i? can be expressed as the rate of change of velocity p,i? relative to the ship\u2019s coordinates by using the vector form of the Law of Coriolis given in [8] as\np a (-1 = + Gab x (-1. Therefore,\np v ( p r R = p s ( p J Q + w s X ( P r R ) = .7 (3 2 )\nwhere\np , = time derivative relative to ship\u2019s coordinate frame\nT= specific force, F/m.\nDefine the scalar velocity relative to the reference frame in the zt and \u2018I directions as being zi and 6 , respectively.\np,i? = iwzi + iz+.\nEquations (31) and (32) can now be written in scalar form where the subscripts on the derivative operator p need no longer be included since the Coriolis law was imposed on the vector form of the equations.\np1i - w,il = fu (334\np s f w,zi = fr (33b)\np w , = h (334\nwhere h = TSI, specific torque. Dimensionless coefficients can be used to normalize force, torque, mass, and inertia by introducing the mass density of water and the ship characteristics of length, area, and ship\u2019s velocity. In general, these coefficients are defined as\nT Ctorquc =\n- d l P\n2\n- A P P\n2\n(34a)\n(34b)\n(34c)\n(34d)", "522 IEEE TR.4h\u2019SACTIOh\u2019S ON AUTOMATIC CONTROL JULY\nxhere Since the differentials in the above equations are all\np = mass density of water A =total lateral plane area of ship, including rudder\n1 =length of loaded ship a t waterline V= ship\u2019s velocity (assumed constant).\nBy combining (34a) and (34c), specific force in the u direction (F/m), is (( P / Z ) Cforee/Cmass)u, n-hich is denoted by ( Vz/Z)Cu acting in the negative direction. Similar relationships for jl , and h can be obtained. The specific forces and torques in (33) can therefore be expressed in terms of dimensionless coefficients.\n1\u20192\nI fu = - - c, + fp\nwhere\nC, =coefficient of total, less propulsive, longitudinal\nC, =coefficient of total transverse specific force C, =coefficient of total specific torque f, = specific propulsive force, F,/m.\n13-hen the ship is moving straight ahead at constant speed, the net force in the u direction is zero xith the specific propulsive force . f p equal to the forward specific resistive force V2CU/Z. Variations in the forward velocity z i will not be considered, since i t is only the transverse motion which is of interest in track-keeping. The nominal values of fo and h are zero in still water. In general, the dimensionless hydrodynamic coefficients C, and C, are considered to be functions of the three variables: 6, the dimensionless quantity ( h S / and 1c, (or - 6 / 17) ; so that the specific forces and torques in (35) can be u-ritten in differential form as\nspecific force\nof curvature of the ship\u2019s path, and is called the space angular * (&/V) is the ratio of ship\u2019s length to the instantaneous radius velocity.\nsmall deviations about their nominal zero value, they can be replaced by their actual value.\nwhere the sign of the terms containing C,.a and C,, are changed to make all the coefficients positive for ships having standard characteristics. In particular, the dimensionless coefficients are\nac, d(Zw,/V) Cz,=-- - derivative of transverse specific force co-\nefficient with respect to ship\u2019s space angu-\nlar velocity\nac, c,*; =- a+\n= derivative of transverse specific force co-\nefficient with respect to yaw angle\nCr6=- = derivative of transverse specific force co- dC,\nas efficient with respect to rudder angle\ne,,=-- - derivative of specific torque coefficient\nwith respect to ship\u2019s space angular ve-\n- ac, a (Ius/ 1,\u2019)\nlocity\nC .=- =derivative of specific torque coefficient\nwith respect to yaw angle\nac, a+\ndC, c,s = ~\nas =derivative of specific torque coefficient\nwith respect to rudder angle.\nThese coefficients include the effect of the kinetic energy imparted to the surrounding water by the accelerating ship, which is normally accounted for in hydrodynamics [4] by considering the existence of an apparent or \u201cadded\u201d mass. The hydrodynamic coefficients of the ship change progressively more as d and ws increase, so that inaccuracies are to be expected xvhen constant coefficients are used for large values of t and ws. Keedless to say, the simple approach associated with a linear system n-ith constant coefficients recommends its use wherever possible. For the track-keeping analysis, the linear approach appears to be quite valid, provided t and ws do not get too large. The coupling of zi into the 1: equation is eliminated by approximating w,zi in (33b) by ws V , which is valid in track-keeping. Then, (33) and (37) can be combined giving a constant coefficient matrix relating d and w,." ] }, { "image_filename": "designv11_20_0001557_ical.2007.4339029-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001557_ical.2007.4339029-Figure1-1.png", "caption": "Fig. 1 shows the geometric model of the 6-6 Stewart platform with planar base and moving platform. All the joints of its base and moving platform are located in respective planes. The six inputs necessary to describe the location and orientation of the upper platform are the leg lengths controlled by each prismatic joint. For a general case, the absolute local frame system O1-X1Y1Z1 and the relative moving frame system O2-X2Y2Z2 are fixed to the arbitrary points O1 on the base platform and O2 on the moving platform, respectively. The direct kinematics problem is to find the position and orientation of the moving platform supposing that the pose of the base platform is known and values for the six constraints connecting to the base and the platform are given.", "texts": [ " Substituting x, y, w into (15), \u2026, (20), solutions of r1, r2, r4, r5, u, v can be gained easily. For one solution of x, y, w there will be one solution of r1, r2, r4, r5, u, and v. By (12), (13), and (14), we can get the solutions of z, r3, and r6. For one solution of x, y, w, there will be two groups opposite sign solutions of each z, r3, and r6. . NUMERICAL EXAMPLE A set of input data for the direct kinematics of the 6-6 Stewart platform with planar base and moving platform is considered. With reference to Fig. 1, the coordinates of the base attachment points Ai (i =1, 2, 3, 4, 5, 6) with respect to the absolute local frame system O1-X1Y1Z1, the coordinates of the moving platform attachment points Bi (i =1, 2, 3, 4, 5, 6) with respect to the relative moving frame system O2-X2Y2Z2 and the actuator lengths Li (i =1, 2, 3, 4, 5, 6) are given in Table I. The 40 sets of solutions in the complex domain shown in Table are obtained. All solutions have been verified by the inverse positions analysis. TABLE I INPUT DATA FOR THE NUMERICAL EXAMPLE i xi yi pi qi Li 1 9 3 3 1 13/36205 2 6 8 2 3 65/1886302 3 0 14 1 5 65/1014653 4 -8 13 -3 4 237 5 -7 -6 -2 2 462 6 -3 -5 -1 -4 65/466706 TABLE SOLUTIONS FOR THE NUMERICAL EXAMPLE No" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000235_bf02905937-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000235_bf02905937-Figure2-1.png", "caption": "Figure 2. Warp and weft directions of the fabric on a membrane element", "texts": [], "surrounding_texts": [ "It is obvious that the architectural forms used are not often minimal surface areas. If we also take into account the orthotropic behaviour of the fabric, it is unnecessary to pursue this criterion in form finding. The following example represents a roof with a surface comparable to a Chinese coolie hat. This roof has a conical shape deviated outwards, two sides and a corner of fabric are embedded. The other two sides are cabled and free, see Figure 10. The initial form Goini is a square with a hole in the middle representing the hoop. We imposed a displacement of the hoop and one of the corners to their final positions, a stress state such as \u03c3of = 24 MPa with r = 2.5 and T = 106 N for cable tension. The warp directions are parallel to the sides of the square. Figures 11 and 12 represent the actual distribution of the stress field in Go following a nonlinear elastic calculation. The values of these stresses vary by \u00b110% when compared to the values set initially. We observed a stress concentration around the hoop and on the edges, which is normal. The second example reveals a swimming pool roof. This roof includes two rows of three hoops inclined at a \u00b110\u25e6 angle with respect to the horizontal direction. There are twenty anchorage points and twenty boltrope cables. This roof measures 37 \u00d7 7 \u00d7 7 m. Figure 13 represents the surface in equilibrium obtained by the force density method for densities equal to the unit overall except on the edge cables where they are worth 1000. Figure 14 represents the surface in equilibrium obtained by the stress ratio method with r = 1, \u03c3ot = 5 MPa and tension in the edge cables of T = 104 N . The latter presents a more flexible appearance with more curvature than the first. This example shows that a representation of the fabric behaviour through a network of bars can be used to obtain a shape in equilibrium, which often appears extremely rigid with minimized curvatures. The aspect of this form can be altered by adding more curvatures through the force densities (from configuration 13 to configuration 14). This equilibrium is however closely related to the meshing density. It is therefore practically impossible to extract the stress field related to this form through the density and meshing information. The stress ratio method offers basic advantages without requiring additional resources or creating dilemmas for the definition of data. The forms in equilibrium obtained through this method present perfect coupling between the geometry and the stress state imposed." ] }, { "image_filename": "designv11_20_0002492_iemdc.2009.5075175-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002492_iemdc.2009.5075175-Figure7-1.png", "caption": "Fig. 7 Incorrect finite-element solution: 4 blocks conducting", "texts": [ " As mentioned earlier, 2D finite-elements are sometimes used in eddy-current problems, in which the zero net current condition is imposed (usually without proof) individually across each conducting region by means of a fictional external circuit connection with a high resistance shorting the two ends. This technique requires a voltage-driven solution and a special formulation in which the circuit equations are incorporated in the field-solution matrix, [12]. Notwithstanding the excellent mathematical literature on the finite-element method, a review of ten books on finite-elements in electrical engineering fails to provide any practical reassurance of the validity of this method. Fig. 7 shows the problem of calculating eddy-currents with a conventional 2D solver using current excitation (and no external circuit connection). The problem is to calculate the no-load eddycurrent losses in the magnets at high speed. The particular example in Fig. 7 uses the time-stepping algorithm of CrankNicholson, and predicts a total loss of 2@06 W, but this result is incorrect because the \u201cinfinite length\u201d assumption permits induced current in magnet A to return in magnet B, and likewise current in magnet C to return in magnet D. Since there is no electrical connection between the magnets, this is a false result. (It is not the only example of the possibility of a false result obtained from the finite-element method by the unwary user). The simple finite-element formulation satisfies (43) only over the whole solution domain, and not individually in each magnet region" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001679_gt2008-50305-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001679_gt2008-50305-Figure5-1.png", "caption": "Figure 5. Model of rotor group, including impeller (A), bearing shaft (B) and turbine rotor (C).", "texts": [], "surrounding_texts": [ "The connection between the turbine rotor and the instability was more ambiguous, as it occurred in only ~15% of the engines. The rotordynamic modeling supported the connection between radial turbine flow induced instability and the resulting vibration measurements, but this would occur more frequently if it was part of the design. The investigation team interpreted this to indicate that the problem was not inherent in the design of the engine and its aerodynamic components. Engines that were apparently identical had nonidentical results. The team initiated an extensive review of the turbine rotor fleet and made comparisons of the good and bad turbine rotors. The turbine rotor of the microturbine is similar to other small turbomachinery products, as seen in Figures 5 and 6. There is a Mar-M 247 cast radial turbine wheel with an Inconel 718 tiebolt inertia welded to the backface, visible in Figure 6. The final machining is completed on this compound, welded assembly. The good and bad rotors, with respect to sub-synchronous vibration, were consistent for meeting the existing drawing tolerances, with the critical machined features held to within 5 \u00b5m (~.0002 in.) and the cast features conforming to typical aerospace tolerances for investment castings. It was then noted that the majority of the bad rotors were received with a high initial unbalance. It must be emphasized that the quality of the balance work and the amount of residual unbalance at time of assembly was identical for all rotors, good or bad. As a temporary measure, initial unbalance limits were applied to all incoming rotors. This was effective in protecting the Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/27/2016 T manufacturing process and the end user, but cost prohibitive due to a high scrap rate of incoming parts. The root cause had to be identified and corrected now that the problem was quarantined. It was theorized that the casting was not well centered in the workholding tool during the machining of the critical datum features used in subsequent machining steps, such as blade profile grinding and location of the pilot fit. This would create an effect where the cast aerodynamic features, as seen in Figure 7, were not symmetric about the rotating center of the part. As the casting, namely the massive hub, was eccentric, the initial unbalance of the worst components was quite high, generally proportional to the centering error. But it was still possible, even with badly unbalanced parts, to measure the machined dimensions to be in print relative to the machining datums. This was because the datums themselves were mis-aligned with the casting center of mass. It was clear at this point that there was a deficiency in the component definition that allowed this to occur. In addition to solving the manufacturing problem, the subsequent challenge was how to quantify the casting-tomachining eccentricity or variation, and how to appropriately impose this requirement on the inspection document (i.e. the blueprint). Attempts were made with CMMs of basic and advanced design to measure the cast surface of the rotor\u2019s hubline between blades with respect to the blade outside diameter profile and the machined pilot diameter, which locates the rotor to the bearing shaft. Due to problems, including overwhelming the memory of the controller of the CMM manufacturer\u2019s 4 Copyright \u00a9 2008 by ASME erms of Use: http://www.asme.org/about-asme/terms-of-use machine, no absolute measurements were made to verify the theory. In the interest of continuous improvement goals, an investment was made in a 3-D optical imaging scanner. With this device, the position of the cast surfaces with respect to the center line of the pilot fit could be made with relative ease. Figures 8 and 9 show the results of a typical bad rotor and good rotor, respectively. These images are actual components captured with the 3-D scanner, then aligned with the solid model using the machined datum features as reference points. The colors show the surface error with respect to the theoretically perfect model. The magnitudes and directions of the displacements are shown in the scale on the right side of the figures. A positive value is away from the surface, in the normal direction. This part is made with an investment cast process. Simply put, a sacrificial mold is created in a ceramic material that can endure the pouring temperatures of the metal. After the part is poured and cooled, the mold is forcibly removed from the metal part that has been created inside. The process may include chipping with hammers, blasting with media, or treating in chemical solutions. Due to the mold removal step, the blade tips in Figure 9 were bent with respect to the nominal blade shape, but evenly and symmetrically. The hub line, and therefore the flow volume of the blade pairs, is also even and symmetric about the rotor. For this reason, this part performed Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/27/201 well with respect to rotordynamics, but most likely had a reduction in aerodynamic performance. The bad rotor clearly shows that the cast surfaces are displaced approximately 750 \u00b5m (~.030 in.) radially from the axis defined by the machined diametral datum. The resulting asymmetry in the flow created the Alford\u2019s-type forces, which destabilized the rotor at the engine\u2019s highest power levels. Asymmetric features as small as 250 \u00b5m (~.010 in.) were found to be significant for some engines, although not in all. Some turbine rotors with small displacements of the cast hubline, and thus relatively low initial imbalance values, were still responsible for high sub-synchronous vibration. Further investigation using the 3-D scanner demonstrated that the rotor\u2019s axis of rotation and its casting axis were mis-oriented. Even though the axes were not highly displaced (they could even intersect), the angle between them was relatively large. This could lead to a similar aerodynamic instability without the very large initial imbalance. It should be noted that, being a complex system, there are some other factors outside of this study which affect the final stability of the complete turbomachine. 5 Copyright \u00a9 2008 by ASME 6 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded From: 6 Copyright \u00a9 2008 by ASME http://proceedings.asmedigitalcollection.asme.org/ on 01/27/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 209187. D" ] }, { "image_filename": "designv11_20_0000262_tmag.2004.824897-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000262_tmag.2004.824897-Figure1-1.png", "caption": "Fig. 1. Process of initial mesh creation. (a) Conventional 3-D mesh (without skew). (b) Deletion of elements in air gap of rotor region. (c) Skewing of rotor core. (d) Virtual tube region. (e) Mesh after connection.", "texts": [ " In our method, the auto-mesh generator is necessary; however, it is used for only creating the initial mesh, and the computational cost is very small because the conventional mesh modification method at each rotation angle is available [4]. In this paper, the method is investigated and it is applied to the 3-D magnetic field analysis of a skewed squirrel-cage induction motor. A. 3-D Finite-Element Mesh Modification Method for Skewed Motor Analysis Our newly developed method for the skewed motor analysis is carried out as follows. First, the conventional 3-D mesh for the motor analysis is prepared, which is without skew and is made by building up the 2-D mesh as shown in Fig. 1(a). The air gap between the stator and the rotor is divided into two areas. One is in the stator region and another is in the rotor region. Next, the elements in the air Manuscript received July 1, 2003. The authors are with the Department of Information Science, Gifu University, Gifu 501-1193, Japan (e-mail: yamachu@info.gifu-u.ac.jp). Digital Object Identifier 10.1109/TMAG.2004.824897 gap of the rotor region are deleted, as shown in Fig. 1(b). Then, the rotor is skewed as shown in Fig. 1(c). Instead of the deleted air gap region, the virtual tube region as shown in Fig. 1(d) is prepared. The distribution of the nodes on the outer surface of the virtual tube region is lattice, and that of the nodes on the inner surface of the virtual tube region is skewed. The elements in the virtual tube region are generated by the auto-mesh generator using the Delaunay method. The elements in the virtual tube region can be divided regularly as the nodes in the virtual tube region are distributed regularly. Therefore, the virtual tube region can be divided into some small parts, which have the same shapes, regularly", " If the analyzed region is not full model, the elements and nodes out of the analyzed region should be deleted taking into account the periodicity. It can be done easily so that the elements and nodes in the virtual tube region distribute regularly. Thus, the elements and nodes in the virtual tube region in the analyzed region remains as the new air gap of the rotor region. At last, the skewed rotor region and the stator region are connected each other using the new air gap of the rotor region as shown in Fig. 1(e). In this process, new elements, nodes and edges are generated in the air gap of the rotor region. The number and the coordinates of nodes on the surface of the new rotor region correspond to those on the surface of the rotor region. If the tetrahedral elements are used, it is necessary to consider the relationships of edges between the rotor and stator regions as well as those of nodes. Using the initial mesh created by the above method, the conventional mesh modification at each rotation angle is performed [4]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002332_j.mechmachtheory.2009.05.007-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002332_j.mechmachtheory.2009.05.007-Figure8-1.png", "caption": "Fig. 8. Photograph of kinematic models from Redtenbacher model collection at Karlsruhe.", "texts": [ " The Redtenbacher Machine Models Collection at Karlsruhe In contrast to the present use of computer simulation and animation in teaching kinematics of machines, engineering professors developed precise 3D mechanical models to help students visualize complex motions in machines as well as illustrate the mathematical curves and functions necessary for scientific design of machines. There is evidence that Willis used such models as early as 1840 in Cambridge, as did Johann Andreas Schubert at Dresden around the same time [14,16,17]. Likewise Redtenbacher developed one of the most extensive model collections of his day at Karlsruhe (see Fig. 8) with as many as 100 different models which were only surpassed after his death by his student Franz Reuleaux at Berlin. Most important for the evolution of mechanical engineering is the fact that Redtenbacher documented the construction of these teaching models in his two monographs mentioned. This offered an opportunity for others to copy these models as did a famous commercial model-maker of the time, the Schr\u00f6der company of Darmstadt. Unlike Reuleaux\u2019s original 800 model collection at Berlin that was destroyed in World War II, the Redtenbacher\u2019s Collection is remarkably preserved at University of Karlsruhe" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002824_jp808865z-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002824_jp808865z-Figure2-1.png", "caption": "Figure 2. Cross section of the novel SERRS cell design. High precision rotation of the electrode allows for minimal-invasive observation of the protein by low-power laser irradiation. Permanent electrical contact is achieved by a mercury contact connected by a conductive bolt to the Ag-RDE. The agile drive shaft is sealed against the cell body by an axial face sealing.", "texts": [ " However, in time-resolved measurements a master signal provided by a function generator (20 MHz 8021, Tabor Electronics) was used to trigger the periodic step potential of the potentiostat as well as the function generator controlling the AOM. The triggering process was screened by an oscilloscope (9354AM, LeCroy), and a synchronized electrochemical and optical excitation of the protein was achieved (Figure 8). SERRS Measuring Cell. A custom-made measuring cell was used designed for spectro-electrochemistry (Figure 2). It was based on a rotating disk electrode (RDE), turned upside down for illumination by means of a confocal microscope. To prevent photodegradation of the protein, the RDE was mounted on top of a rotating axis, which was belt-driven at a constant speed of 800 rpm by a DC-motor (FAULHABER 3863H012C 38A 1:5). In order for the surface of the electrode to stay exactly in the focal plane of the confocal microscope at all times, it had to be machined very precisely (deviation <10 \u00b5m). Moreover, the 10 mm drive shaft was inserted in the cell body by two highprecision ball bearings (GRW Gebr" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001554_eej.20585-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001554_eej.20585-Figure1-1.png", "caption": "Fig. 1. Schematic diagram of the PMIG.", "texts": [ " Therefore, clearly understanding the electrical and magnetic characteristics when creating connections to this kind of grid is extremely important in terms of the operation and design of a PMIG. In particular, in a PMIG, what kinds of effects the negative sequence rotating magnetic field generated in the stator has due to the PM rotor must be taken into consideration. Thus, in this paper the authors use circuit network analysis based on the symmetrical coordinate method and two-dimensional finite element analysis (2D-FEA) to clarify the electrical and magnetic characteristics of a PMIG when the grid voltage is unbalanced. Figure 1 shows a schematic diagram of a PMIG. As can be seen in the figure, the PMIG consists of a stator using a three-phase winding and a squirrel-cage rotor with a PM rotor inside it. The squirrel-cage rotor is directly connected \u00a9 2007 Wiley Periodicals, Inc. Electrical Engineering in Japan, Vol. 161, No. 4, 2007 Translated from Denki Gakkai Ronbunshi, Vol. 126-D, No. 8, August 2006, pp. 1126\u20131133 Contract grant sponsors: High-Tech Research Enterprise in the Ministry of Education, Culture, Sports, Science and Technology (2004\u20132006) and a Japan Society for the Promotion of Science Research Grant [2002\u20132005 Basic Research (C)]", "1 Estimation of the coil position and rotor position for current maximum In order to perform the 2D-FEA using a static magnetic field, the position of the coil at which the primary and secondary current reach a maximum, the position of the squirrel-cage rotor and PM rotor, and the secondary current distribution must be estimated as input information. Thus, first the phasor for the gap voltage, primary and secondary current are calculated using V . uv as a reference based on the symmetric equivalent circuit in Figs. 2(a) and 2(b). The phasor line when generating power (s12 < 0) is given in Figs. 3(a) and 3(b) with respect to Fig. 1. (9) (10) (11) (12) (8) (13) (14) (15) (16) In the phasor line diagrams, there is a correlation between the magnitude and the phase for the space vector diagrams during normal operation. As a result, Figs. 3(a) and 3(b) can be redrawn as Figs. 4(a) and 4(b). Consequently, the position of the coil when the primary and secondary current for each symmetric component is at a maximum and the position of the rotating magnetic field can be estimated for any instant in which the PMIG is rotating at a constant speed" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000309_s10569-004-1508-z-Figure15-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000309_s10569-004-1508-z-Figure15-1.png", "caption": "Figure 15. Family of the L1 halo to L2 halo transfers.", "texts": [ " Although it must be mentioned that perhaps it is not the most effective solution of the transfer problem with free exit of and insertion into the halo orbit. The suggested method also can be applied to analysis of the family of transfers between two given positions r0 and r1 in time T 2 \u00bdT0;T1 with given T0;T1 and step DT. In order to find this family the transfer in time T \u00bc T0 is determined first and then the transfer orbit for each time T \u00bc T0;T0 \u00fe DT;T0 \u00fe 2DT; . . . is used as a reference orbit for finding next transfer in time T\u00fe DT. Figure 15 shows the family of the halo-to-halo transfer for T0 \u00bc 180 day, T1 \u00bc 230 day and step Dt \u00bc 5 day. Parameters of the halo orbits are the same as in the example of the L1 halo to L2 halo transfer in Section 5.4. The transfer parameters, for instance the DVs necessary for the transfer, can be analyzed using this family. Figure 16 shows DV0 of launch from the L1 halo, DV1 of insertion into the L2 halo, and the total DV \u00bc DV0 \u00fe DV1 versus the transfer time. As is seen in Figure 16, the total DV reaches its minimum in the 212-day transfer" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000426_elan.200302850-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000426_elan.200302850-Figure2-1.png", "caption": "Fig. 2. Schematic diagram of the electrochemical flow cell used in the voltammetric and potentiometric measurements in flow injection system. A: 1) polyurethane resin block; 2) reference electrode (Ag/AgCl); 3) platinum electrode; 4) potentiometric sensor; 5) polyethylene tubing. B: modified carbon paste was packed into an electrode body.", "texts": [ " First, the electrochemical properties of the modified electrode in lithium solution by cyclic voltammetry in flow-through configuration were investigated. Experimental parameters, such as pHof the carrier solution, flow rate, injection sample volume, the selectivity for Li against other alkali metal and alkaline-earth ions and the response time of the potentiometric sensor were evaluated Cyclic voltammetric and potentiometric measurements were carried out with anAutolab/PGSTAT30 (Eco Chimie) under computer control. The potentiometric measurements were performed in a three-electrode cell (Fig. 2) using carbon paste electrodemodified (CPEM)with a spinel-type manganese oxide as a working electrode (indicator electrode), Ag/AgCl reference electrode and platinum auxiliary electrode (grounding electrode). For cyclic voltammetry, the potential range was from 0.4 to 1.1 V (vs. Ag/AgCl) at scan rate of 10 mV s 1, and during the measurements the solution in the cell was not flowed. The potential differences between the indicator (CPEM) and reference electrodes were measured using the GPES software (Eco Chimie) by chronopotentiometry (zero current). The body of the electrochemical flow cell (Fig. 2A) was fabricated from polyurethane resin from vegetable oil [24] (size, 48 mm 38 mm 33 mm). The effective volume of flow cell was of 77 L. All solutions were prepared using a water purified in a Millipore (Milli-Q) system. All chemicals were analytical reagent grade and used without further purification. The supporting electrolyte used for all experiments was a 0.1 mol L 1 Tris(hydroxymethyl)aminomethane (Tris) buffer solution (pH 8.3). 0.01 mol L 1 solution of lithium ions were prepared daily by dissolving lithium chloride (Merck) in 100 mL of Tris buffer", " KGaA, Weinheim decantation with deionized water, filtered and dried in air at 90 C. The modified carbon paste electrode was prepared by carefully mixing 50% (m/m) of graphite powder (1 \u00b1 2 m particle size, Aldrich), 25% (m/m) spinel-type manganese oxide and 25% (m/m) of mineral oil (Aldrich). This mixture was prepared by magnetic stirring in a becker (50 mL) containing 20 mL of hexane. The final paste was obtained with the evaporation of the solvent. The modified carbon paste was packed into an electrode body (see Fig. 2B), consisting of a plastic cylindrical tube (o.d. 7 mm, i.d. 4 mm) equipped with a stainless steel staff serving as an external electric contact. Appropriate packing was achieved by pressing the electrode surface (surface area of 12.6 mm2) against a filter paper. Before the use, the electrode was activated in a 0.01 mol L 1 lithium ions in 0.1 mol L 1 Tris buffer solution (pH 8.3) by cyclic voltammetry. Theelectrochemical cellwas inserted into anone-channel flow injection system schematically shown in Figure 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002449_j.mechatronics.2009.06.012-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002449_j.mechatronics.2009.06.012-Figure1-1.png", "caption": "Fig. 1. Motorized spindle bearing structure.", "texts": [ " The thermal expansions of the spindle, housing and bearings are calculated based on predicted temperature distributions and are used to update the bearing preloads depending on the operating conditions. Which are again used to update the thermal model [4\u20136]. Although a good spindle design with proper bearing preload has been considered acceptable with a small amount of thermal growth, the bearings loading will increase rapidly due to centrifugal force from the increase of working time and spindle speed [7\u2013 9]. The higher speed of spindle revolution, the higher centrifugal force will be created. Fig. 1 illustrates a typical spindle bearing structure. A pair of angular contact bearings is used to support the lower portion of the spindle. It is also a proven design and good solution for higher revolution speeds as the angular contact bearings are capable of absorbing both axial and radial thrust typical ll rights reserved. x: +886 4 2560 8017. ang), jjchen@cc.ncue.edu.tw of a normal machining [7,8]. In Fig. 2, as the load on the bearings increases during a normal cutting from many different directions, the resultant decrease in the gap of bearings will cause rapid temperature rising which is nonlinear [9]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000333_jra.1987.1087071-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000333_jra.1987.1087071-Figure3-1.png", "caption": "Fig. 3. Two types of actuator arrangement. (a) Relative coordinate system (RCS): The muscle-bone system of living things including animals has this type of actuator arrangement. (b) Body coordinate system (BCS): BCS actuator arrangement is limited to artificial legs.", "texts": [ " 2) The influence of the difference in the arrangement of actuators on specific resistance is discussed. 3) The effect of the inertia reaction force of the ground is discussed in order to satisfy the approximate balance of forces and moments when an overall walking machine assembly is assumed as a free body. 4) The effect of nondimensional height 6 is examined. A . Considerations on Arrangement of Actuators The arrangement of actuators of walking machines are classified into two types as shown in Fig. 3. One type shown in Fig. 3(a) is such that the rotation of the thigh actuator drive shaft makes the fixed side of the shank actuator rotate at the same angle. This type is called a relative coordinate system (RCS) arrangement in this paper. Another type shown in Fig. 3(b) is such that regardless of the rotation of the thigh actuator drive shaft, the fixed side of shank actuator points in the same direction. This type is called body coordinate system (BCS) arrangement in this paper. The muscle-bone system of living things including animals and the legs of various walking machines have RCS actuator arrangement. As indicated in Fig. 3(b), BCS actuator arrangement is limited to artificial legs. Although the specific resistance of BCS actuator arrangement has been examined [6], the specific resistance of RCS actuator arrangement has not been examined except the 21 massless model (m2 = 0) [7]. For example, let us assume a case in which supporting force F acts as the reaction force of the ground, with the mass of the legs assumed zero, as shown in Fig. 3. In the case of RCS arrangement, magnitude of the moment acting on the hip joint is expressed by the product of the supporting force F and A x (Ax is the horizontal distance between the supporting point and the hip joint). In the case of BCS arrangement, the magnitude of the moment acting on the hip joint appears as if a vertical force F acted on the knee joint. Therefore the difference in the actuator arrangements should impose a decisive influence on the specific resistance. B. Assumption for Analysis Now we will discuss an n-legged walking model (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001915_j.triboint.2008.09.003-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001915_j.triboint.2008.09.003-Figure1-1.png", "caption": "Fig. 1. Schematic of a typical radial lip seal.", "texts": [ " Radial lip seals for rotating shafts have been widely used in many different industrial applications since 1940s. These seals function to prevent leakage of a lubricant and to exclude contaminants. A radial lip seal is composed of an oil resistant elastomeric body bonded to a metallic case and is designed to have an interference fit with the shaft. A garter spring is often used to provide additional force which compensates for the load and flexibility loss that occurs when rubber materials are exposed to hot oil for extended periods of time. Fig. 1 shows the schematic of an installed lip seal and its components. The sealing mechanism of this seemingly simple device is now better understood through modern computational techniques and the accumulated results of fifty years of experimentation. It has been shown by the pioneering work of Jagger [1] that, under steady operational conditions, a thin lubricating oil film of a few micro-meters thick separating the lip from the shaft is build up in a successful lip seal to reduce wear and prevent mechanical and thermal damage to the lip" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003086_tsmca.2010.2076405-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003086_tsmca.2010.2076405-Figure1-1.png", "caption": "Fig. 1. Six-DOF Stewart platform.", "texts": [ "ndex Terms\u2014Independent component analysis (ICA), kinematics, Nelder\u2013Mead (NM) algorithm, Stewart platform. I. INTRODUCTION As a major type of parallel manipulator, the six-degree-of-freedom (DOF) Stewart platform has been receiving extensive attentions since the early 1990s and been applied in manufacturing, medical robots, and flight simulation systems [1]. As shown in Fig. 1, a six-DOF Stewart platform consists of six identical telescopic links, a mobile platform, and a base platform. Each link connects the base and the mobile platforms with a universal joint and a spherical joint, respectively, and the link length is controlled by a cylindrical joint. The base platform is fixed as a base, and the mobile platform has six-DOF motion with respect to the base platform. In the forward kinematics analysis of the Stewart platform, the link lengths are used to solve the position and orientation of the mobile platform relative to the base platform while, in the inverse kinematics, the position and orientation of the mobile platform is used to determine the link lengths", " 2, as follows: 1) Rotate the system Oxyz about the z-axis by \u03b1, and the x-axis now lies on the line ON ; 2) rotate the system Oxyz again about the new rotated x-axis by \u03b8, and the z-axis is now in its final orientation; and 3) rotate the system Oxyz a third time about the new z-axis by \u03b2. R can be easily derived, and it is given by (2), as shown at the bottom of the page. For the convenience of expression, \u03b8, \u03c8, and \u03d5 are called transformation angles from the coordinate system Oxyz to Ox\u2032y\u2032z\u2032. For the convenience of forward kinematics analysis, three Cartesian coordinate systems have been set up for the Stewart platform. In Fig. 1, the B-frame (x, y, z) is fixed at the center O of the base platform with the z-axis directed vertically out of the plane. The T -frame (x\u2032, y\u2032, z\u2032) is fixed at the center O\u2032 of the mobile platform with the z\u2032-axis directed orthogonally out of the plane. An extensible virtual bar is introduced between the two centers, and the bar is denoted as the M -bar. As a result, another coordinate system, M -frame (x\u2032\u2032, y\u2032\u2032, z\u2032\u2032), is also introduced. The origin of the M -frame is O, and the z\u2032\u2032-axis is along the centerline of the M -bar" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002095_20080706-5-kr-1001.00301-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002095_20080706-5-kr-1001.00301-Figure1-1.png", "caption": "Fig. 1. Quadrotor model", "texts": [], "surrounding_texts": [ "Several works deal with the quadrotor modelling (see for example Alabazares [2007] and Castillo et al. [2005]). In this section we recall a sketch of modelling as well as the notations are introduced. In order to model the four-rotor rotorcraft dynamics two frames are defined i .e. Ri(O, \u2212\u2192 E 1, \u2212\u2192 E 2, \u2212\u2192 E 3) is an inertial frame attached to the earth and Rb(G, \u2212\u2192e 1, \u2212\u2192e 2, \u2212\u2192e 3) is a body fixed frame attached to its center of mass. The UAV model can be deduced from the rotation dynamic Newton-Euler law (1) and the translation dynamic Newton-Euler law (2) ( dI\u2212\u2192\u03c9 dt ) Rb + \u2212\u2192\u03c9 \u2227 (I\u2212\u2192\u03c9 ) = \u2212\u2192 M (1) m ( d\u2212\u2192v G dt ) Rb + \u2212\u2192\u03c9 \u2227 (m\u2212\u2192v G) = \u2212\u2192 F (2) where I is the inertia matrix and \u2212\u2192\u03c9 = (p, q, r)\u22a4 is the angular velocity both expressed in the body fixed frame. \u2212\u2192 M represents the torque derived from the differential rotors thrusts, m is the vehicle mass, \u2212\u2192v G = (u, v, w)\u22a4 is the center of mass velocity expressed in the body fixed frame and \u2212\u2192 F is the sum of the four rotor thrust\u2212\u2192 T = \u2212T\u2212\u2192e 3 and the weight \u2212\u2192 P = mg \u2212\u2192 E 3. In order to express the weight with respect to the body fixed frame, the attitude matrix R must be used. By means of Euler angles \u03a6(roll) , \u03b8(pitch) , \u03a8(yaw), the attitude matrix can be written as R = ( c\u03b8c\u03a8 s\u03a6s\u03b8c\u03a8 \u2212 c\u03a6s\u03a8 c\u03a6s\u03b8c\u03a8 + s\u03a6s\u03a8 c\u03b8s\u03a8 s\u03a6s\u03b8s\u03a8 + c\u03a6c\u03a8 c\u03a6s\u03b8s\u03a8 \u2212 s\u03a6c\u03a8 \u2212s\u03b8 s\u03a6c\u03b8 c\u03a6c\u03b8 ) (3) where c. = cos(.) and s. = sin(.). From the rotation dynamic Newton-Euler law (1) the dynamics of the angular velocity is given by p\u0307 = \u2212 Izz \u2212 Iyy Ixx qr + \u03c4\u03a6 Ixx q\u0307 = \u2212 Ixx \u2212 Izz Iyy pr + \u03c4\u03b8 Iyy r\u0307 = \u2212 Iyy \u2212 Ixx Izz pq + \u03c4\u03c8 Izz (4) where Ixx , Iyy and Izz are the inertia matrix terms expressed in the principal inertia axis, \u03c4\u03a6 , \u03c4\u03b8 and \u03c4\u03c8 represent the control torques due to the differential rotors thrusts. Using the center of mass dynamics equation (2) and the attitude matrix (3), the translation dynamics with respect to the body frame is u\u0307 = \u2212qw + rv \u2212 g sin \u03b8 v\u0307 = \u2212ru+ pw + g sin\u03a6 cos \u03b8 w\u0307 = \u2212pv + qu+ g cos \u03a6 cos \u03b8 \u2212 T m (5) where T is the total thrust which represents a control input. The time derivative of the Euler angles are related to the angular velocity \u2212\u2192\u03c9 by the following relation \u03a6\u0307 = p+ q tan \u03b8 sin\u03a6 + r tan \u03b8 cos \u03a6 \u03b8\u0307 = q cos \u03a6 \u2212 r sin\u03a6 \u03c8\u0307 = q sin\u03a6 cos \u03b8 + r cos \u03a6 cos \u03b8 (6) The dynamic model of the four-rotor rotorcraft is then given by equations (4), (5) and (6)." ] }, { "image_filename": "designv11_20_0000694_1.2000269-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000694_1.2000269-Figure3-1.png", "caption": "Fig. 3 Pico slider with a crown of 25.4 nm and a camber of 2.5 nm. The base recess is 2.5 m.", "texts": [ " We observe that the fly height value is slightly reduced when intermolecular forces are considered, which is expected as the intermolecular forces are attractive in nature. But since the fly heights are more then 10 nm, the reduction in fly height is not much, as the attractive intermolecular forces are very weak at such large separations. We also observe that the intermolecular forces slightly increase the magnitude of the pitch and roll angles. Static simulations were also carried out at 3500 disk rpm for the slider design shown in Fig. 3 at three different radial positions. The fly height, pitch, and roll with and without intermolecular forces are shown in Table 2. Since this is a relatively low flying slider, there is significant reduction in the fly height due to the intermolecular forces. Due to the attractive nature of the intermolecular forces their inclusion reduces the fly height at all three radial positions. Figure 4 shows that the magnitude of the intermolecular force is maximum at the OD and minimum at the MD radial position and hence there is a ce on the static performance of the ce on the static performance for the for for of Use: http://www", " The solver converges to a stable fly height when the intermo- 200 / Vol. 128, JANUARY 2006 rom: http://tribology.asmedigitalcollection.asme.org/ on 01/28/2016 Terms lecular forces are not considered, but the inclusion of the intermolecular forces results in large fly height modulations, which implies instability of the system. Dynamic simulations were also carried out with intermolecular forces at the disk rpm of 7200 as shown in Fig. 7. The solver converges to a stable fly height value at this disk rpm. The fly height diagrams for the slider shown in Fig. 3 are plotted in Fig. 8 for three different radial positions. We observe that at the disk rpm of 3500, there exist multiple equilibrium points at all three radial positions. Hence the simulations started with some perturbation about the steady state results in large modulations in fly height as shown in Fig. 6. However, only one stable equilibrium exists at a disk rpm of 7200. Hence the system always converges to a steady state for all small perturbations about the equilibrium point as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.131-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.131-1.png", "caption": "Fig. 14.131 Dozer with blade position controlling system for precision leveling of terrain", "texts": [ "7 m3 undercarriage, is hydraulically driven. Boreholes 846 mm in diameter can be drilled. Besides whole machines, computerized systems for controlling construction machines, allowing one to automate partially the operation of machines, are also available. Offered control systems can be installed on excavators, dozers, graders, and asphalt pavers. They incorporate tachometers, global position system (GPS) receivers or laser surveyor\u2019s levels. A simple example here is a control system for setting the position of the dozer\u2019s blade (Fig. 14.131) so that smooth, leveled soil surfaces can be obtained. A rotary laser on a tripod produces a horizontal reference plane. The control system with a signal receiver, mounted on the dozer, automat- Part B 1 4 .8 ically adjusts the elevation of the blade as the laser receiver follows the adopted reference plane. A computerized control system for a grader works as follows. The task to be performed is stored in the memory of the machine\u2019s PC in accordance with a digital work execution scheme. The machine\u2019s actual location relative to the design data is determined by a tachometer or the GPS, which compares it with the design terrain elevation at a given point and on this basis sends signals to the machine\u2019s hydraulic system" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003648_icems.2011.6073664-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003648_icems.2011.6073664-Figure8-1.png", "caption": "Fig. 8 Prototype of duble layer FSPM motor 12/10.", "texts": [], "surrounding_texts": [ "In order to predict the maximum short-circuit current, the short-circuit current against the rotor velocity is calculated analytically and then verified by simulations as well as by experiments. The results are shown in the Fig. 7. It is found that at high rotor velocity, the short-circuit peak currents trend to be constant, because at high rotor velocity, the influence of resistance and the mutual inductances can be neglected, and the short-circuit current can be obtained by the no-load phase flux linkage divided via the phase self inductance, which leads to a short-circuit current independent on the rotor velocity. Furthermore, because the self inductance of the double layer FSPM motor is approximately two times higher than that of the single layer FSPM motor, the short-circuit current of the single layer could be two times lower than that of the double layer FSPM motor. This lower short-circuit current is essential for critical applications, because to compensate the resistant torque due to the short-circuit currents, lower normal currents are needed. This leads to lower copper losses and also lower disturbance of the system.\nIV. CONCLUS\u0130ON\nIn this paper, between a double layer FSPM motor and a single layer FSPM motor the comparison in terms of self and mutual inductances as well as the maximum short-circuit current against the rotor velocity is carried out. The results have shown that the single layer FSPM motor has higher fault tolerance capability, because it has higher self inductances while with lower mutual inductances. Thus, under faulty mode as short-circuit, for example, the higher self inductances limit the increase of short-circuit currents and the lower mutual inductances, which means lower magnetic coupling, prevent the faults from infecting the normal phases. \u0130t should also be noted that the increase of self inductances will decrease the power factor, while for the critical applications, this is not the main topic and has not been studied in this paper. For the single layer FSPM motor, it should be noted that, due to the non-symetric structure, there will be a strong normal force. This would lead to a high vibration level. Thus, in practical applications, it is proposed to use the FSPM motor with 24 stator teeth and 20 rotor poles. Like this, for each phase, we can have two opposite teeth. As a result, the sum of normal force versus rotor position could be decreased considerably, and the vibration could be limited to a low level.", "V. APPEND\u0130X\nSTRUCTURAL DATA OF FSPM MOTOR 12/10 (PROTOTYPE)\nNumber of stator/rotor poles (Ns/Nr) 12/10 Stator outer radius (rs outer) 77.2 mm Rotor outer radius (rr outer) 43 mm Stack length (La) 60 mm Air-gap length (e) 0.2 mm Number of phase coils (N) 50 Filling factor (kb) 0.4 Remanent flux density (Br) (200\u00b0C) 0.8 T Self inductance (L0) 2.6 mH Mutual inductance (M0) 1.02 mH No-load phase flux ( ) 51 mWb Moment of inertia (J) 1.8\u00d710-3 kg\u00b7m\u00b2 Coefficient of friction (f) 20\u00d710-6 Nm\u22c5s\nVI. REFERENCES\n[1] E. Hoang, A. H. Ben-Ahmed, and J. Lucidarme, \"Switching flux\npermanent magnet polyphased synchronous machines,\" in in Proc. 7th Eur. Conf. Power Electron. Appl., 1997, p. 903\u2013908.\n[2] Z. Q. Zhu et al., \"Analysis of electromagnetic performance of fluxswitching permanent magnet machines by non-linear adaptive lumped parameter magnetic circuit model,\" IEEE Trans. Magn., vol. 41, no. 11, p. 4277\u20134287, Nov. 2005.\n[3] M.-A. Shamsi-Nejad, B. Nahid-Mobarakeh, and S. Pierfederici, \"Fault tolerant and minimum loss control of double- star synchronous,\" IEEE Trans. Ind. Electron., vol. 55, no. 5, p. 1956\u20131965, may 2008.\n[4] N. Takorabet et al., \"Study of Different Architectures of Fault Tolerant Actuator Using a Double-Star PM Motor,\" in IEEE Industry Applications Society Annual Meeting, 2008. IAS'08, 2008.\n[5] N. Ertugrul, W. Soong, G. Dostal, and D. Saxon, \"Fault tolerant motor drive system with redundancy for critical applications,\" in Power Electronics Specialists Conference, IEEE 33rd Annual, 2002, pp. 1457- 1462.\n[6] N. Bianchi, S. Bolognani, M.D. Pre, and G. Grezzani, \"Design considerations for fractional-slot winding configurations of synchronous machines,\" IEEE Trans. Ind. Appl., vol. 42, no. 4, pp. 997-1006, July-Aug. 2006.\n[7] A.S. Thomas, Z.Q. Zhu, R.L. Owen, G.W. Jewell, and D. Howe, \"Multiphase Flux-Switching Permanent-Magnet Brushless Machine for Aerospace Application,\" IEEE Trans. Ind. Appl. , vol. 45, no. 6, pp. 1971- 1981, Nov./dec. 2009.\n[8] R.L. Owen, Z.Q. Zhu, A.S. Thomas, G.W. Jewell, and D. Howe, \"FaultTolerant Flux-Switching Permanent Magnet Brushless AC Machines,\" in Industry Applications Society Annual Meeting, 2008. IAS '08. IEEE, 5-9 Oct. 2008, pp. 1-8." ] }, { "image_filename": "designv11_20_0003014_j.dyepig.2010.10.003-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003014_j.dyepig.2010.10.003-Figure1-1.png", "caption": "Fig. 1. Schematic view of the fluorescence measuring set-up.", "texts": [ " Additional filters in front of the camera enable a reduction of the stray light and a pre-selection of the detection wavelength region [18]. The excitation light can be also directed to the backside of the cell for adsorption measurements. Such a modular set-up made the measuring console very flexible to various measuring regimes. The emitted radiation is taken by a deep cooled CCD camera. The measuring arrangement, a fluorescence reader (Sensovation AG, Stockach), consists of a lamp, a shutter, filters for excitation and emission (Fig. 1) and a 3-stage Peltier cooled CCD image sensor with integrated micro-lenses and a high quantum efficiency (up to 90%) for detecting the fluorescence signal. For the preparation of TEOS/CHIT/CPO/GOx; GOx (1 mg), TEOS/ CHIT (100 mL) and CPO (75 mL, 1 mg/mL) were mixed and spotted to the cleaned glass surface and allowed to dry at room temperature for 5 min. After that, spots were treated with glutaraldehyde solution (2.5% in potassium phosphate buffer, 50 mM, pH 7.5) for 2 min and washed with distilled water and phosphate buffer solution, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000057_s0069-8040(08)70029-3-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000057_s0069-8040(08)70029-3-Figure8-1.png", "caption": "Fig. 8. Typical tubular electrode assemblies. (a) Integral construction. A, Generator electrode; B, detector electrode; C, reference electrode; D, counter electrode; E, porous frits; F, ball and socket joints; G, epoxy resin. (b) Demountable type. A, Generator electrode; B, counter electrode; C, Teflon spacers; D, reference electrode; E, Teflon cell body; F , brass thread. (From ref. 128.)", "texts": [ " Unfortunately, if a peristaltic pump is employed, then sheathing the tubes in the pump is not possible: these are normally of silicone rubber, which is exceptionally permeable to oxygen. Where deoxygenation is necessary, the peristaltic pump is arranged to suck the solution and the cell is placed as close to the reservoir as possible. A note of caution is that, in this conformation, air can enter the solution relatively easily so that junctions must be well sealed. The arrangement of the flow systems is very similar to that employed in flow-injection analysis [ 135,1361. For use with tubular and channel electrodes (Fig. 8), no extra special 395 References pp. 434-441 difficulties arise so Iong as, if they are of the sandwich type, there are no leaks, especially round 0 ring joints. The same problem may occur with wall-jet cells (Fig. 9). If there are any small leaks caused by incorrectly fitting 0 rings, then this can usually be solved by a piece of Teflon tape on the screw thread. Additionally, for wall-jet cells, care needs to be exercised in filling the cell with solution so that trapping of air bubbles is avoided" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001944_pesc.2008.4591895-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001944_pesc.2008.4591895-Figure2-1.png", "caption": "Fig. 2. The stator and rotor flux linkages in various reference frames.", "texts": [ " The IPMSM model in the stator flux reference frame is firstly described in the subsequent section, followed by the explanation of the proposed method. Subsequently, the modeling and experimental comparisons between the classical and modified DTC schemes are presented. Then, the speed estimation scheme based on the stator flux linkage and torque angle is described. Finally, the experimental results of the sensorless drive are presented. The stator flux linkage vector, s and the rotor flux linkage vector, f can be drawn in the rotor flux (d-q), stator flux (x-y) and stationary ( - ) frames. The various reference frames are depicted in Fig. 2. 978-1-4244-1668-4/08/$25.00 \u00a92008 IEEE As elaborated in [3], the machine equations in the d-q reference frame are as follows. d d s d r q q q s q r d d v R i dt d v R i dt (1) d d d f q q q L i L i (2) where dL , qL and sR are the machine dq axes inductances and stator resistance respectively. With (3), these equations can be transformed to the x-y reference frame. cos sin sin cos x d y q F F F F (3) where is the load angle and F represents voltage, current or flux linkage. Furthermore, it can be shown that the torque T of an IPMSM in the x-y reference frame is given by 3 2 s yT P i (4) where P is the number of pole pairs in the machine" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000522_s11249-006-9062-3-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000522_s11249-006-9062-3-Figure2-1.png", "caption": "Figure 2. Model of dry, frictionless point contact used in the simulations.", "texts": [ " To illustrate the differences the various surfaces evaluated, contour plots of selected ground, honed, and polished surfaces are shown in figure 1. Surface texturing or *To whom correspondence should be addressed. E-mail: a-martini@northwestern.edu 1023-8883/06/0900-0243/0 2006 Springer Science+Business Media, Inc. DOI: 10.1007/s11249-006-9062-3 patterning has been shown to significantly impact contact pressure and subsurface stress distributions [3,5\u20138]. Each of the 59 surfaces was used in a contact simulation of dry, frictionless point contact. The simulation model is shown in figure 2. The material properties, operating conditions, and numerical parameters for the simulation are summarized in table 2. The elastic-perfectly plastic contact model used here is the same as that described by some of the authors in a previous publication [10]. However, in this case the simulations are of isothermal contact and therefore do not incorporate the thermal displacement effect. Basically, the logic performed by the simulation code is to satisfy the constraints of zero pressure where there is no contact, pressure between zero and a pressure limit at contact points, and load balance" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001101_ac50016a047-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001101_ac50016a047-Figure1-1.png", "caption": "Figure 1. Development of the EFL used to interrogate the Gran\u2019s plot obtained from the potentiometric titration of 0.125 mmol of bromide in 0.1 M KN03 (ionic strength adjustor solution) using 0.05 M AgN03 titrant", "texts": [ " Excess HC104 was evaporated, then the solution was diluted with deionized water and adjusted to pH 1.2 with HC104. Procedure. Most titrations were carried out in a circulating water bath controlled at 25 OC f 0.5\u201d. The starting volume was 50 mL, 0.1 N in KN03 ISA. Approximately 0.05 N titrant was used to titrate -0.125 mmol of Br-, C1-, or SCN-, except for the data in Table 111. Other experimental procedures were similar to those described previously (6). The EFL. The technique for determining the EFL is shown in Figure 1, where bromide is determined with a Br- ISE and silver nitrate titrant. Figure l a shows the titration curve and its ANALYTICAL CHEMISTRY, VOL. 49, NO. 8, JULY 1977 1251 corresponding Gran plot. Figure Ib illustrates the use of cursors to select a window segment size (data between cursors 1 and 3) and a step size (distance between cursors 1 and 2) for each successive calculation. The calculations are terminated when cursor 3 moves past cursor 4. For each segment, a linear least-squares fit to the data points is computed and extrapolated to the z axis to estimate the volume of titrant required to reach the equivalent point" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003680_j.jcis.2011.11.031-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003680_j.jcis.2011.11.031-Figure1-1.png", "caption": "Fig. 1. Sketch of two spherical drops with equal radii, a, which stay at a closest distance between them, h. The drops move toward each other with given approaching velocity V. The z-axis of Cartesian coordinate system Oxyz coincides with the axis of revolution and the plane z = 0 represents the plane of symmetry of the system.", "texts": [ " The role of interfacial rheology, Marangoni effect, bulk and surface diffusivities is accounted for in the model equations (Section 2). For small deviation of adsorption from its equilibrium value the exact solution of the problem in bipolar coordinates is described (Sections 3 and 4). Some useful asymptotic solutions are reported and the numerical results for different systems are discussed in Section 5. We consider two identical drops (bubbles) approaching each other with a given velocity, V (Fig. 1). The physicochemical parameters of the fluid phase inside the drops (the disperse phase) will be denoted with the subscript \u2018\u2018d\u2019\u2019, while the respective parameters of the continuous fluid phase \u2013 with the subscript \u2018\u2018f\u2019\u2019. For example, the mass densities are denoted as qd and qf, the dynamic viscosities \u2013 as gd and gf, the vectors of the fluid velocity are vd and vf, and the pressures in the respective phases are pd and pf. We are interested in the hydrodynamic interactions between small bubbles and drops when the distances between them and the relative fluid velocity are small", " For small particle sizes, a, the capillary numbers are small and the possible deformations of the interfaces can be neglected when the leading order solution of the problem is derived. Thus, the considered problem is steady state, the drainage velocity is V = dh/dt, and the analytical solution of the simplified problem for nondeformed droplets (bubbles) can be obtained in bipolar coordinates. First, one introduces Cartesian coordinate system Oxyz, in which the z-axis coincides with the axis of revolution and plane z = 0 represents the plane of symmetry of the system (Fig. 1). The respective cylindrical coordinate system has radial coordinate r, polar angle u, and x \u00bc r cos u; y \u00bc r sin u: \u00f03:1\u00de The bipolar coordinates, x1 and x2, are defined with respect to the positions of the poles at the axis of revolution with coordinates z = \u00b1c (see Fig. 2) as [18,19] r \u00bc c b sin x2; z \u00bc c b sinh x1; \u00f03:2\u00de b cosh x1 cos x2: \u00f03:3\u00de The plane of symmetry (z = 0) corresponds to x1 = 0 in these coordinates. The part of the axis of revolution, for which |z| > c, is described by x2 = p, while that, for which |z| < c \u2013 by x2 = 0 (Fig", " Therefore, the ratio C1/C0 is small because of the assumption for small Peclet number of the considered problem. In the volume of the disperse phase, which contains the pole A (Fig. 2), Cd,n(x1) is proportional only to exp( knx1). Using boundary condition (2.12) and Eqs. (4.1) and (4.2) we obtain cd;1 \u00bc Va Ds @cd @ ln C 0 \u00f02b\u00de1=2 sinh\u00f0e=2\u00de cosh e X1 n\u00bc0 gn exp\u00bdkn\u00f0e x1\u00de Pn\u00f0n\u00de: \u00f04:3\u00de To obtain the solution for cf,1 in the volume of the continuous phase together with boundary condition, Eq. (2.12), one applies the condition for symmetry with respect to the plane x1 = 0, which corresponds to z = 0 (Fig. 1). Hence, Cf,n(x1) is proportional only to cosh(knx1) and one derives cf ;1 \u00bc Va Ds @cf @ ln C 0 \u00f02b\u00de1=2 sinh\u00f0e=2\u00de cosh e X1 n\u00bc0 gn cosh\u00f0knx1\u00de cosh\u00f0kne\u00de Pn\u00f0n\u00de: \u00f04:4\u00de In general case of solutions containing one surface-active component, two sets of unknown constants have to be determined: u1, u2, . . . including in the serial expansion for the surface velocity (see Section 3.1); g0, g1, g2, . . . appearing in the right-hand side of the expression for the deviation of adsorption, Eq. (4.2). To calculate them we will use two boundary conditions: (i) the mass balance equation for adsorption, Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001657_bfb0119410-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001657_bfb0119410-Figure4-1.png", "caption": "Figure 4: Three Degree-of-Freedom Manipulator", "texts": [ " The suspension of the rover is constructed from aluminum square tubing, and includes integral motor mounts. The body is made from a formed aluminum sheet and supports the system electronics and sensors. The body also serves as an attachment point for a manipulator arm and stereo cameras. A mechanical differential allows the body to \"split the difference\" of the two rocker angles. The total cost of construction of the mechanical structure was approximately $2500. The three degree-of-freedom manipulator mounted on the front of the rover is shown in Figure 4. The manipulator's light weight (approximately 16 ounces) is achieved by using low-weight aluminum members and small, highly geared motors. The joints are driven by MicroMo DC motors with gear ratios of 2961:1, 3092:1, and 944:1 at the trunk, shoulder, and elbow joints, respectively. With the high gear reduction, the manipulator is capable of exerting large forces. In some configurations, it can exert a force equal to one-half the rover weight. This highforce capability could be useful for manipulator-aided mobility or trap recovery" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001222_tro.2005.858850-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001222_tro.2005.858850-Figure7-1.png", "caption": "Fig. 7. Four-arm parallel mechanism with colinear end-effector joints and", "texts": [ " This characteristic greatly contributes to the lack of stability of this mechanism, as is evident in the following analysis. We now define the conditions where (45) maps the nonzero vector [ 1 2 3 4 5 6 v1 v2 v3] T to the zero vector. To facilitate the analysis, we make the following decompositions: p26 = p12 + p15 + p56 p37 = p13 + p15 + p57 p48 = p14 + p15 + p58 These are substituted into (45), and the equations corresponding to rows 4\u20136 are examined x = 1h1p15+ 4( h4p13+h4p15 h4p75)+v2p75p85+v3p75hn: (46) From Fig. 7, h1 k h4, thus if 4 = 1, then x = 1( h4p13 + p75h4) + v2p75p85 + v3p75hn: (47) Equivalently x = [h4p13 p75h4 p75p85 p75hn] 1 v2 v3 = A1 1 v2 v3 : The following transform is applied to A1: pT75 (pT75) ? 1 (pT75) ? 2 A1 = A11: (48) The first row ofA11 is [pT75(h4p13) 0 0]. The first element is zero when the vector p75 is parallel to the base, or equivalently, if end-effector rotation about the y axis is zero. In this condition, matrix A1 is rankdeficient, and maps the nontrivial vector [ 1 v2 v3]T to x = [0 0 0]T ", " We now find the condition where the following is true: 0 0 0 = [h1p15 h5p48 h6p48 p85p75 p85hn] 1 5 6 v1 v3 : (49) These equations correspond to the last three rows of JC . Having found values for 1; v3, this equation is rewritten as 1h1p15 v3p85hn = [h5p48 h6p48 p85p75] 5 6 v1 : (50) When the matrix on the right side of this equation is full rank, unique values for 5; 6, and v1 can be found, and the null direction for JC is completely determined. The mechanism exhibits single-arm passive joint self-motion if this matrix is rank-deficient. The analysis of this mechanism JC has revealed that in the configuration of Fig. 7, it is unstable at all poses where the y component of the XY Z fixed angles is zero. The flaws in this architecture are apparent from this analysis, and include the following. 1) The collinearity of three end-effector spherical joints results in y1h = [0 0 0], which allows the extra DOF in (50). 2) Joints 1 and 4 being parallel allows a simple cancellation of the h1p15 by selecting 4 = 1. Rotating one of these joints relative to the other reduces the number of unstable poses. 3) The collinearity of the end-effector spherical joints 4, 5, and 6, and the proportionality of end-effector and base vectors allows the values of 1; v2; v3 of (48) to simultaneously satisfy (49)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001628_ias.2007.344-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001628_ias.2007.344-Figure1-1.png", "caption": "Figure 1. Modeled bearing fault", "texts": [ " Thus, the detection of bearing fault can be realized by analyzing the stator current spectrum and investigating its harmonics at the frequencies given by Equation (1). III. EXPERIMENT RESULTS Bearing fault experiments of a 3kW, 380V, 50Hz, 4-pole, typed Y100L-2 induction motor have been completed. The motor bearing is typed 6206 with the following parameters: bn =9, BD=10mm, PD=46mm and \u03c6 =0 . Here only cites the results for outer race defect. This type of bearing fault is artificially modeled with electro-erosion (1mm hole performed in outer race), displayed in Fig. 1. Fig. 2 provides the stator a-phase current and its spectra under the condition that the test motor is idle, and healthy. Fig. This work is supported by National Natural Science Foundation of China (50407016). 0197-2618/07/$25.00 \u00a9 2007 IEEE 2277 3 provides that under the condition that the test motor is idle, and with bearing outer race fault. Here, 1.176\u2248vf Hz, and 1.226,1.126\u2248bearingf Hz for k=1. 0 -3 -2 -1 0 1 2 3 1 2 3 4 5 6 7 8 9 10 Time / s St at or a -p ha se c ur re nt / A Frequency / Hz A m pl itu de / A 0 50 100 150 200 250 0 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001753_2013.36770-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001753_2013.36770-Figure1-1.png", "caption": "FIG. 1 Diagram illustrating the overturning motion of a wide-frontend tractor.", "texts": [ " From the above discussion, it is noted that only a portion of the tractor tips about the first tipping axis while the entire tractor tips about the second tipping axis. Since the kinematics to fol low treats the motion of the center of gravity of the portion of the tractor which is tipping, a distinction is made between the center of gravity (CGTP) of the part of the tractor rotating about the first axis and the center of gravity (CGT) of the entire tractor in its orien tation for tipping about the second axis. Fig. 1 illustrates the geometry of both overturning motions. An x, y, z axis system with associated orthogonal unit vectors l, j , and k is located with its origin at the \"contact point\" of the rear tire remaining on the ground during the overturning motion. Since the initial tipping motion does not take place around either the x, y, or z axes, it is convenient to define the skew coordinate axis about which the tractor is assumed to tip. Following the method of Pershing and Yoerger (1964), A let j be a unit vector in the direction of the first tipping axis", " yly and z1 are the coordinates of the hinge point of the front axle, \u00bb\u00bb n n '\u00bb j = ' ( x i i + y i J + z i k V B [i] where B = \\ : + yf+z Let x t , y t , z t be the coordinates locating the CGTP when the tractor is in full , undisturbed contact with the 1974 - TRANSACTIONS of the ASAE A A A ground. Then r t p = x t p i + y t p j + z t p k is the posit ion vector from the origin t o the CGTP. As the t rac tor tips sideways, the CGTP is constrained to move in a circu lar path (see circled i tem 1 in Fig. 1) about a center of ro ta t ion de te rmined by the intersect ion of the t ipping axis with the plane of ro ta t ion . T o find the equat ion of this t ipping plane, let T be a vector from the CGTP to any general point (x, y , z) in the t ipping plane (see circled i tem 2) . T = ( x - x t P ) * + (y - y t p ) j + ( z - z t P ) k [2] Since the vector T is in a plane perpen dicular t o the t ipping axis, it follows tha t j and T are perpendicular to each o ther so that their do t (or scalar) prod uct (j \u00b0 T) is zero" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002032_neco.2007.19.3.730-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002032_neco.2007.19.3.730-Figure2-1.png", "caption": "Figure 2: Single pendulum swing-up task. \u03b8 : joint angle, T : input torque, l: length of the pendulum, m: mass of the pendulum.", "texts": [ " We could simultaneously acquire both control and estimation policies for the biped walking task. 4.1 Application to a Single Pendulum. The proposed method was applied to pendulum dynamics, ml2\u03c9\u0307 = \u2212\u00b5\u03c9 \u2212 mgl sin \u03b8 + T, (4.1) where \u03b8 is the angle from the bottom position, \u03c9 is the angular velocity, T is the input torque, \u00b5 = 0.01 is the coefficient of friction, m = 1.0 kg is the weight of the pendulum, l = 1.0 m is the length of the pendulum, and g = 9.8 m/s2 is the acceleration due to gravity (see Figure 2). We define the state vector as x = (\u03b8, \u03c9)T and the control output as u = T . Here we consider one-dimensional output y. Thus, the dynamics is given by the system x(k + 1) = f(x(k), u(k)) + n(k), (4.2) y(k) = Cx(k) + v(k), (4.3) where n(k) and v(k) are noise inputs with parameters = diag{0.01, 0.01} t and R = 1.0 t, and the discrete-time dynamics of the pendulum is f(x(k), u(k)) = x(k) + \u222b (k+1) t k t F(x(s), uc(s))ds, (4.4) F(x(t), uc(t)) = ( \u03c9 \u2212 g l sin \u03b8 \u2212 \u00b5 ml2 \u03c9 ) + ( 0 1 ml2 ) uc, (4.5) where t is a time step of the discrete-time system, and uc is the control input in continuous time" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001925_0954406jmes476-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001925_0954406jmes476-Figure1-1.png", "caption": "Fig. 1 Friction force and sliding velocities acting on the area dA of an elliptical contact surface", "texts": [ " 221 Part C: J. Mechanical Engineering Science at University of Bath - The Library on June 11, 2015pic.sagepub.comDownloaded from A process for evaluating the heat generated from the ball bearing is developed in this paper. Then a two-dimensional model for evaluating the temperature of the ball bearing has been set up, assuming that the heat source is moving. The model is more consistent with the actual working condition so that it can well compute the heat generated and the temperature of the ball bearing. Figure 1 shows the friction force and sliding velocities acting on area dA of an elliptical contact surface. The various stress, force, and moment can be written as [12] \u03c4 = 3\u00b5Q 2\u03c0ab ( 1 \u2212 (x a )2 \u2212 ( y b )2 )0.5 (1) dF = \u03c4 dA (2) F xmj = \u222b \u03c4xmj dA (3) F ymj = \u222b \u03c4ymj dA (4) dMs = r cos(\u03c6 \u2212 \u03b8)dF\u0304 (5) Ms = 3\u00b5Q 2\u03c0ab \u222b+a \u2212a \u222b+b(1\u2212(x/a)2)0.5 \u2212b(1\u2212(x/a)2)0.5 (x2 + y2)0.5 ( 1 \u2212 (x a )2 \u2212 ( y b )2 )0.5 cos(\u03c6 \u2212 \u03b8)dy dx (6) F\u0304dj = \u03c0\u03beCvD2(dm\u03c9oj) 1.95 32 g (7) F\u0304CL = \u03b7 \u03c0\u03c9CRcmdCR(\u03c9c \u2212 \u03c9m) 1 \u2212 (d1/d2) (8) The heat generated is derived from the friction in the high-speed ball bearing" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003140_iros.2009.5354270-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003140_iros.2009.5354270-Figure3-1.png", "caption": "Fig. 3. Calculation method of the robotic swing angle.", "texts": [ " Similarly, the robotic shoulder comprises free joints, but in this study, two constraint components are attached to restrain the rotation at the shoulder in order to simplify the learning. On the back of the robot, a communication module with Bluetooth is attached as shown in Fig. 2 (b). By using this module, motor commands are transmitted via a radio communication based on RS-232C. As for the measurement system, a PSD-sensor system (model: C5949, Hamamatsu Photonics) is employed in order to obtain the robotic position. Using measured position data, the robotic swing angle can be calculated, as shown in Fig. 3. In this experimental system, the sampling rate is set to 250 ms in order to completely drive the motors to the desired position within a sampling step. B. Dynamics Simulator This study also created a dynamics simulator of the developed compact humanoid robot by using Open Dynamics Engine, a free physical-calculation simulator. Fig. 4 shows a three-dimensional graphics of the robot in the developed simulator. In this simulator, the robotic parameters, such as the motor characteristics and damping effect generated by the friction between the horizontal bar and arm components, are adjusted to those of the real robot as far as possible" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002504_robio.2009.4913206-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002504_robio.2009.4913206-Figure1-1.png", "caption": "Fig. 1 Developed BWS system with pelvic support (a), Manufactured seat (b), Ball joint with plate (c)", "texts": [ " INTRODUCTION Recently, the number of bedridden elderly people in society is increasing, and they need care givers and even robots that can help them lead independent lives [1][2]. Walking is a basic essential ability for living independent life and for maintaining physical and mental health. Since any mechanism that supports body weight is useful, several such mechanisms that support various parts of human body have been developed (see Table I). These mechanisms have a disadvantage from the viewpoint of pain and numbness. However, our mechanism (see Fig. 1) includes the painless support by the ischium. In addition, this support also allows a small mechanism that unloads body weight from below, unlike other systems that lift up from above. By integrating our BWS system with a mobile base system, we plan to develop a walking support robot [3]. Compared with other support systems, it is difficult to support by the ischium; thus, both a suitable supporting mechanism attached to the end of the arm of a motor actuated device and a force controller for BWS need to be considered. Previously we developed a prototype of a pelvic support mechanism. The mechanism has a seat with a surface shape of hip to hold pelvic posture (see Fig. 1 b) and a ball joint with a range of movement limited to 2 passive DOF by the plate, so that the range of pelvic movement can be adjusted (see Fig. 1 c), taking into consideration the ease of walking. The pelvic support mechanism has been reported in detail [4]. In this paper, we focus only on the force control we have adopted for our BWS system. Many different methods of providing unloading force exist. According to previous study [5], many BWS systems which is available on the market use a winch [6], counter weight [7], or elastic spring [8], and lift up body weight from above via a harness connected to a wire (Fig. 2 a). Lokolift is the system that is composed of a passive elastic spring element to take over the main unloading force and an active closed-loop controlled electric drive to generate the exact target unloading force [5] (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003074_j.engfailanal.2010.03.009-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003074_j.engfailanal.2010.03.009-Figure8-1.png", "caption": "Fig. 8. Possible vibration modes.", "texts": [ " One possible explanation for the impeller failure was that the blade with a defective weld fatigued due to resonance, then severe vibrations occurred due to imbalance after the blade was thrown, which in turn resulted in other cracks being formed. The root cause analysis needed to be continued since a source of stress to drive fatigue crack growth in the thrown blade had not been identified. In seeking a source of stress, two vibration modes were conceptually identified: a blade \u2018flapping\u2019 mode and a torsional mode (see Fig. 8). Thus, the first step was to calculate and measure the natural frequencies for structural modes of vibration of the blades, and the torsional frequency for the impeller (front ring relative to back plate). In addition, an impulse response procedure was used to confirm the calculated values. Initial calculations indicated the three lowest \u2018\u2018flapping\u201d modes were 438, 567 and 1037 Hz. These compared well with the measured modes of 422, 586 and 1055 Hz respectively. The torsional mode for the whole impeller was calculated approximately as 71 Hz" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002815_1.4002527-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002815_1.4002527-Figure7-1.png", "caption": "Fig. 7 Shaft element", "texts": [ " Assuming that damping of the disk is of a viscous kind, the equations of motion of the disk Eqs. 23 \u2013 26 can be rewritten into a matrix form MDEx\u0308DE + BDE + GDE x\u0307DE + GDExDE = fDE 27 where MDE, GDE, and BDE are the mass, gyroscopic, and external damping matrices of the disk, fDE is the vector of general forces constraint and applied acting on the liberated disk, and x\u0307DE and x\u0308DE are vectors of the disk general velocities and accelerations. 3 The Motion Equation of a Shaft Element The shaft finite Fig. 7 element has two nodes and eight degrees of freedom two displacements in y and z directions, two rotations about axes y and z at each node . It can be deformed by deflection in two mutually perpendicular planes and in addition, it turns about its center line. In the following text, the geometric and Transactions of the ASME 6 Terms of Use: http://www.asme.org/about-asme/terms-of-use m o J r a d a a w s p a A d w t a o w e n y a t t t b a J Downloaded Fr aterial parameters of the shaft element are noted: L is the length f the element, S is the area of the cross section of the element, SD and JSP are the quadratic moments of the element cross section elative to its diameter and to its axis of rotation, and E, , and V re Young\u2019s modulus, density, and the coefficient of viscous amping of the element material, respectively. Deflection of the center line of the element in the directions of xes y and z of the element fixed frame of reference Fig. 7 are pproximated by the polynomial shape functions y = a0y + a1yx + a2yx2 + a3yx3,z = a0z + a1zx + a2zx 2 + a3zx 3 28 here x is the x-coordinate of the point on the center line of the haft element, y and z are the deflections displacements of the oints on the shaft center line in directions y and z, and aiy and aiz re the coefficients of the shape functions i=0 ,1 ,2 ,3 . Let the following column matrices be introduced a = a0y,a1y,a2y,a3y,a0z,a1z,a2z,a3z T 29 py x = 1,x,x2,x3,0,0,0,0 T, pz x = 0,0,0,0,1,x,x2,x3 T 30 ry x = 0,1,2x,3x2,0,0,0,0 T, rz x = 0,0,0,0,0,1,2x,3x2 T 31 fter expressing the coefficient vector a by means of the nodal eformation parameters a = P\u22121xHE 32 here xHE = y1,z1, y1, z1,y2,z2, y2, z2 T 33 P = py 0 ,pz 0 ,\u2212 ry 0 ,rz 0 ,py L ,pz L ,\u2212 ry L ,rz L T 34 he relations for displacements y and z and rotations y and z of ny cross section of the shaft element along or about axes y and z f its fixed frame of reference take the forms y x = py T x P\u22121xHE, z x = pz T x P\u22121xHE 35 y x = \u2212 rz T x P\u22121xHE, z x = ry T x P\u22121xHE 36 here xHE is the vector of general displacements of the shaft lement, y1 and z1 are the displacements in y and z directions at ode 1, y1 and z1 are the rotations about axes y and z at node 1, 2 and z2 are the displacements in directions y and z at node 2, nd y2 and z2 are the rotations about axes y and z at node 2. Lagrange equations of the second kind are used for the derivaion of the equation of motion. Let a small cylinder of an infiniesimal thickness be specified by two section planes perpendicular o the center line of the shaft element Fig. 7 . This cylinder can e considered as an infinitely thin circular disk. Moreover, it is ssumed that its center of mass is slightly shifted from the shaft ournal of Applied Mechanics om: http://appliedmechanics.asmedigitalcollection.asme.org/ on 01/28/201 center line. During vibration, the elementary disk performs a sliding motion given by the movement of the disk center of gravity and a spherical motion about its center of gravity. Its position is then defined by displacements of the center of gravity yT and zT and by two Euler\u2019s angles 1 and 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002783_j.electacta.2010.12.020-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002783_j.electacta.2010.12.020-Figure1-1.png", "caption": "Fig. 1. Redox and protonation equilib", "texts": [ " Coupled Raman\u2013EIS data were obtained using a SOTELEM potentiostat and a home-made numerical Frequency Response Analyzer described in [23] allowing up to eight independent channel inputs. In both cases, the signal amplitude was 30 mV peak-to-peak for which linearity condition was confirmed. 3. Results and discussion 3.1. Mechanism implying 3 species and 2 electrochemical reactions In the pH domain covered in this study, and considering that overoxidation is excluded as experimentally verified, the most complete reaction scheme ascribable to different proton and electron exchanges for PANI is displayed in Fig. 1 where A2\u2212 are the inserted or expelled counter ions corresponding to sulphate to maintain electroneutrality in the film. The two polaronic or bipolaronic forms for the oxidized and protonated state (ES) were ria be a w b p i t t h l i t b w p t L L E I e c c c i b t i a xp[2 lso considered. The divalent character of the counter ion involves eaker coulombic interaction with polarons of the same chain or ipolarons and entails the possibility of pairing with interchains olarons. This was suggested in the positioning of A2\u2212 counter ions n Fig. 1. According to this scheme, all reactions involved in Fig. 1 with he two electrochemical reactions and the chemical one have been aken into account to predict the different transfer functions. One can write the possible kinetic reactions with the following ypotheses: the film is assumed to be infinitely thin in direct equi- ibrium with the solution and by neglecting the ions diffusions both n the solution and in the film. The protons and counter ion concenrations (here sulphate) will thus be set equal to their values in the ulk solution and so they will contain not modulated component ith potential" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002239_j.euromechsol.2008.06.008-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002239_j.euromechsol.2008.06.008-Figure8-1.png", "caption": "Fig. 8. Planar 1-link WMM.", "texts": [ " The final values of N p and Na are retained to initiate the stochastic process. Fig. 7(d) gives an example showing a trial path and the final path obtained after convergence of the process. Although initially a trial path might present undesirable distortions, they will be attenuated progressively as the process converges. We now give some numerical results obtained on a 1.6 GHz Intel Centrino Duo computer. The first example compares the pro- posed RPA to another method based on PMP, using a 1-link WMM (see Table 1 and Fig. 8). Most of the other examples will concern a planar 2-link WMM (see Table 2 and Fig. 9). The last example will deal with a spatial 3-link WMM. Unless otherwise specified, we have made use of a clamped cubic-spline model with Nm = 3 free nodes for the motion function \u03bb(\u03be) and a fourth-order B-spline model, respectively, with N p = 7 free nodes for the platform path P p(\u03bb) and Na = 5 free nodes for the manipulator path P a(\u03bb). Parameters of the simulatedannealing algorithm have been set as given in Haddad et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002020_tmag.2008.2002613-Figure13-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002020_tmag.2008.2002613-Figure13-1.png", "caption": "Fig. 13. Slider dynamic modulation as measured by LDV when flying over the e-beam spoke pattern at different TFC clearances. Dynamic modulation of 2.5 nm is measured.", "texts": [ " A smooth transition from the large optical patterns (micrometer scale) to the e-beam pattern is desired for good recording studies at various recording densities. Our initial approach was to minimize a density mismatch between the large- and the small-scale pattern. Fig. 12 shows the clearance change with pattern density as measured by pulse thermal protrusion. Each data point is obtained on different media with the same slider repeatedly. Again, the slope (similar to Fig. 9) is proportional to the hole depth (42 nm). In Fig. 13, the actual slider dynamic modulation (as measured by LDV on the slider trailing edge) is shown when flying from the large scale (optically patterned media) across the small e-beam spoke pattern. The large-scale pattern density varies and, hence, the FH is different on each of the three media. The LDV maps shown in Figs. 13 and 14 were measured without TFC applied for the three different disks shown in Fig. 12. The radii (tracks) were scanned at steps of 38 m ranging from 20 to 28 mm at 7200 r/min" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.64-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.64-1.png", "caption": "Fig. 14.64 Two-mast material and equipment lift with rack-and-pinion hoisting gear", "texts": [], "surrounding_texts": [ "Table 14.3 Specifications of model ladder lifts\nParameters\nLifting capacity (kg) 150\u2013200\nLength of track (m) 13\u201330\nSpeed of platform (m/min) 18\u201340 Guide bend angle adjustment (\u25e6) 20\u201345; in some ladder lifts: 0\u201370\nNumber of guide sections 9 to 15\nElectric supply (V) 230 single-phase alternating current\nMotor power (kW) 0.75\u20131.5\nControl Electrical \u2013 through control buttons\nMass of drive unit (kg) 40\u201355 Mass of guide section (kg) \u2248 10\nTotal mass of ladder lift (kg) 160\u2013210\nMast Material and Equipment Lifts with Rack-and-Pinion Hoisting Gear\nIn material and equipment lifts with a rack-and-pinion hoisting gear the platform used for transporting materials climbs a rack running along the mast.\nThe lift consists of a platform and a mast with a base. The mast structure is segmental. The individual segments have the form of a cuboidal truss structure with a rack attached to one side. Thanks to its segmental structure and the use of appropriate assembly accessories the lift is a self-assembly device. The mast is extended by adding and joining mast sections. The mast is anchored, usually by means of tubular elements, enabling the adjustment of its position relative to the wall of a building. The drive unit mounted on the platform\nBucket with tipping device\nBucket with side discharge\nBox with guards for transporting bricks\nLoad box with hinged side guards\nPlatform for ransporting panels\nFig. 14.62 Load fastening accessories\ntypically consists of an electric motor with a builtin electromagnetically released spring brake, a flexible clutch, a worm gear, and a rack wheel that mates with the mast\u2019s rack. In the case of two-mast lifts, two drive units are employed. Examples of material lifts with a platform raising rack-and-pinion hoisting gear are shown in Figs. 14.63, 14.64 and Table 14.4.\nThe handled materials are loaded and unloaded at stops with transport stages. The mast\u2019s base is enclosed by fencing with a control box attached to it. There is an electrically controlled pivoting gate in the fencing. On the entry side all platform stops have sliding barriers which may be equipped with an electric barrieropening monitoring system for additional protection, so that persons at the stop cannot fall out. Similarly as ca-\nPart B\n1 4 .4", "ble lifts, lifts with a rack-and-pinion lifting gear may be equipped with a set of wheels to enable them to be easily transported to a new work site. Instead of the platform for handling lump and sacked materials, a special bucket for transporting concrete mix can be installed.\nMultisectional mast\nLift\u2019s base\nPlatform with drive unit\nAnchoring connectors\nFencing\nFig. 14.63 Single-mast material and equipment lift with rack-and-pinion hoisting gear\nA comparison of lifts with a rack-and-pinion hoisting gear with lifts with a cable hoisting gear shows that the former lifts have several advantageous properties such as: greater lifting height, easier and safer assembly, and simpler operation.\nPart B\n1 4 .4", "Shaft Material and Equipment Lifts Shaft material and equipment lifts (Fig. 14.65) are used for the vertical transport during the construction of medium- and high-rise building structures.\nA shaft lift consists of the following main parts (Fig. 14.65 and Table 14.5):\n\u2022 A shaft\u2022 An upper beam\u2022 Guides\u2022 A bottom cable pulley\u2022 The platform\u2019s upper beam\u2022 A platform\u2022 A winch (typical lifting winches with an electric or diesel drive can be used)\nThe shaft has a spatial truss structure. The load-bearing platform is made of steel sections. The cable is guided by the bottom and top cable pulleys. The end of the cable is fixed to the upper beam (Fig. 14.65).\nThere is also a shaft lift design in which the shaft is made of only flat frames anchored to the building\u2019s wall. Shaft lifts are equipped with similar safety devices as other material lifts.\nMaterial lifts are equipped with the following safety devices:\n\u2022 A gripping device which stops the platform as it descends whenever it exceeds the maximum allowable rate of descent.\u2022 Protection against disengagement of the drive wheel from the mast\u2019s rack. As standard, sliding guides are used. They keep the load platform on the mast even if the roller guides fail.\u2022 An emergency lowering system used in the case of a prolonged power failure. Some lifts are equipped\nwith emergency lowering systems with speed selfstabilization \u2013 the speed stabilizes below the speed at which the gripping device is actuated.\u2022 The upper and lower limit switches, automatically stopping the platform at the mast\u2019s highest and lowest levels.\u2022 Switches and locks for stop doors or barriers, preventing their accidental opening when the platform is outside the stop zone or in motion.\u2022 Stops to ensure that the platform will be brought to a stop if the limit switches fail.\u2022 An induction sensor that monitors mast presence during mast assembly.\u2022 A sound system signalling the start of a platform ride.\u2022 Protection against electric shock.\u2022 Overload protection of the electric motors.\u2022 Switches actuated when the working platform skews in two-mast lifts.\nThe operation of the cable-driven material lifts described above typically consists of the control of the movement of the carriage by pushing buttons on the control panel at the lower station. It is also possible to switch to control from the platform during assembly and maintenance of the lift.\nMaterial and equipment lifts with access to personnel are intended for the vertical transport of persons and materials during construction/assembly works and repairs of mainly high-rise buildings in housing and industrial construction. Their design is usually similar to that of material and equipment lifts with a rack-andpinion hoisting gear.\nA person and material lift consists of a cabin with a rack-and-pinion drive, moving on a mast secured at the bottom to the lift\u2019s base and anchored to the building\u2019s wall, and transport stages (stops) between which transport takes place. The mast has a segmental structure and can be extended by adding mast sections. It is anchored by means of a system of tubes, which makes it possible to adjust the mast\u2019s position relative to the building\u2019s wall.\nThe lift can have two cabins, each with its own drive system, whereby the transport of persons and materials can be doubled. The lift\u2019s cabins move on a common mast independently of each other. Examples of person\nPart B\n1 4 .4" ] }, { "image_filename": "designv11_20_0002461_s11633-008-0138-4-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002461_s11633-008-0138-4-Figure1-1.png", "caption": "Fig. 1 A simple network", "texts": [ " In the computing procedure, c2 = 1, the step length for \u03b1 is 0.01, iteration number is 2 461 and processing time is 127.8280 seconds. Let c1 = 0.5 and c2 = 2. Performing Algorithm 2 using 1) of the LMI feasibility problem, we found that the system is regular, impulsive-free, and PS with respect to (c1, c2, d), for a maximum dmax = 1.9928, obtained for \u03b1 = 1.0. In the computing procedure, d = 4, the step length for \u03b1 is 0.01, iteration number is 2 518 and processing time is 147.0460 seconds. Consider a simple circuit network shown in Fig. 1, where the voltage source VS(t) is the control input, R, L, and C stand for the resistor, inductor, and capacity, respectively, and their voltages are denoted by VR(t), VL(t), and VC(t), respectively. Assume that the voltage source is perturbed by some exogenous disturbance \u03c9 = 0.3 sin(t). Choose the state vector and the input vector as follows: x = [ I(t) VL(t) VC(t) VR(t) ]T , and u = VS(t). Let R = 1, L = 1, and C = 1. Then, the system can be described in the form of the linear descriptor system (17), with E = \u23a1\u23a2\u23a2\u23a2\u23a3 1 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 \u23a4\u23a5\u23a5\u23a5\u23a6 , A = \u23a1\u23a2\u23a2\u23a2\u23a3 0 1 0 0 1 0 0 0 \u22121 0 0 1 0 1 1 1 \u23a4\u23a5\u23a5\u23a5\u23a6 B = G = [ 0 0 0 1 ]T " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001071_3.58552-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001071_3.58552-Figure1-1.png", "caption": "Fig. 1 Engine cross section.", "texts": [ " 3) Aeroelastic instability\u2014mechanical deflections produce a pressure perturbation that, in turn, reinforces the mechanical deflection. The interaction persists at a mechanical natural frequency which has a deflected shape (mode shape) most sympathetic with the aeroelastic coupling. This discussion concentrates on the aeroelastic instability phenomena and relates a recent example confronted in the F100 turbofan engine. The seal on which the aeroelastic instability encountered is at the high compressor discharge of the F100, a high thrust-toweight turbofan engine (Fig. 1). The knife-edges are on the seal rotor, the land is stationary. The land surfaces are faced with an abradable material. The seal was located remote from D ow nl oa de d by U N IV E R SI T Y O F M IC H IG A N o n M ar ch 4 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /3 .5 85 52 JULY 1979 AEROELASTIC INSTABILITY IN F100 LABYRINTH AIR SEALS 485 the compressor discharge so the radius of the seal and land are relatively small to minimize leakage. The knife-edged rotor was made to be removed and refurbished at minimum cost, so rather than being integral with the compressor rear hub, it is a thin conical shell secured to the hub radially by snaps and axially by a stack-up secured by the compressor aft bearing nut. The seal land is a conical forging attached to the i.d. of the engine burner case (Fig. 1). Engine Operation Affects Vibration Characteristics The vibratory characteristics of the thin knife-edge shell are normally dominated by the more massive compressor hub. An exception occurs during a maximum rotor speed transient from idle to maximum speed. The knife-edge shell is heated by compressor discharge air. The hub is cooled on the i.d. by bleed from a forward snap on the knife-edge shell to grow temporarily loose from the hub. During this brief interval, a maximum of 30 s, the fixity of the knife-edge shell is altered, permitting a unique family of vibratory shell modes to occur" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002882_1077546309349906-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002882_1077546309349906-Figure1-1.png", "caption": "Figure 1. Journal bearing configuration.", "texts": [ " The following dimensionless parameters are defined: X \u00bc x L , Y \u00bc y L , H \u00bc h hm , P \u00bc p p0 , Lt \u00bc 12ZoL2 h2 mpo , Lx \u00bc 6Zu1L h2 mpo , Ly \u00bc 6Zw1L h2 mpo , ~Q \u00bc 12ZL2pa h3 mp2 0 ra r~v Adopting the assumptions of an adiabatic process and a non-viscous flow, it can be shown that the mass flow rate is given by: _m \u00bc Ap0f ffiffiffiffiffiffiffi 2r0 p0 s j \u00f03\u00de j \u00bc k 2 2 k\u00fe1 \u00f0k\u00fe1\u00de=\u00f0k 1\u00de 1=2 ; p p0 bk k k 1 p p0 2=k p p0 \u00f0k\u00fe1\u00de=k 1=2 ; p p0 > bk 8>>< >>: \u00f04\u00de where A is the cross-sectional area of the orifice, k is the ratio of the specific heat, p0 is the supplied pressure, f is the coefficient of the mass flow rate through the orifice, and bk \u00bc pc p0 \u00bc 2 k\u00fe1 k=\u00f0k 1\u00de . Figure 1 presents the gas journal bearing configuration considered in the present study. It is observed that two sets of eight orifices are arranged evenly around the circumference of the bearing at quarter-station positions. The boundary conditions are shown as follows: 1. The atmosphere boundary condition: P \u00bc pa p0 2. The periodic boundary condition: P\u00f0Z\u00de \u00bc P\u00f0Z \u00fe 2p\u00de 3. The symmetric boundary condition: qP qX \u00bc 0 Two different hybrid methods are applied to determine the overall gas pressure, and the load capacity of the bearing can then be established by means of force balance equations" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001064_j.conengprac.2006.05.002-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001064_j.conengprac.2006.05.002-Figure2-1.png", "caption": "Fig. 2. Cross-section of Fig. 1 in the x\u2013y plane. Motor coupling not shown.", "texts": [ " While modeling of AMBs is relatively established (Schweitzer et al., 1994), their dynamics are briefly reviewed here to establish notation which is used in control design. The system has five DOF and consists of a horizontal shaft onto which two journals and a disk are mounted. The journals form part of the radial magnetic bearings, and the disk is part of the axial magnetic bearing. The entire shaft assembly is coupled to a DC motor via a helical coupling. Fig. 1 shows the configuration of the shaft and magnetic bearings. Fig. 2 shows a schematic of a crosssection of the system. Assuming a rigid shaft the dynamic equations are (Matsumura & Yoshimoto, 1986): m \u20acx \u00bc Fx, m \u20acy \u00bc Fb;y \u00fe F f ;y \u00fe Fc;y \u00femgy, m\u20acz \u00bc Fb;z \u00fe Ff ;z \u00fe F c;z \u00femgz, Jz \u20acc \u00bc \u00f0lf ;a \u00fe x\u00deF f ;z \u00f0lb;a x\u00deFb;z Jxo_y\u00fe lcFc;z, Jy \u20acy \u00bc \u00f0lb;a x\u00deFb;y \u00f0lf ;a \u00fe x\u00deFf ;y \u00fe Jxo _c lcF c;y, (2) T.R. Grochmal, A.F. Lynch / Control Engineering Practice 15 (2007) 95\u2013107 97 where x; y; z denote the coordinates of the center of mass cm relative to the origin O of the inertial frame", " The integral in (7) is used to compensate for static offset that may result from constant disturbances (Mizuno & Bleuler, 1995) or static load change (Flowers, Sza\u0301sz, Trent, & Greene, 2001). Similar expressions to (7) for vy; vz; vc; vy are used to stabilize the tracking errors in y; z;c; y. The tracking control law requires knowledge of x; y; z;c; y, their derivatives, and integrals. The axial displacement x is measured and y; z;c; y are obtained indirectly from measurement of the shaft displacement in the sensor planes x \u00bc lb;s and x \u00bc lf ;s (see Fig. 2). These measurements are denoted V 13;W 13;V24, and W 24. The approximate relations between x; y; z;c; y and the measurements V 13;W 13;V 24;W 24 are W 24 \u00bc y\u00fe \u00f0lb;s x\u00dey; V 24 \u00bc z \u00f0lb;s x\u00dec, W 13 \u00bc y \u00f0lf ;s \u00fe x\u00dey; V 13 \u00bc z\u00fe \u00f0lf ;s \u00fe x\u00dec. \u00f08\u00de Solving (8) for y; z;c; y gives y \u00bc \u00f0W 24 W 13\u00de lb;s \u00fe lf ;s ; c \u00bc \u00f0V13 V24\u00de lb;s \u00fe lf ;s , y \u00bcW 13 \u00fe \u00f0lf ;s \u00fe x\u00dey; z \u00bc V24 \u00fe \u00f0lb;s x\u00dec. \u00f09\u00de For implementation the velocities _x; _y; _z; _c; _y are approximated by numerical differentiation, and integration is approximated by numerical quadrature" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001571_s1560354707030045-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001571_s1560354707030045-Figure1-1.png", "caption": "Fig. 1. The Snakeboard.", "texts": [], "surrounding_texts": [ "MSC2000 numbers: 70F25, 70E55, 70E60, 70E18 DOI: 10.1134/S1560354707030045\nKey words: Snakeboard, Gibbs\u2013Appell equations, dynamics, analysis of motion\n1. INTRODUCTION\nThe Snakeboard is one of the modifications of a well-known skateboard. It allows the rider to propel himself forward without having to make contact with the ground even if a motion occurs uphill. The motion of the snakeboard becomes possible due to a specific features of its construction and due to the special coordinated motions of legs and a torso of the rider. The first snakeboard has appeared in 1989 and from this moment till now it has found a lot of fans among the amateurs of extreme sports. Soon after the invention of the snakeboard the first attempts were made to give a mathematical description of the basic principles of human snakeboarding. The basic mathematical model for the snakeboard investigated by various methods in many papers [1\u201313] was proposed by Lewis et al. [1]. In our paper we give the further development of the model proposed in [1].\nThe Snakeboard (see Figs. 1\u20132) consists of two wheel-based platforms upon which the rider is to place each of his feet. These platforms are connected by a rigid crossbar with hinges at each platform to allow rotation about the vertical axis. To propel the snakeboard the rider first turns both of his feet\n*E-mail: kuleshov@mech.math.msu.su\n321", "in (see Fig. 3). By moving his torso through an angle, the snakeboard moves through an arc defined by the wheel angles. The rider then turns both feet so that they point out, and moves his torso in the opposite direction. By continuing this process the snakeboard may be propelled in the forward direction without the rider having to touch the ground.\nThe mathematical model of the snakeboard considered in this paper is represented in Fig. 4. We assume that the snakeboard moves on the xy plane, and let Oxy be a fixed coordinate system with origin at any point of this plane. Let x and y be the coordinates of the system center of mass (point G) and \u03b8 the angle between the central line of the snakeboard and the Ox-axis. In the basic model treated in [1] platforms could rotate through the same angle in opposite directions with respect to a central line of the snakeboard (by other words, for this model \u03d5f = \u2212\u03d5b = \u03d5, see Figs. 3\u20134). We suppose that platforms can rotate independently and their positions are defined by two independent variables \u03d5f and \u03d5b (Fig. 4). The motion of the rider is modeled by a rotor, represented in the form of a dumb-bell in Fig. 4. Its angle of rotation with respect to the crossbar is denoted by \u03b4.\nREGULAR AND CHAOTIC DYNAMICS Vol. 12 No. 3 2007" ] }, { "image_filename": "designv11_20_0002358_jahs.53.282-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002358_jahs.53.282-Figure2-1.png", "caption": "Fig. 2. Test gears.", "texts": [ " The nominal oil supply pressure was 80 psi, and the nominal flow rate was 1.0 gpm for both the test section and slave section. Oil inlet temperature was set at 85\u25e6F. An external vacuum pump connected to the oil tank worked as a scavenge system to remove the oil from the test gearboxes and bearing cavities and direct it to the sump. Also, the oil system was equipped with an oil-debris monitor as well as a 3-\u00b5m filter. The design parameters for the pinions and face gears used in the tests are given in Table 1. A photograph of the test specimens is shown in Fig. 2. The set was primarily designed to fail in surface pitting fatigue mode. The set had a reduction ratio of 3.842:1. The set also had a diametral pitch of 10.6 teeth/inch, roughly similar to the previous TRP design and current AH-64 replacement design. The face width of the face gears was 0.6 inch. The face width of the spur pinions was 0.8 inch, significantly greater than the face gear to allow for backlash adjustment and optimization of tooth contact. The shaft angle was 90 deg to accommodate the facility" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002480_acc.2007.4282986-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002480_acc.2007.4282986-Figure3-1.png", "caption": "Fig. 3. Control strategy to capture a target.", "texts": [ " In this section, we focus on a simple yet effective strategy to capture a full-observed target. Several control algorithms can be implemented such as a proportional controller and a more sophisticated leaderfollower approach. In the latter case the target to capture is seen as the leader to follow. We describe an effective approach that allows a pursuer to capture a target along a trajectory of minimal length [8]. We assume that the target is moving in a straight line with constant velocity and intercepting along that line. This strategy depicted in Fig. 3 is based upon the geometry of the problem and taking into account the kinematic constraints of the pursuer. The strategy attempts to intercept the target at a point \u03b4. The interception point is calculated by determining the time for both the pursuer and target to reach that point. The initial states of the pursuer and target are p0 = (xp, yp, \u03b8p) and \u03c40 = (x\u03c4 , y\u03c4 , \u03b8\u03c4 ), respectively. The interception point is defined as \u03b4 = [ xt + tcvt cos \u03b8t yt + tcvt sin \u03b8t ] and the time to interception becomes tc = r\u03c8 + \u2016c\u2212 \u03b4\u2016 cos\u03b1 vp , (20) where the distance traveled by the pursuer is the distance along the arc p0p1 plus the straight line distance between p1 and \u03b4" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002374_12.817405-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002374_12.817405-Figure4-1.png", "caption": "Fig. 4 \u2013 Finite element mesh. If deposition takes place in the mid-plane of the substrate, a symmetry plane exists and the problem is solved in half the domain, taking a zero heat flux through that plane.", "texts": [ " The model was used to study the influence of the deposition parameters on the microstructure, hardness, Young\u2019s modulus and residual stress distributions in Ti-6Al-4V walls (width = 0.50 mm; length = 2.50 mm; height = 1.25 mm) produced by overlapping 10 layers of Ti-6Al-4V using the same material as substrate. The deposition was assumed to take place in the midplane of the substrate, so that the problem is symmetrical in relation to this plane, and to reduce computational time only half of the geometry was considered (Fig. 4). The deposition parameters used in the simulation were: scanning speed = 2.5 mm/s, laser beam power = 725 W, laser beam radius = 1.5 mm, idle time between the deposition of consecutive layers = 1 s and initial substrate temperature = 27 \u00baC. The distribution of phases along the wall height middle line is shown in Fig. 5, and the distribution of Young\u2019s modulus and hardness in Fig. 6. The final part presents a non-uniform distribution of properties which is a consequence of the non-uniform microstructure in the deposited material" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002115_s0070-2153(07)81008-8-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002115_s0070-2153(07)81008-8-Figure10-1.png", "caption": "Figure 10 Branching modeled by Stokes fluids deformed by a fixed force. Epithelium and mesenchyme each have constant viscosity. Growth is omitted. Deformations depend on viscosity ratio, surface tension, and clefting force (arrows). Results suggest that forces from epithelium alone are insufficient to generate observed shape changes when epithelium is embedded in mesenchyme. Figure schematic after simulations from (Lubkin and Li, 2002).", "texts": [ " In the interests of parsimony and clarity, we focused solely on the tissues\u2019 mechanical response to forces, and ignored many of the known features of branching systems, such as growth. Since branching morphogenesis is robust, it is fair to assume that the effect of forces far from the epithelio-mesenchymal interface is small; we used periodic boundary conditions on the outside of the mesenchyme/culture medium. Finally, the clefting force is input to the model as a point force at a specified value in specific locations and held constant in magnitude and direction for the duration of clefting (Fig. 10). This model demonstrated several mechanical features: 1. The clefting force strongly affects the time course of branching morphogenesis. Decreasing the clefting force by half an order of magnitude increases the clefting time by 2 orders of magnitude. 2. Mesenchymal viscosity strongly affects the time course of branching morphogenesis. Increasing the viscosity of the material the epithelium is embedded in by an order of magnitude increases the clefting time by half an order of magnitude. 3. The ratio of viscosities of epithelium and mesenchyme strongly affects morphology of the clefts, with more viscous mesenchyme yielding wider clefts" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002904_1077546309104878-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002904_1077546309104878-Figure6-1.png", "caption": "Figure 6. The change of .", "texts": [ " We observe that the mass M follows the orbit determined above in Section 2.1.2, while the mass mo, due to the system of the spring/damper, oscillates with respect to M (see Figure 5). The above-mentioned oscillation is describedby the following equation: mo m ko m co m 0 (24) which has the solution m t e ot D1 sin ot D2 cos ot (25) where o o and o are given by equations 5 and 6. At instant t (loss of touch), the normal distance between M and mo changes to m and the following equation is valid (see Figure 6): z m o or, finally, m o z (26) Thus, the initial conditions for the oscillation of mo are m 0 x t o x z t m 0 x t z t (27) which finally give at SIMON FRASER LIBRARY on November 17, 2014jvc.sagepub.comDownloaded from D1 o x t x t o o x oz t z t o D2 x t o x z t (28) Therefore, at the instant of impact the velocity of mo because of its oscillation will be m t e ot [ o D1 o D2 sin ot o D1 o D2 cos ot ] (29) 3. The third component of the impact velocity, , is the velocity of the vibrating beam x t " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002027_978-1-4020-4535-6_27-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002027_978-1-4020-4535-6_27-Figure3-1.png", "caption": "Figure 3. 3D-mesh for one pole of the AFPM machine. (a) Stator with armature winding (b) Rotor.", "texts": [ " Drawbacks However, this configuration has some problems that can be summarized as follow: Lower power density respect to regular PM machine due to the reduction on the amount of magnet. gsymmetrical flux density distribution over the stator teeth introduces additional saturation over the stator and rotor yoke. This is because of their trapezoidal shape. In order to determine the effectiveness of the proposed configuration a 3D-FEA is carried out. Rotor and stator domain and 3D-mesh used to evaluate the topology is shown in Fig. 3. Stator winding representation is depicted. One detail has to be incorporated in the model. Lamination core for the magnetic circuit has the property to carry flux mainly in tangential and axial directions. However, due to the interlamination airgap radial flux is reduced considerably. To take into account this effect in the model, additional radial airgaps are introduced in the 3D-FE model. In this manner, flux is forced to flow in the ordinary directions given by the iron permeability and lamination" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000952_s1064230706030014-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000952_s1064230706030014-Figure1-1.png", "caption": "Fig. 1. An inverted double pendulum.", "texts": [ " The pendulum motion is controlled using only one torque with a limited absolute value applied at the suspension point or at the interlink joint. The domains of controllability of the system are found in both cases. The laws of change of both torques for which the pendulum is stabilized at the top unstable equilibrium position are found in a feedback form. For the constructed control, the bazin of attraction of the required equilibrium position is close (in the linear approximation) to the maximum possible. DOI: 10.1134/S1064230706030014 338 JOURNAL OF COMPUTER AND SYSTEMS SCIENCES INTERNATIONAL Vol. 45 No. 3 2006 FORMAL\u2019SKII Figure 1 shows the planar physical double pendulum with the motionless suspension point O . The suspension is performed using the ideal (without friction) cylindrical joint. The same joint at the point D joins the links of the pendulum, which are absolutely rigid. The axes of the joints at the points O and D are perpendicular to the plane of the drawing. The center of mass of the first link is on the segment OD . Denote by \u03d5 1 and \u03d5 2 the counterclockwise angles of deviation from the vertical line for the first (segment OD ) and second links, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003271_s00170-011-3475-3-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003271_s00170-011-3475-3-Figure6-1.png", "caption": "Fig. 6 Tool paths and points on the tooth profile curve", "texts": [ " am and hm were also described as the cutting parameters of the gear. These values were inputted as constant values to the programme while the gear was cut, where hm and am were described as the cutting depth along the Z-axis and the cutting angle of the tooth profile curves (or increment value in angle an) on the XY-plane, respectively (see Figs. 2, 3 and 5). In the programme, an, bk and bn were described as variables, and initial values were taken as zero. In the previous study, the tool paths were designed as shown in the Fig. 6, where A-B and B\u2019-A\u2019, B-C and C\u2019-B\u2019, C-D and D\u2019-C\u2019, and E-F intervals were described as the dedendum circle, the end mill radius, the radial line segment and an arc of the outside circle, respectively. In a related paper, D-E and F-D\u2019 intervals are the right and left involute curves of the spur gear, respectively. These two intervals were constituted from many points. Points on these two intervals were linearly cut by the end mill. The programme was run according to values such as am=0.25\u00b0, x=0, N=10, f=20\u00b0, m=6 and hm=0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001583_1077546307078829-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001583_1077546307078829-Figure3-1.png", "caption": "Figure 3. Gear kinematic model. (a) Key roll angles (b) gear involute coordinate and parameters.", "texts": [ "comDownloaded from For convenience, the above velocity distributions are expressed as follows where the superscript * denotes the complex conjugation: y x t r 1 r x S0 Qr dr r L [ r exp r t r exp r t ] 2 j (17a) r r j dr r r j dr (17b, c) r x A exp r x B exp r x C exp j r x D exp j r x (17d) A 1 r 2 B 1 r 2 (17e, f) C 1 j r 2 D 1 j r 2 (17g, h) By applying the velocities of the beam to (5a), the sound pressure is finally derived as: p x t b 2 d n r 1 S0 Qr dr r L 2 r exp[ r t d c ] 2 j Gr1 2 r exp[ r t d c ] 2 j Gr2 (18a) Gr1 A exp r r sin c L 1 r r sin c B exp r r sin c L 1 r r sin c C exp j r r sin c L 1 j r r sin c B exp j r r sin c L 1 j r r sin c (18b) Gr2 A exp r r sin c L 1 r r sin c B exp r r sin c L 1 r r sin c C exp j r r sin c L 1 j r r sin c B exp j r r sin c L 1 j r r sin c (18c) at Bibliothekssystem der Universitaet Giessen on May 31, 2015jvc.sagepub.comDownloaded from A meshing spur gear pair is shown in Figure 3 along with key parameters on the gear base circles Rb where represents the gear roll angle and subscripts p and g denote pinion and gear respectively. Further, subscripts m, 0 and L indicate mesh point, mesh start points of pinion and gear respectively and r is a radial distance to mesh point from the gear centers. The line of action (LOA) and off-line of action (OLOA) are also denoted here by x and y. The length of gear involute curve with a circle of radius R is known as follows where s is the involute curve coordinate as shown in Figure 3(b): s ds dx d 2 dy d 2d 1 2 R 2 (19a) x R cos R sin y R sin R cos (19b, c) As the gears roll, the mesh location of pinion and gear at the involute coordinates are: sp 1 2 Rpb[ p0 pm]2 sg 1 2 Rgb[ gL gm]2 (20a, b) at Bibliothekssystem der Universitaet Giessen on May 31, 2015jvc.sagepub.comDownloaded from , 1st teeth pair , 2nd teeth pair. The radii of involute curvature at the mesh are: pm Rpb[ p0 pm] gm Rgb[ gL gm] (21a, b) The sliding velocity at mesh is expressed as follows where pmx and gmx are the velocities of gear and pinion in OLOA direction respectively: S pmx pmx (22a) pmx rpm sin pm p Rpb[ p0 pm] p (22b) gmx rgm sin gm g Rgb[ gL gm] g (22c) Further, the moment arm of the first and second teeth for friction torques are as follows where the subscript c represents mesh cycle: xp1m Rpb[ p0 pc pm] xg2m Rgb[ gL gc gm] (23a, b) xp2m Rpb[ p0 pm] xg2m Rgb[ gL gm] (23c ,d) Given constant operating speeds, the mesh locations on involute coordinates and sliding velocities of a typical gear pair are shown for one mesh cycle in Figure 4" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003288_amr.308-310.1513-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003288_amr.308-310.1513-Figure3-1.png", "caption": "Fig. 3. Elevator car featuring brake block.", "texts": [ " Instantaneous safety gears, the simplest type, where the force increases as function of the distance travelled after its application. The rated speed of elevators using this type of safety is limited to 0.63 m/s. 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: 158.42.28.33, Universidad Politecnica de Valencia, Valencia, Spain-12/05/15,14:10:39) In the studied system the brake block is fixed below the lower members of the moving car frame as seen in figure 3. The cylinder (roller) is connected to the process bar with a lever arm. When the car\u2019s velocity exceeds 15% of the nominal speed an emergency brake happens and the roller remains motionless which causes the roller to move relatively opposite direction to the brake block. The roller (cylinder) moves into the narrowing region between the brake block and guide rail as seen in figure 4. At that moment, by the help of this mechanical clamping equipment, braking can be assured exploiting high friction forces" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000988_pedes.2006.344270-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000988_pedes.2006.344270-Figure1-1.png", "caption": "Fig. 1. Power generation control methods: (a) pitch control, (b) active stall control.", "texts": [ " Using this phenomenon, enhancement of the rotor speed stability of a constant speed WTG with active stall is explained in the next section. WTG The active stall control based constant speed WTG (referred as Type-A2 wind turbine technology) has control of pitch in the negative direction (i.e. between \u221290o to 0o) with respect to pitch control based WTG (Type-A1). The rate of negative pitch control is normally less than 5o per second. Although the pitch rate may exceed 10o per second during emergencies [5]. Fig. 1 shows the difference in the direction of blade rotation between the pitch controlled and the active stall controlled constant speed WTGs. In the figure, the chord line is the straight line connecting the leading and trailing edges of an airfoil. The plane of rotation is the plane in which the blade tips lie as they rotate. The pitch angle (\u03b2) is the angle between the chord line of the blade and the plane of rotation. And, the angle of attack (\u03b1), is the angle between the chord line of the blade and the relative wind or the effective direction of air flow" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002821_j.sna.2010.07.002-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002821_j.sna.2010.07.002-Figure6-1.png", "caption": "Fig. 6. Normal forces acting on a mass element of the frictional layer in down portion of roller.", "texts": [ " 5 normal ng for lateral preload is not illustrated. deformation of frictional layer in down portion of roller relative to stator is obtained as follows: Uk z (x) = Uz(x) + a + \u221a r2 \u2212 x2 \u2212 H, \u2212 L \u2264 x \u2264 L Uk z = A cos ( 2 x ) + a + \u221a r2 \u2212 x2 \u2212 [ A cos ( 2 L ) + r cos + a ] , \u2212 L \u2264 x \u2264 L L = r sin (6) where \u201cH\u201d is vertical distance of the center of roller to x-axis (natural axes of undeformed stator). Velocity and acceleration of the frictional layer are obtained as follows: U\u0307k z = A 2 vw sin ( 2 x ) (7) F R U According to Fig. 6 Newton\u2019s dynamic law can be applied for ig. 13. The theoretical and experimental result of mechanical characteristic of IUSM. ( )2 ( ) \u00a8 k z = \u2212A 2 vw 2 cos 2 x (8) Fig. 15. Photographs of the normal element of frictional layer in down portion of roller. mdxU\u0308k z (x) = p(x)bdx \u2212 cNUk z (x)dx \u2212 dNU\u0307k z (x)dx (9) fabricated prototype. nd Ac p ( F w r fi v I v f i a o m m t U U U fi v w z v U c U U i o o U Y. Hojjat, M.R. Karafi / Sensors a (x) = A b [ (cN \u2212 m\u03c92) cos ( 2 x ) + dN \u03c9 sin ( 2 x ) + cN A ( \u221a r2 \u2212 x2 \u2212 H) ] (10) The unknown boundary (L) can be found iteratively by using Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001679_gt2008-50305-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001679_gt2008-50305-Figure1-1.png", "caption": "Figure 1. Cross section of 65 kW microturbine", "texts": [ " Kxy, cross coupled stiffness \u03c4, torque \u03b2, efficiency factor from Alford\u2019s equation Dp, blade diameter of stage H, blade height ATP, acceptance test procedure DN, diameter (mm) * rotor speed (rpm) CMM, Coordinate Measurement Machine MTBF, Mean Time Between Failures APU, Auxiliary Power Unit FFT, Fast Fourier Transform RPM, Revolutions per Minute ISO, International Organization for Standardization The gas turbine of interest to this case study is a commercially available Capstone C65 MicroTurbine\u2122. The system produces 65 kW of electrical energy at ISO conditions. Uniquely, this gas turbine is completely oil-free, supported only on foil bearings. The 65 kW microturbine was a completely new model with more aggressive power densities than previous models. Pressure ratios, tip speeds, and temperatures were increased to values seen in aerospace gas turbines, such as APU\u2019s. A cross section of the engine, as installed within the recuperator, is seen in Figure 1. Figure 2 is a schematic wnloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/27/2016 T representation of the rotor\u2013bearing system from the rotordynamic model. The design consists of a two-part rotor installed and rotating as a single assembly. The first part is the powerhead, consisting of a radial compressor and radial turbine supported on a common bearing shaft. The second part of the rotor is a permanent magnet generator-motor shaft supported by two radial foil bearings. The two parts are connected through a patented flexible coupling that eliminates internal friction, as part of the careful control of instability drivers" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001838_0094-114x(75)90072-5-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001838_0094-114x(75)90072-5-Figure4-1.png", "caption": "Figure 4. Graves' theorem.", "texts": [ " Hence the light ray h, tangent to the ellipse g, is reflected in P on the confocal ellipse p into another tangent t2 of g; obviously this is the root of Poncelet's theorem. There is another famous theorem, due to Ch. Graves (1841) and closely connected with the previous facts: If a closed rope, longer than the perimeter of an ellipse g, is slung around g and spanned by a pin P, then P can be led along an ellipse p, confocal with g. In fact the system is in equilibrium, if the spanning force T applied at P is acting in the direction of the internal bisector of the straight parts of the rope, as the tension forces T,, T2 in these parts have the same value (Fig. 4). Hence the pin P can move only orthogonally to this bisector if the direction of T is changed, which means that the external bisector t of T~ and 1\"2 is the tangent of the path p of P.+ According to Poncelet's theorem, p is an ellipse. The classical \"gardener's construction\" of the ellipse appears as a limit case of Graves' theorem, i.e. when the elliptical base g is reduced to the focal segment EF. Without pretending that these heuristic considerations are a rigorous proof of Graves' theorem--an exact proof will be indicated in section 3--we can complete Fig. 4 by adding its mirror image with respect to the tangent t and thus obtain a mechanism equivalent to the pair of elliptic wheels pj, p2: it consists of two equal elliptic disks gl and g2, confocal with p, and pz, respectively, and connected by a closed crossed belt which transmits the rotation about E, into a rotation about F~ (Fig. 5). Of course the transmission ratio of the angular velocities is not constant. The reasoning was essentially based upon the fact that the rolling curves p, and p: are always symmetrically situated with respect to their common tangent t" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003118_iros.2009.5354676-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003118_iros.2009.5354676-Figure2-1.png", "caption": "Fig. 2. Forces and torques acting on each link of the snake robot.", "texts": [ "00 \u00a92009 IEEE 3615 This section gives a brief introduction to a mathematical model of a planar snake robot previously presented in [10]. A feedback linearized form of this snake robot model is developed in Section III in order to simplify the controllability analysis presented in Section V. A. Notations and defined identities The snake robot consists of n links of length 2l interconnected by n\u22121 joints. The mathematical identities defined in order to describe the kinematics and dynamics of the snake robot are described in Table I and illustrated in Fig. 1 and Fig. 2. All n links have the same length, mass, and moment of inertia. The total mass of the snake robot is therefore nm. The mass of each link is uniformly distributed so that the link CM (center of mass) is located at its center point (at length l from the joint at each side). The following vectors and matrices are used in the subsequent sections: A := \u23a1\u23a2\u23a31 1 . . . . 1 1 \u23a4\u23a5\u23a6,D := \u23a1\u23a2\u23a31 \u22121 . . . . 1 \u22121 \u23a4\u23a5\u23a6 where A \u2208 R(n\u22121)\u00d7n and D \u2208 R(n\u22121)\u00d7n. Furthermore, e := \u00a3 1 . . 1 \u00a4T \u2208 Rn E = \u2219 e 0n\u00d71 0n\u00d71 e \u00b8 \u2208 R2n\u00d72 sin \u03b8 := \u00a3 sin \u03b81 ", " Performing the matrix multiplication and assembling the friction forces on all links in matrix form gives fR=\u2212 \u2219 ct (C\u03b8) 2 + cn (S\u03b8) 2 (ct \u2212 cn)S\u03b8C\u03b8 (ct \u2212 cn)S\u03b8C\u03b8 ct (S\u03b8) 2 + cn (C\u03b8) 2 \u2219\u0327 x\u0307 y\u0307 \u00b8 (10) where fR= \u00a3 fTR,x fTR,y \u00a4T \u2208 R2n. Note that in the case of uniform friction (ct = cn = c), the expression for the friction forces is reduced to (8). This section presents the equations of motion of the snake robot in terms of the acceleration of the link angles, \u03b8\u0308, and the acceleration of the CM of the snake robot, p\u0308. These coordinates describe all n+ 2 DOFs of the snake robot. The forces and torques acting on link i are visualized in Fig. 2. The force balance for link i in global frame coordinates is given by mx\u0308i = fR,x,i + hx,i \u2212 hx,i\u22121 my\u0308i = fR,y,i + hy,i \u2212 hy,i\u22121 (11) while the torque balance for link i is given by J\u03b8\u0308i = ui \u2212 ui\u22121 \u2212l sin \u03b8i(hx,i + hx,i\u22121) + l cos \u03b8i(hy,i + hy,i\u22121) (12) Through straightforward calculations, it is shown in [10] that (11) and (12) may be rewritten for all links and combined into the following complete model of the snake robot: M\u03b8\u0308 +W\u03b8\u0307 2 \u2212 lS\u03b8NfR,x + lC\u03b8NfR,y = DTu nmp\u0308 = ET fR (13) where \u03b8 and p represent the n + 2 generalized coordinates of the system, \u03b8\u0307 2 = diag(\u03b8\u0307)\u03b8\u0307, and M := JIn\u00d7n +ml2 (S\u03b8V S\u03b8 + C\u03b8V C\u03b8) W := ml2 (S\u03b8V C\u03b8 \u2212 C\u03b8V S\u03b8) N := AT \u00a1 DDT \u00a2\u22121 D V := AT \u00a1 DDT \u00a2\u22121 A (14) This section transformes the model from [10], which was summarized in the previous section, to a simpler form through partial feedback linearization" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002592_1.2772634-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002592_1.2772634-Figure3-1.png", "caption": "Fig. 3 FE model of a blisc with a CR damper", "texts": [ " 30 and 31 is negligible compared with the computational expense necessary for forced response calculation, and in practical calculations the sensitivity analysis does not cause a noticeable increase of calculation time. As an example, forced response of an ADTurbII blisc see Ref. 18 with cottage-roof CR underplatform dampers UPDs is analyzed. The blisc consists of 24 blades, and a finite element model of its one sector contains 21,555 DOFs. The UPD is modeled by a damper model developed in Ref. 19 . Finite element FE blisc and UPD models are shown in Fig. 3. In the numerical studies presented here a cyclically symmetric bladed disk is considered and it is assumed that the stochastic variation of UPD parameter keeps the structure cyclically symmetric. As a result of this assumption, specific effects caused by blade and UPD mis- tuning e.g., see Ref. 20 are not considered here. MARCH 2008, Vol. 130 / 022503-5 6 Terms of Use: http://www.asme.org/about-asme/terms-of-use m r o s d = 3 s d f o b d i o u i F t v F p d r s O 0 Downloaded Fr The UPD model is formulated for a general case of multiharonic forced response analysis" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002997_bf02476900-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002997_bf02476900-Figure2-1.png", "caption": "Figure 2. Coordinate systems. The horizontal canal lles in the (x', y ')-plane; the anterior vertical canal, in the (y~, z~)-plane; and the posterior vertical canal, in the (x ~, y~)-plane. Only the projection of the AVC onto the (y', z ')-plane has been depicted. A similar projection should be visualized for AVC onto the (x', y')-plane, and for PVC onto the (x', y ')-plane and the (y', z')-plane. A similar spatial figure has to be imagined for the other labyr in th on the", "texts": [ " The second reference system, (x, y, z) rotates with the turntable about y, and its origin 0 coincides with the origin 0o, and 6---B.M.B, labyrinth in the head, we introduce a third system of coordinates (x', y', z') fixed in the head so that its origin 0' = 0 = 0o, the y'-axis lies in the sagittal plane, and the horizontal semicircular canals lie in the (x', z')-plane. In addition, z' = z. Any point of the labyrinth, and of the endolymph, is specified as P ( x ' , y ' , z') and biometrically well defined. The coordinates :~', y ' , z' of any P are, of course, constant in time (Fig. 2). The tilt which the head, and therefore any point P ( x ' , y ' , z ') undergoes results from a rotation of (x' , y ' ,z') around z'. That is, P goes through an angular displacement relative to (x, y, z), and a tangential linear velocity, v-~, corresponds to the tilt which depends on the rotation about z'. Summarizing, a point P ( x ' , y ' , z ') in the endolymph rotates about y due to a certain angular velocity w~, and moves relative to (x, y, z) with a tangential velocity va by virtue of a certain angular velocity wz, =- wz. According to the foregoing conditions, the direction angles between the system (x, y, z) and the system (x', y', z') are (Fig. 2) Tf 7r al = w~t; a2 = -~ - w # ; a 8 = -~; = + 2t; = 2t; = 2 ; (11) ~'r Tr ?/1 ~ ~ ~ 2 ' ?/2 ---~ ~ ; ?/3 = 0 . Henceforth we can transform the coordinates in (x', y', z') of P into coordinates in (x, y, z) by means of the matrix COS tozt T z = sin wzt 0 - s i n ~oz~ 0 ] cos eOzf 0 0 1 Consequently the coordinates of P in (x, y, z) = T~ y' Z will be functions of time. (12) (13) MECHANICAL FORCES IN THE SEMICIRCULAR CAI~AL 273 prolongation of x ' through 0 274 M, VALENTINUZZI IV . Geometrical description of the canals in the head coordinates", " Since the vertical canals form an angle ~ with the (x', y')-plane, it will be convenient to use their projections onto the (x', y')-plane and the (y', z')-plane in the following calculations. Take the plane in which the canal is actually located, and use an auxiliary set of coordinate axes (x\", y\", z\") so tha t Y'IlY', x\" and z' lie in the (x', z')-plane, the anterior vertical semicircular canal lies in the (y\", z\")-plane, and the posterior vertical canal lies in the (x\", y\")-plane. Therefore, x\", z\" are the intersections of the (x', z')-plane with the (x\", y\")-plane and the (y\", z\")-plane respectively. Each x\"-axis and z\"-axis forms an angle 0 with x' (Fig. 2). In the (x ~, y\", z\") system the anterior canal circumference is described by: x\" = O; ] y\" = y~ + Rs in O; z\" = z'~ + R cos 0, (15) where 0 is the angle formed by the radius R of the canal and the z\"-axis. For the posterior canal circumference we have x c + R cos 0; i y \" = y ~ + R s i n 0 ; z\" = O, (16) where 0 is the angle formed by the canal radius R and the x\"-axis. MECHANICAL FORCES IN THE SEMICIRCULAR CANAL 275 In the (x', y')-plane the expressions (15) yield Xo + zc cos ~) + R cos 0 cos 0; f y' y~ + R sin 0, (17) with x~ for the origin of the auxil iary (x\", y~, zU)-system in (x', y', z')", " (80) M E C H A N I C A L F O R C E S I N T H E SEMICIRCULAR CANAL 285 With the conditions of this ease, the Coriolis acceleration formula (25) is \"-)\" COyco\u2022 a -- ~ ~b've sin ~) cos wzt. (81) The mean force and mean pressure are fco = - 2~P~coyco~Svc sin 0 cos cozt; (82) Pco = - 2r sin 0 cos co~. (83) The fact that the Coriolis acceleration (and force and pressure) gives for the vertical canal an integral effect along the endolymph depending on cos co~t with, consequently, a maximum for co~t = 0 and zero for cozt = ~r/2, can be predicted and understood if we take two opposite points in the (y', z')-projection of the canal (say P1 and P2, Fig. 2) and calculate the resultant torque from the Coriolis force at those points. This torque is a function of cos co~t. By adding the results in each of the cases just analyzed (Table II), we obtain, according to (19), the resulting mean acceleration, mean force and mean pressure (Table III). VI I . Discussion. The present analysis shows that the area of the semicircular canal (SHe, Svc) is an important factor on which the mean acceleration and, consequently, the mean force and the mean endolymphatic pressure depend" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000043_s0167-8922(08)70194-9-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000043_s0167-8922(08)70194-9-Figure2-1.png", "caption": "Figure 2 Slipline field for normal indenter plus radial traction force.", "texts": [], "surrounding_texts": [ "The i n f l u e n c e of l u b r i c a n t c o n t a m i n a t i o n and subsequent s u r f a c e damage on r o l l i n g b e a r i n g f a t i g u e h a s formed t h e b a s i s of s e v e r a l s t u d i e s over r e c e n t y e a r s . Webster e t a 1 (1) c a l c u l a t e d t h e e l a s t i c s u b s u r f a c e stress f i e l d s from r e a l d e n t p r o f i l e s and used t h e s e a s i n p u t t o a f a t i g u e l i f e model t o de te rmine t h e r e d u c t i o n i n l i v e s . I n t h i s paper a s l i p l i n e f i e l d a n a l y s i s h a s been used t o c a l c u l a t e t h e s u b s u r f a c e r e s i d u a l stresses f o r d i f f e r e n t i d e a l i s e d d e n t / r o l l e r combina t ions . These stress f i e l d s were superimposed upon t h o s e c a l c u l a t e d from a d r y c o n t a c t a n a l y s i s of t h e o v e r r o l l i n g of t h e dent which used novel m u l t i - l e v e l t e c h n i q u e s t o a c c e l e r a t e convergence, and t h e r e s u l t a n t stress f i e l d s provided i n p u t t o t h e f a t i g u e l i f e model. The i n f l u e n c e of t h e d e n t i t s e l f on l i f e a p p e a r s t o be s m a l l and o n l y becomes s i g n i f i c a n t on i n c l u s i o n o f t h e r e s i d u a l stresses. These have a p a r t i c u l a r l y marked e f f e c t a s t h e r o l l e r r a d i u s and l o a d a r e reduced, s u g g e s t i n g e x p e c t e d l i v e s may not i n c r e a s e a s r a p i d l y with d e c r e a s i n g l o a d a s would be expec ted from convent iona l models. 1 INTRODUCTION Lubricant contaminat ion and subsequent s u r f a c e damage i s i n c r e a s i n g l y recognised a s having a s i g n i f i c a n t e f f e c t on b e a r i n g f a t i g u e l i v e s which would n o t normally be p r e d i c t e d by t h e t r a d i t i o n a l Lundberg and Palmgren model (2). I n ( 3 ) I o a n n i d e s and H a r r i s p r e s e n t e d a n i m p o r t a n t g e n e r a l i s a t i o n o f t h e c l a s s i c a l model i n which t h e s t ress p e r t u r b a t i o n s r e s u l t i n g from non-per fec t ly smooth c o n t a c t i n g s u r f a c e s can be inc luded . E s s e n t i a l l y t h e new model is a n e l e m e n t a l form o f t h e e a r l i e r a n a l y s i s where t h e p r o b a b i l i t y of f a i l u r e i s c a l c u l a t e d from t h e cumulat ive c o n t r i b u t i o n of stresses above a t h r e s h o l d v a l u e i n s m a l l volume e lements of t h e m a t e r i a l . Therefore t o a p p l y t h e model a comple te h i s t o r y of t h e stress f i e l d underneath t h e d e n t a s it p a s s e s through t h e c o n t a c t i s r e q u i r e d . T h i s f i e l d w i l l be e f f e c t i v e l y made up of two components t h e r e s i d u a l stress f i e l d from t h e i n d e n t a t i o n and t h e EHD stress f i e l d from t h e subsequent o v e r r o l l i n g of t h e d e n t . U n f o r t u n a t e l y t h e s e a r e n o t n e c e s s a r i l y independent as f u r t h e r p l a s t i c deformation of t h e d e n t edges and some shakedown w i l l p r o b a b l y occur , b u t a u s e f u l f i r s t approximat ion can b e made by t r e a t i n g them a s such . T h i s problem was f i r s t t a c k l e d by Webster, I o a n n i d e s and S a y l e s (l), where t h e y measured t h e p r o f i l e s o f d e n t s f rom a r e a l b e a r i n g s u r f a c e . This in format ion was t h e i n p u t t o a numerical c o n t a c t model coupled t o a f i n i t e e lement a n a l y s i s t o d e t e r m i n e t h e s t a t e o f stress i n a n i n d e n t e d b e a r i n g raceway. The stress i n f o r m a t i o n was used i n conjunct ion wi th t h e modif ied f a t i g u e model t o determine t h e r e d u c t i o n i n f a t i g u e l i f e . I n ( 4 ) a p p r o x i m a t e s o l u t i o n s w e r e p r e s e n t e d f o r t h e i n t e r f a c i a l p r e s s u r e s and d e f l e c t i o n s r e s u l t i n g from debris p a r t i c l e s b e i n g squashed i n t h e i n l e t t o an EHD c o n t a c t . A p l a n e s t r a i n approach was used t o a l l o w t h e a p p l i c a t i o n of 2 d imens iona l e x t r u s i o n t h e o r y . The problem i s s i m i l a r t o s t r i p r o l l i n g , b u t i n t h i s c a s e t h e r o l l e r s cannot b e e x p r e s s e d a s c i r c u l a r a r c s a s t h e e l a s t i c d e f l e c t i o n s due t o t h e EHD and e x t r u s i o n p r e s s u r e s a r e s i g n i f i c a n t . The s o l u t i o n was f o u n d by numer ica l ly i t e r a t i n g between t h e p r e s s u r e and d e f l e c t i o n e q u a t i o n s u n t i l convergence. As t h e r o l l i n g s u r f a c e s c o n t i n u e t o a p p r o a c h one a n o t h e r i n t h e i n l e t , t h e i n t e r f a c i a l p r e s s u r e s may i n c r e a s e s u f f i c i e n t l y such t h a t p l a s t i c deformat ion of t h e raceways can o c c u r . The c r i t i c a l v a l u e of p r e s s u r e may b e e s t i m a t e d from i n d e n t a t i o n exper iments or t h e o r e t i c a l c o n s i d e r a t i o n s such a s ( 4 ) . By a p p l y i n g t h e s e c r i t e r i a t o t h e model , t h e a p p r o x i m a t e d e n t s h a p e s a n d p r e s s u r e d i s t r i b u t i o n s may b e f o u n d a n d compared w i t h t h o s e measured e x p e r i m e n t a l l y ( 5 ) . The measured p r o f i l e s o f t h e s e d e n t s c l o s e l y approximate a c i r c u l a r a r c s i m i l a r t o t h e i m p r e s s i o n l e f t by a r i g id c y l i n d r i c a l i n d e n t e r . Dumas and Baronet ( 6 ) produced a f i n i t e e l e m e n t s o l u t i o n t o t h e c i r c u l a r i n d e n t e r problem and found t h a t a t s i g n i f i c a n t d e p t h s of d e p r e s s i o n , where t h e s u b s u r f a c e zone w a s e n t i r e l y p l a s t i c t h e c a l c u l a t e d p r e s s u r e p r o f i l e w a s f a i r l y c o n s t a n t o v e r most o f t h e i n t e r f a c e . A t s m a l l e r d e p t h s o f d e p r e s s i o n t h e p r e s s u r e p r o f i l e i s n o t f l a t , b u t r e a c h e s a maximum a t t h e c e n t r e where t h e p r e s s u r e i s approximate ly 50% g r e a t e r t h a n it i s n e a r t h e p e r i p h e r y o f t h e d e n t . T h i s s u g g e s t s t h a t f o r r e l a t i v e l y s m a l l i n d e n t a t i o n s t h e i n t e r f a c i a l p r e s s u r e i s s t i l l i n c r e a s i n g t o w a r d s t h e c e n t r e of t h e d e n t , t h u s g e n e r a t i n g a r a d i a l e x t r u s i o n f o r c e on t h e d e b r i s p a r t i c l e r e s u l t i n g i n a s u r f a c e t r a c t i v e f o r c e on t h e b e a r i n g raceway. A t g r e a t e r d e p t h s o f i n d e n t a t i o n t h e p r e s s u r e p r o f i l e f l a t t e n s and t h e r e s u l t a n t s u r f a c e t r a c t i o n f o r c e d i m i n i s h e s . O u t s i d e t h e i n d e n t e d zone where t h e d e f l e c t i o n s a r e e n t i r e l y e l a s t i c , d e b r i s d e f o r m a t i o n w i l l c o n t i n u e t o o c c u r and t h e p r e s s u r e p r o f i l e w i l l b e d e f i n e d by t h e e x t r u s i o n a n d e l a s t i c i t y e q u a t i o n s a s d i s c u s s e d i n ( 4 ) . The r e s i d u a l stresses r e s u l t i n g from t h e i n d e n t a t i o n e s p e c i a l l y i n t h e p r e s e n c e of a s u r f a c e t r a c t i o n f o r c e may encourage f a t i g u e c rack i n i t i a t i o n a s p o s t u l a t e d by Olver ( 7 ) . I n t h i s paper a s l i p l i n e f i e l d model based on t h a t deve loped by Olver was used t o e s t i m a t e t h e p l a s t i c stresses i n t h e s u b s u r f a c e and t h e r e s u l t a n t r e s i d u a l stresses on unloading. A f t e r i n d e n t a t i o n t h e s q u a s h e d d e b r i s p a r t i c l e i s l o s t and t h e damaged s u r f a c e r e p e a t e d l y p a s s e s t h r o u g h t h e EHD c o n t a c t . Model l ing t h i s p r o c e s s i s n o t p r a c t i c a b l e a t p r e s e n t , however a s t h e f i l m t h i c k n e s s i s s m a l l compared t o t h e e l a s t i c deformat ion of t h e s u r f a c e and w i l l p r o b a b l y b e f u r t h e r reduced around t h e edges of t h e d e n t , a d r y c o n t a c t a p p r o x i m a t i o n may n o t b e t o o u n r e a l i s t i c . C l e a r l y , a l t h o u g h t h e r e s i d u a l stress d i s t r i b u t i o n i s f i x e d r e l a t i v e t o t h e d e n t t h e super imposed c o n t a c t stress f i e l d w i l l b e changing d u r i n g o v e r r o l l i n g and t h e f u l l h i s t o r y w i l l be r e q u i r e d f o r t h e l i f e c a l c u l a t i o n s . 1.1 Notation A c o n s t a n t b ha l f -wid th of Her tz ian c o n t a c t , m dB incrementa l a r e a of t h e c r o s s - s e c t i o n of t h e r i n g , mm2 e l i f e exponent H d imens ionless f i l m t h i c k n e s s H O O c o n s t a n t i n d imens ionless f i l m k y i e l d s h e a r stress, N/m2 N f a t i g u e l i f e P dimens ionless p r e s s u r e R reduced r a d i u s of c u r v a t u r e , m S p r o b a b i l i t y of s u r v i v a l x , x ' c o o r d i n a t e s i n r o l l i n g d i r e c t i o n , m X , X ' d imens ionless c o o r d i n a t e s y c o o r d i n a t e i n a x i a l d i r e c t i o n , m z c o o r d i n a t e normal t o t h e s u r f a c e , m z ' stress weighted average depth , mm R domain a' extended domain t h i c k n e s s equat ion 2 SLIPLINE FIELD I n t h e s l i p l i n e f i e l d a n a l y s i s t h e d e b r i s i n d e n t a t i o n of t h e b e a r i n g s u r f a c e i s model led by a r i g i d d ie i m p r e s s i n g a r i g i d - p l a s t i c h a l f s p a c e c o u p l e d w i t h a t r a c t i o n f o r c e e x t e n d i n g from t h e c e n t r e l i n e . I f t h e t r a c t i o n f o r c e i s z e r o t h i s r e d u c e s t o t h e c l a s s i c a l f i e l d f o r a r i g i d i n d e n t e r a s p r e s e n t e d by H i l l ( 9 ) i n t h e deforming zone. I n ( 7 ) Olver p r e s e n t e d a s o l u t i o n f o r a s i n g l e u n i - d i r e c t i o n a l t r a c t i o n f o r c e a p p l i e d t o t h e i n t e r f a c e , f i g u r e 1. A s t h e s l i p l i n e s w i l l n o t m e e t t h e i n t e r f a c e a t 45' a n e x t r a f a n must be drawn i n ( r e g i o n BCF). I n t h e r e g i o n DCFE t h e l i n e s w e r e c a l c u l a t e d n u m e r i c a l l y by a f i n i t e d i f f e r e n c e t e c h n i q u e proposed by H i l l ( 9 ) and t h e e l a s t i c b o u n d a r y ( B E ) was d e r i v e d a n a l y t i c a l l y by Johnson ( 8 ) . I f t h e t r a c t i o n f o r c e i s made e q u a l t o z e r o t h e s l i p l i n e s m e e t t h e s u r f a c e a t 45O and t h e f i e l d reduces t o t h a t proposed f o r a p u r e l y normal i n d e n t a t i o n . If t h e t r a c t i o n f o r c e e x t e n d s r a d i a l l y f rom t h e c e n t r e t h e n t h e zone below t h e c e n t r e l i n e w i l l be p l a s t i c and t h e s l i p l i n e s m u s t m e e t a t 90\u00b0, f i g u r e 2 . From symmetry a s q u a r e s e c t i o n BEFG can be drawn and as b e f o r e a f a n i s c o n s t r u c t e d between BC and BE. The s l i p l i n e s i n r e g i o n s CDHE and EHIF c a n b e c a l c u l a t e d n u m e r i c a l l y by f i n i t e d i f f e r e n c e s . Although t h e s l i p l i n e f i e l d could be ex tended i n d e f i n i t e l y t h e a c t u a l zone o f p l a s t i c i t y w i l l o n l y e x i s t d i r e c t l y below t h e i n d e n t e r . D e f i n i n g t h e e l a s t i c / p l a s t i c boundary e x a c t l y i s n o t p o s s i b l e , b u t an approximat ion c a n b e made by a p p l y i n g t h e i n t e r f a c e p r e s s u r e s t o an e l a s t i c h a l f s p a c e and c a l c u l a t i n g where t h e y i e l d s h e a r stress is exceeded, a s s u g g e s t e d by Olver ( 7 ) . S i g n i f i c a n t p l a s t i c deformat ion w i l l p r o b a b l y o n l y o c c u r i n t h e r e g i o n OADCB, whereas i n t h e remainder o f t h e f i e l d t h e m a t e r i a l may be p l a s t i c b u t w i l l no t deform. Having c o n s t r u c t e d t h e f i e l d t h e stresses may b e c a l c u l a t e d i n t h e normal manner see ( 7 ) . During unloading , a t e n s i l e normal f o r c e e q u a l i n magnitude t o t h e p l a s t i c p r e s s u r e s i s e f f e c t i v e l y a p p l i e d t o t h e i n t e r f a c e . C o n s e q u e n t l y t h e r e s i d u a l stresses c a n b e found by summing t h e stress d i s t r i b u t i o n r e s u l t i n g from t h i s e l a s t i c t e n s i l e f o r c e t o t h e p l a s t i c stress d i s t r i b u t i o n . The e l a s t i c s tresses w e r e c a l c u l a t e d a n a l y t i c a l l y a s d e s c r i b e d by Ford ( 1 0 ) . A map of t h e p r i n c i p a l r e s i d u a l s tress d i s t r i b u t i o n s f o r t r a c t i o n c o e f f i c i e n t s of 0 and 0 . 2 are shown i n f i g u r e s 3 ( a ) a n d ( b ) a n d a c o n t o u r map o f t h e o r t h o g o n a l r e s i d u a l s h e a r stresses i n f i g u r e ( 4 ) . The e f f e c t of t h e s u r f a c e s h e a r stress is t o r e d u c e t h e t e n s i l e r e s i d u a l s tresses immediately below t h e s u r f a c e b u t t o i n c r e a s e t h o s e below t h e dent s h o u l d e r . 3 DRY CONTACT EQUATIONS To s o l v e t h e d r y c o n t a c t p r o b l e m i t i s n e c e s s a r y t o f i n d a p r e s s u r e d i s t r i b u t i o n P ( X ) , t h a t s a t i s f i e s t h e f i l m t h i c k n e s s e q u a t i o n ( l ) , i e s o l v e t h e i n t e g r a l e q u a t i o n ( 2 ) f o r t h e p r e s s u r e P . H ( X ) = 0 XE a (1) i n t h e one d i m e n s i o n a l c a s e , t h i s e q u a t i o n r e a d s : H, + - - - P(X') K(X , X') dX' = 0 XE a XL 2 I t ' I n (2) where K ( X , X ' )=In 1 X-X' I U n f o r t u n a t e l y t h e domain on which e q u a t i o n (1) h o l d s i s g e n e r a l l y n o t known i n a d v a n c e . Therefore (1) is extended t o ( 3 ) such t h a t (1) h o l d s i n t h e o r i g i n a l domain and P ( X ) = O X E ~ ' P ( X ) > 0 H ( X ) = 0 XE a P ( X ) = 0 H ( X ) > 0 X E R ' ( 3 ) 191 A second c o n d i t i o n t h a t has t o be s a t i s f i e d i s t h e f o r c e b a l a n c e e q u a t i o n : b ( X ) d x = f n ( 4 ) The i n t e r a c t i o n of t h e s e e q u a t i o n s i s i d e n t i c a l t o t h e l u b r i c a t e d c o n t a c t case , see f o r i n s t a n c e (11). I n t h i s d r y c o n t a c t a n a l y s i s t h e f i l m t h i c k n e s s i s set t o zero over t h e domain (a ) and t h e n o n l y t h e p r e s s u r e d i s t r i b u t i o n has t o b e c a l c u l a t e d . The problem h a s t r a d i t i o n a l l y b e e n s o l v e d b y d i r ec t m e t h o d s (Newton Raphson) , r e s u l t i n g i n l o n g computing t i m e s f o r l a r g e problems. AS a l a r g e number o f s o l u t i o n s i s r e q u i r e d i n a p p l y i n g t h e l i f e model t h e c o m p u t i n g t i m e c a n become e x c e s s i v e l y long . To a l l e v i a t e t h i s problem a n a l t e r n a t i v e i t e r a t i v e t y p e s o l u t i o n was used which h a s been d e s c r i b e d i n (12) . Convergence w a s f u r t h e r a c c e l e r a t e d by t h e a p p l i c a t i o n of novel m u l t i - l e v e l t e c h n i q u e s which have been a p p l i e d p r e v i o u s l y t o t h e s o l u t i o n o f d i f f e r e n t i a l e q u a t i o n s . The computing t i m e f o r t h e u s u a l i t e r a t i v e s o l u t i o n t e n d s t o be dominated by t h e i n t e g r a l c o m p u t a t i o n , s o t h e s o l u t i o n t i m e i s p r o p o r t i o n a l t o o r d e r n2 where n i s t h e number of p o i n t s , a s can b e seen i n Table 1. However it i s p o s s i b l e t o reduce t h e comput ing , t ime t o o r d e r n logn by t h e a p p l i c a t i o n of m u l t i l e v e l t e c h n i q u e s , s p e c i f i c a l l y M u l t i l e v e l M u l t i - I n t e g r a t i o n ( M L M I ) . The b a s i c i d e a was g i v e n i n (13) and worked out i n d e t a i l i n ( 1 2 ) . A s can be seen from Table 1, column 3 s i g n i f i c a n t t i m e s a v i n g s c a n be o b t a i n e d from l e v e l 6 onwards, s o t h e approach i s most u s e f u l f o r p r o b l e m s w i t h many g r i d p o i n t s . The c a l c u l a t i o n of t h e s u b s u r f a c e stresses i s a t a s k s i m i l a r t o t h e f i l m t h i c k n e s s s o l u t i o n , when one c o n s i d e r s o n l y one p a r t i c u l a r v a l u e of t h e d e p t h z, a t a t i m e . P l o t s of t h e d r y c o n t a c t s u r f a c e p r e s s u r e a n d a s s o c i a t e d s u b s u r f a c e o r t h o g o n a l s h e a r s t r e s s d i s t r i b u t i o n are shown i n f i g u r e 5 . 4 L I F E PREDICTIONS I n t h e t r a d i t i o n a l Lundberg a n d Pa lmgren b e a r i n g f a t i g u e l i f e model, t h e p r o b a b i l i t y of f a i l u r e c a n be e x p r e s s e d i n t e r m s o f t h e stressed volume and t h e magnitude and d e p t h below t h e s u r f a c e of t h e maximum o r t h o g o n a l s h e a r stress. I n ( 2 ) I o a n n i d e s and Har r i s p r o p o s e d a g e n e r a l i s e d model i n which t h e s t r e s s e d volume i s d i v i d e d i n t o d i s c r e t e volume e lements i n which t h e maximum stress i s c a l c u l a t e d a c c o r d i n g t o some stress r e l a t e d f a t i g u e c r i t e r i o n . I n common w i t h s t r u c t u r a l f a t i g u e l i f e p r e d i c t i o n s f o r s teels i n r e v e r s e d b e n d i n g or t o r s i o n , a t h r e s h o l d stress v a l u e i s d e f i n e d below which f a i l u r e w i l l n o t o c c u r . Each e l e m e n t i s w e i g h t e d a c c o r d i n g t o i t s depth below t h e s u r f a c e and t h e p r o b a b i l i t y o f f a i l u r e i s e x p r e s s e d i n terms o f t h e i n t e g r a l o f t h e e l e m e n t a l stresses over t h e e n t i r e volume. A modif ied l i f e c r i t e r i o n h a s been used t o compute t h e Ll0 b e a r i n g l i v e s i n t h e p r e s e n t c a s e . The maximum s h e a r stress ampl i tude z, i s c a l c u l a t e d d u r i n g t h e o v e r r o l l i n g of t h e d e n t . I n common w i t h s t r u c t u r a l f a t i g u e , t h e f a t i g u e stress t h r e s h o l d z, i s m o d i f i e d a c c o r d i n g t o t h e a b s o l u t e v a l u e of t h e s h e a r stress. z, i s assumed t o remain unchanged i f T,, d o e s n o t e x c e e d t h e y i e l d s t ress a n d t o d i m i n i s h l i n e a r l y t o z e r o f o r Tmax v a r y i n g between r e a n d t h e f r a c t u r e s t r e n g t h zf. A s t h e c rack might be e x p e c t e d t o b e c r e a t e d more e a s i l y i n t h e p r e s e n c e of a t e n s i l e r a t h e r t h a n compressive s t ress f i e l d , a n a d d i t i o n a l h y d r o s t a t i c weight ing was i n c l u d e d i n t h e model. I n t h i s t h e c r i t i c a l stress Ta was modi f ied to fa+ a.Hp, where Hp is e q u a l t o t h e h y d r o s t a t i c p r e s s u r e and a i s t a k e n as a=0.3.Using t h e s e v a l u e s , t h e p r o b a b i l i t y of s u r v i v a l of t h e i n n e r r i n g can be e x p r e s s e d a s : The e f f e c t i v e p e r t u r b a t i o n on t h e g l o b a l p r e s s u r e d i s t r i b u t i o n by t h e d e n t w i l l depend v e r y much on t h e r a t i o of t h e d e n t wid th t o t h e Her tz c o n t a c t s i z e . To a s s e s s t h i s e f f e c t f o u r d e n t / r o l l e r combinat ions w e r e chosen . An a r t i f i c i a l c i r c u l a r d e n t of 200 micron wid th and 3 micron d e p t h was o v e r r o l l e d by r o l l e r s of 2 , 4 , 8 and 16mm r a d i u s . The o v e r r o l l i n g of t h e d e n t is s i m u l a t e d u s i n g 9 d i f f e r e n t p o s i t i o n s o f t h e r o l l i n g element wi th r e s p e c t t o t h e d e f e c t , i n o r d e r t o p i c k up t h e maximum stresses. The p o s i t i o n of t h e c e n t r e x c of t h e r o l l i n g element i s g i v e n by: xc = b (n-5) /2 f o r n=1,2 ,..., 9 . The stress h i s t o r y i n e a c h p o i n t i s a n a l y s e d wi th r e s p e c t t o t h e s e n i n e p o s i t i o n s and t h e n t h e l i f e i n t e g r a l i s c a l c u l a t e d . A s t h e number of p o s i t i o n s i n t i m e and space a r e r e l a t i v e l y smal l , 9 and 49x17 r e s p e c t i v e l y , t h e v a l u e s of t h e l i f e i n t e g r a l s a re r a t h e r jumpy a n d c o n s e q u e n t l y t h e numer ica l r e s u l t s s h o u l d be i n t e r p r e t e d w i t h c a r e . 5 RESULTS The l i f e i n t e g r a l s were c a l c u l a t e d f o r f o u r d i f f e r e n t v a l u e s o f t h e r e d u c e d r a d i u s o f c u r v a t u r e a n d f o r s i x d i f f e r e n t l o a d s ( c o r r e s p o n d i n g t o H e r t z i a n p r e s s u r e s r a n g i n g f rom 2 . 0 t o 3 . 3 GPa) . T h r e e cases were examined, a smooth raceway, a raceway wi th one d e n t a n d a r a c e w a y w i t h o n e d e n t a n d a s s o c i a t e d r e s i d u a l stress f i e l d . The e f f e c t of t h e s e stress f i e l d s on p r e d i c t e d l i v e s c a n be g r a p h i c a l l y e x p r e s s e d i n t e r m s of r i s k maps. I n t h e s e a s e c t i o n of t h e x, z p l a n e i s drawn on a gr id w i t h t h e ' f a t i g u e c r i t e r i o n ' stress e x p r e s s e d as t h e y c o - o r d i n a t e . Each map i s n o r m a l i s e d t o t h e smooth c a s e by a s c a l i n g f a c t o r . The smooth c a s e r i s k map i s shown i n f i g u r e 6 ( a ) where a s e x p e c t e d t h e h i g h e s t r i s k o c c u r s a t t h e p o s i t i o n of t h e maximum o r t h o g o n a l s h e a r stress, 0 .8b below t h e b e a r i n g s u r f a c e . AS t h e d e n t ( f i g u r e 6 ( b ) ) and t h e d e n t p l u s r e s i d u a l stresses ( f i g u r e 6 ( c ) ) are i n c l u d e d t h e map i s m o d i f i e d , 192 p a r t i c u l a r l y a round t h e d e n t s h o u l d e r s . A s a r e s u l t t h e s c a l e f a c t o r , e f f e c t i v e l y a measure of t h e i n c r e a s e d r i s k , i n c r e a s e s d r a m a t i c a l l y , by a f a c t o r of a lmost 50 on t h e i n c l u s i o n of t h e r e s i d u a l stresses. The p r e d i c t e d l i v e s f o r each d e n t / r o l l e r combination a r e p l o t t e d i n f i g u r e 7 . From t h i s d a t a an approximate map of r e l a t i v e l i v e s can b e c o n s t r u c t e d i n t e r m s of t h e d e n t s i z e , c o n t a c t s i z e and r o l l e r r a d i u s o f c u r v a t u r e ( f i g u r e 8). A s can be s e e n t h e l i f e of t h e smooth r a c e w a y i n c r e a s e s r a p i d l y w i t h d e c r e a s i n g l o a d a n d t h e l i f e g e n e r a l l y i n c r e a s e s w i t h i n c r e a s i n g r a d i u s . The i n f l u e n c e of t h e dent wi thout t h e a s s o c i a t e d r e s i d u a l stresses on l i f e i s m i n i m a l . A s i g n i f i c a n t l i f e r e d u c t i o n o n l y occurs f o r t h e s m a l l e s t r a d i u s u n d e r t h e t h r e e l i g h t e s t l o a d s . T h i s c h a n g e s d r a m a t i c a l l y when t h e ( t e n s i l e ) r e s i d u a l stress f i e l d below t h e d e n t i s t a k e n i n t o a c c o u n t . The r e s i d u a l stresses were o b t a i n e d assuming no r a d i a l t r a c t i o n f o r c e a t t h e i n t e r f a c e , i e a f l a t p r e s s u r e d i s t r i b u t i o n . Under h i g h l o a d s t h e i n f l u e n c e o f t h e r e s i d u a l stress f i e l d s a r e r e l a t i v e l y s m a l l , b u t a s t h e l o a d i s reduced t h e l i v e s d e c r e a s e m a r k e d l y r e l a t i v e t o t h e smooth c a s e s . A s t h e r a d i u s o f t h e c o n t a c t i s i n c r e a s e d t h e i n f l u e n c e o f t h e r e s i d u a l s t resses a n d o f t h e d e n t g e o m e t r y i s diminished. I n s p e c t i o n o f t h e o r t h o g o n a l s h e a r stress c o n t o u r s ( f i g u r e 9) h e l p s e x p l a i n why t h i s might b e s o . A t h i g h e r l o a d s a l t h o u g h t h e o r t h o g o n a l s h e a r stresses of t h e m o d i f i e d H e r t z i a n f i e l d a r e of a h i g h e r magnitude, t h e maxima a r e s i t u a t e d w e l l below t h e s u r f a c e . The stress c o n c e n t r a t i o n s from t h e shoulder of t h e d e n t and p a r t i c u l a r l y t h e t e n s i l e stresses of t h e r e s i d u a l stress f i e l d l i e much more c l o s e l y t o t h e s u r f a c e and c a n n o t t h e r e f o r e combine w i t h them t o g e n e r a t e h i g h v a l u e s of t h e f a t i g u e c r i t e r i o n . A s t h e l o a d i s reduced t h e stress c o n t o u r s l i e more c l o s e l y t o t h e s u r f a c e a n d c a n combine w i t h t h e t e n s i l e r e s i d u a l stresses t o cause more damage. The r e s u l t s s h o u l d be i n t e r p r e t e d w i t h some c a u t i o n because of t h e s i m p l i f i c a t i o n s made and t h e c o a r s e g r i d numerics and s t r ic t q u a n t i t a t i v e c o n c l u s i o n s s h o u l d n o t b e made. However , t h e q u a l i t a t i v e r e s u l t s a r e i n t e r e s t i n g enough t o c o n t i n u e t h e r e s e a r c h i n t h i s d i r e c t i o n , i n c o r p o r a t i n g more r ea l i s t i c d e n t shapes , r e s i d u a l stress f i e l d s and more a c c u r a t e c a l c u l a t i o n s on f i n e r g r i d s . 6 CONCLUSIONS The e f f e c t of b o t h d e n t s i z e and s u b s u r f a c e r e s i d u a l stresses have been added t o t h e o r i g i n a l l i f e r e d u c t i o n work. The s l i p l i n e f i e l d a n a l y s i s c a n o n l y b e r e a l i s t i c a l l y a p p l i e d t o deep t r a n s v e r s e i n d e n t a t i o n s where t h e assumptions of p l a n e p l a s t i c s t r a i n can be j u s t i f i e d and t h e d r y c o n t a c t a n a l y s i s i s p r o b a b l y o n l y r e a s o n a b l e f o r r e l a t i v e l y t h i n f i l m c o n d i t i o n s . However t h e t r e n d s i n t h e l i f e r e d u c t i o n f a c t o r s are d i s t i n c t i v e and t h e e x p l a n a t i o n f o r them would s e e m reasonable and a p p l i c a b l e t o a n y d e n t p r o f i l e / r o l l e r combinat ion. The most s t r i k i n g outcome i s t h a t where f a i l u r e i s i n i t i a t e d t h r o u g h s u r f a c e i n d e n t a t i o n and a s s o c i a t e d r e s i d u a l stresses (and c o n s i d e r a b l e e v i d e n c e e x i s t s t o s u g g e s t t h i s i s s o ) t h e e x p e c t e d l i v e s may n o t i n c r e a s e w i t h d e c r e a s i n g l o a d a s r a p i d l y a s would b e p r e d i c t e d by convent iona l models. The r e d u c t i o n i n expec ted l i v e s i s v e r y s e n s i t i v e t o t h e s i z e of d e n t i n r e l a t i o n t o t h e r o l l e r r a d i u s a n d t h i s may w e l l have i m p o r t a n t consequences i n t e r m s o f c r i t i c a l p a r t i c l e s i z e and s a f e and u n s a f e l e v e l s of f i l t r a t i o n . To b e a b l e t o draw q u a n t i t a t i v e c o n c l u s i o n s f u r t h e r r e s e a r c h i s needed t h a t uses more r e a l i s t i c r e s i d u a l stress f i e l d s and f i n e r grids i n t h e l i f e c a l c u l a t i o n s . 7 ACKNOWLEDGEMENTS W e would l i k e t o thank D r Andrew Olver f o r h i s h e l p f u l a d v i c e i n t h i s work and t o register our g r a t i t u d e t o SKF-ERC, The Nether lands, who have s p o n s o r e d t h i s work, and t o D r I a n Leadbetter, Managing D i r e c t o r of SKF-ERC f o r permiss ion t o p u b l i s h . APPENDIX References 31 4 1 91 Webster, M . N . , Ioannides , E . and S a y l e s , R . S . , ( 1 9 8 5 1 , \"The E f f e c t o f T o p o g r a p h i c a l D e f e c t s on t h e C o n t a c t Stress and F a t i g u e L i f e i n R o l l i n g Element Bearings\" , Proceedings of t h e 1 2 t h LeedsLyon Symposium on T r i b o l o g y , Lyon, But te rwor ths , Vo1. 12, pp. 121-131. Lundberg, G . and Palmgren, A . , ( 1 9 4 7 ) , \"Dynamic C a p a c i t y o f R o l l i n g Bear ings \", Acta P o l y t e c h n i c a , Mechanical Engineer ing series , Royal Academy o f E n g i n e e r i n g Sc iences , Vol. 1, No 3, 7 . I o a n n i d e s , E . and H a r r i s , T . A. , (1985) , \"A N e w F a t i g u e L i f e Model f o r R o l l i n g B e a r i n g s \", ASME J o u r n a l of L u b r i c a t i o n Technology, , Vol. 107, pp. 367-378. Hamer, J . C . , Sayles , R. S . and Ioannides , E . , ( 1 9 8 5 ) , \"Deformation Mechanisms and S t r e s s e s C r e a t e d by 3 r d Body D e b r i s C o n t a c t s and T h e i r E f f e c t s on R o l l i n g Bear ing F a t i g u e \", Proceedings of t h e 1 4 t h Leeds-Lyon Symposium on Tr ibology, Lyon, But te rwor ths , Vol. 1 4 . Hamer, J. C . , Sayles , R. S. and Ioannides , E., \" P a r t i c l e Deformation and Counter face Damage When R e l a t i v e l y S o f t P a r t i c l e s a r e S q u a s h e d Between Hard A n v i l s \" , T r a n s ASME/STLE t o be p u b l i s h e d . Dumas, G . and B a r o n e t , C . N . , ( 1 9 7 1 ) , \" E l a s t o - p l a s t i c i n d e n t a t i o n of a h a l f - s p a c e b y a l o n g r i g i d c y l i n d e r \" , I n t e r n a t i o n a l J o u r n a l o f M e c h a n i c a l Sc iences , Vol. 13, 519. Olver , A . V . , (19861, \"Wear of Hard S t e e l i n L u b r i c a t e d , R o l l i n g C o n t a c t \" , Phd Thes is , I m p e r i a l Col lege . Olver , A . V . , Sp ikes , H . A . , Bower, A . and Johnson, K . L . , ( 1 9 8 6 ) , \"The R e s i d u a l S t r e s s D i s t r i b u t i o n i n a P l a s t i c a l l y Deformed Model Asper i ty\" , Wear, Vo1. 107, H i l l , R . , (1950) , \"The Mathematical Theory pp. 151-174. o f P l a s t i c i t y \", Oxford U n i v e r s i t y Press. 1 0 1 Ford , H . , ( 1 9 6 3 ) , \"Advance Mechanics of Mater ia l s\" , Longmans . 113 Lubrecht , A . A . , \"The Numerical S o l u t i o n of t h e Elas tohydrodynamica l ly L u b r i c a t e d L i n e and P o i n t C o n t a c t Problem, Using M u l t i g r i d Techniques\", Phd Thes is , Twente U n i v e r s i t y , l 9 8 7 , The Nether lands . 193 [121 Brandt, A. and Lubrecht, A. A., 'Multilevel Multi-Integration and Fast Solution of Integral Equations\", to be published in the Journal of Computational Physics. [131 Brandt, A. , \"Multigrid Techniques: 1984 Guide with Applications to Fluid Dynamics\", monograph available as GMD studien 85 from GMD postfach 1240 , Schloss Birlinghofen D5205 St. Augustin 1 BRD. coefficient (p) of: (a) p = 0 and (b) p = 0.2. %/A 000 ? Figure 4 Residual orthogonal shear stress (Q contour map. 194 shear stress (2,) distribution during the overrolling of the dent. 195 Figure 6(c) Risk map of indented surface plus residual stresses. Scale factor49.7 196 stresses. 197 n 9 17 33 65 129 257 513 1025 time 0.8 1.5 2.4 3.9 8.4 23.0 79.0 306.0 time * 6.2 12.3 24.3 48.4 Table 1: Computing time as a function of the level (L), the number of points n, in seconds on a VAX 785, for the smooth dry line contact problem, with (*) and without MLMI." ] }, { "image_filename": "designv11_20_0002821_j.sna.2010.07.002-Figure11-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002821_j.sna.2010.07.002-Figure11-1.png", "caption": "Fig. 11. The stator construction.", "texts": [ " The hysical parameter of motor which is fabricated, is illustrated in able 1. The piezoelectric ceramic used in the prototype is PZTb(Zr/Ti)O3 (C-213 model), provided by fuji ceramic Co., Ltd. mplitude of vibration is calculated by strain constant of piezoelecric. Input voltage of piezoelectric is V0\u2212p = 100 V. Fig. 10 shows the olarization pattern of piezoelectric ceramic. The vibration mode f the stator is B (3, 1). The stator is made of yellow brass. The onstruction of the stator is shown in Fig. 11. A PCB circuit is bonded on the steel backing and piezoelectric ing is bonded on the PCB circuit, by epoxy. Electrode patterns n PCB conduct the electricity to the piezoelectric ring segments. ibrating brass ring (stator) is bonded on the piezoelectric ring by ltrasonic epoxy (UHU plus epoxy) which is strong and high fatigue ife cycle epoxy. Fig. 12 shows the cross-section of the RIUSM. The mechanical characteristics of the motor with various freuencies and amplitude of the applied voltage have been calculated sing the proposed analytical model" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000457_bfb0075007-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000457_bfb0075007-Figure9-1.png", "caption": "Figure 9", "texts": [ " Figure 8 is a picture of this slice. The points numbered 1 through 18 on the left are identified with the points labelled 1 through 18 on the right. The thick line represents the \"center line\" of the \"ribbon disc\" on which we must do double Dehn surgery. Notice that points where the thick line crosses itself are not indicated as either overcrossings or undercrossings. This is because, by general position, we can choose them to be either one or the other and we wish to make the choices later. Now we explain how figure 9 is obtained from figQre 8. First, we do the double Dehn surgery. This has the effect of replacing the surgery ribbon disc by a ribbon lying flat in the plane, whose connections to the thin line, at the two \"T\" parts are indicated by the inserts. Note that these \"connections\" are exactly the same as those indicated in figure 7. Rather than represent the boundary of the ribbon as two thin lines we continue to represent it as one thick line. Also, the image of the axis of rotation, represented as a point in figure 7 has been moved to the right and \"laid down\" over the other component of the link. Also certain of the ribbon identifications have been made (1, 6, and 11). Making such an identification may have the effect of introducing a twist in the band - The band as pictured in figure 9 has two full right twists in it. In obtaining figure 9 from figure 8 we have committed ourselves to choosing certain overcrossings. The remaining figures are obtained, one after the other, by making identifications, choosing overcrossings, and keeping track of the number of twists in the band. Since these methods apply to any link all of whose components are unknotted we have the following result. PROPOSITION 3.1. Any link, all of whose components are unknotted, occurs as a sublink of a link whose n components are cyclically permuted by an n-fold rotation", " Let L be branched covering space map preimage of the branch set, any n-component link in 53. There p:s3 s3, branched over a knot k, -1 p (k), contains L as a sublink. is a 2n + 3 to such that the 37 9... 8\u2022\u2022 , 1 ... -. . .. o.... '\" 1,. ' .. 2. .... 3- ... 4\u2022 \u2022 Trivial knot orthogonal to the plane figure 8 38 - -'t-r- ';:::.l:>----+=-et===t:+===I- ?-1- -1 ;==========t==t=-i---tli-_a... ?C=========i==I=t+==1- Ithe band c +2Number of turns of =========- ---= =====t=t====-JE. It- \" - -13l't-:=================_/6 figure 9 39 J \"f ... r> r: h \":, - _ r ;::. :> f - q - 1(; .... . 11 Ir 16 1[' /6 /0 Number of turns of the band = +4 figure 10 40 f(-- '/- J- fr -( \"\"\"\"\"'I r:. - - b I I\"\" , _I.> - 2) with the forward horizontal (Fig. 5). V20 is vertical, and V2 is always smaller than V20. sin A cos5A cos <& ? = sin 0o = sin (E-A) cos2A 2 cos E sin(E-A) [cos E \u2014sin E sin A cos A cos E (2 cos2E-cos2A)] (12) (13) From Eq. (12) and the limits on E and A, 2 is limited to 0\u00b0 <2< 33.74\u00b0 and ir-4L=A0>A. For an optimal F-type transfer, the final length of the primer vector must be less than or equal to one, hence we must have E>/(A) instead of E=/(A), and thus the equality in Eq. (10) is replaced by a \"greater than or equal\" sign. RF Transfer This transfer is similar to the F transfer except the vehicle grazes the forbidden sphere. This has the effect of dividing the trajectory into the two parts, before and after graze. The primer vector is analyzed separately on the two sections. A Fig. 6 The RF transfer. Fig. 7 The PNP transfer. general result is that the graze must occur before the impulse (for the convention that V2 < V } ) . An analysis similar to that for the F transfer applies, and the illustration for the optimal configuration of the single impulse is given in Fig. 6 in which the points with the subscript 0 correspond to the unconstrained optimal F transfer. The direction of the impulse to change from the grazing hyperbolic branch to the final hyperbolic branch of the transfer must be equal to A for the unconstrained F transfer as shown in Fig. 6. The optimality of the transfer requires the point R l to be at a greater distance from M than the point R w. Transfers of Type RF oo This transfer is a combination of the two previous cases: E and A must satisfy Eq. (9), V] and V2 must verify the construction of Fig. 6 and the second impulse (at infinity) has a direction ( 2 ) given by Fig. 5 andEqs. (12) and (13). PNP Transfer The PNP transfer, or biparabolic transfer, is the adaptation to the case of a planet with finite radius of the six infinitesimal impulse transfer8 of the unconstrained transfer problem. The PNP transfer is shown in the accompanying figure (Fig. 7). The vehicle is first put on a grazing trajectory by the customary infinitesimal impulse at infinity. At periapse, at the first graze, a tangential braking impulse of magnitude (U l \u2014L) leads to the parabolic trajectory shown. At infinity, an infinitesimal impulse i2 leads to a circular orbit which is traversed until the proper orientation is reached. Then a second infinitesimal impulse i3 places the vehicle on the grazing parabolic trajectory which has a common periapse with the grazing hyperbola of the desired asymptotic velocity. At this common periapse, a tangential impulse I4 of magnitude (U2\u2014L) accelerates the vehicle and leaves it on the desired departure trajectory. D ow nl oa de d by U N IV E R SI T Y O F O K L A H O M A o n Fe br ua ry 4 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /3 .6 04 95 AUGUST 1975 OPTIMAL TRANSFERS BETWEEN HYPERBOLIC ASYMPTOTES 983 Fig. 8 The Roo transfer. The characteristic velocity of the PNP transfer is (U 1 + U2 \u2014 2L), and is independent of turn angle A. It can also be demonstrated that where a] + a2Vj. These transfers are of six different types which are always grazing: 1) Type PNP: identical to that of the less than optimal deviation case and always optimal for A = 180\u00b0; 2) Type Roo: transfer with one impulse at infinity on the side of the greater velocity; 3) Type ooRoo: transfer with two impulses at infinity, one on either side of the grazing passage; 4) Type RF: transfer with one impulse at a finite distance after the grazing passage; 5) Type RFoo: transfer with two impulses after the grazing passage, one at a finite distance and the other at infinity; and 6) Type ooRF: transfer with two impulses, one at infinity before the grazing passage and the other at a finite distance after the passage. If V2 < V} the types Roo, RF, RFoo, and ooRF become ooR, FR, ooRF and RFoo. Transfer of Type Roo This transfer begins as a grazing hyperbola H and ends at infinity with the impulse7; = V2\u2014 V20\\ Fig. 8. The impulse is easy to calculate. The direction of 11 is located by the angle 4>2 with the forward horizontal and the optimality requires arctan <2 7 ) and 2, with the local forward horizontal (measured clockwise if Hl is the counterclockwise direction and conversely). The optimality requires [2V 10L2/UW(L2+ 2V10 2) (16) Transfers of Type RF Let us use Fig. 10, similar to Figs. 4 and 6. As in the \"Less Than Optimal Deviation Angle Case\" the theory of optimization of Pontryagin4 leads to a very simple result: 1) the grazing passage must occur before the impulse; 2) the velocity Vj =IR j of arrival at the impulse, corresponds to a point R ] nearer to M than R 2 with the velocity of departure given by V2 = IR2. In this case it is possible to have \u00a3= (180-.40/4) Fig. 9 The ooRoo transfer. Fig. 10 The RF transfer. Fig. 11 The RFoo transfer. Fig. 12 The ooRF transfer. ,2 - V 2 ,2 J I O \" V I O + greater than 45\u00b0 but (2 cos 2L \u2014cos2 A) remains positive and the relation \u00a3>/( A) is still valid, hence 0\u00b02 with the local forward horizontal (Fig. 11), 2<35.264\u00b0 (18) D ow nl oa de d by U N IV E R SI T Y O F O K L A H O M A o n Fe br ua ry 4 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /3 .6 04 95 AUGUST 1975 OPTIMAL TRANSFERS BETWEEN HYPERBOLIC ASYMPTOTES 985 I.I55< Fig. 18 The optimal transfers for F7/L = 1.45, A / ) the angle which locates the direction of the first impulse, measured as usual from the local forward horizontal and positive in the clockwise direction if the transfer is in the counterclockwise direction and conversely, (Fig. 12), and let us call X and Y the radial and circumferential components of the velocity of arrival V'} at the second impulse I2 (X and Y are both positive). By the theory of Lawden7 the transfer must satisfy the following two equations: K/0cos ,= V2 cos5A sin A 2 cos \u00a3 sin (E-A) (19) Y cos A sin A I (1- Y2 2L2Ycos/ U10 2(YU10+VW 2~XV]0) (20) Thus any RF transfers which satisfy these equations satisfy the optimality requirements of the ooRF transfer. The optimality also requires that 0\u00b0 < < / > / < arcsin V3/3 = 35.264\u00b0. D ow nl oa de d by U N IV E R SI T Y O F O K L A H O M A o n Fe br ua ry 4 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /3 .6 04 95" ] }, { "image_filename": "designv11_20_0000087_1.1904983-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000087_1.1904983-Figure6-1.png", "caption": "Figure 6. CVs showing the reversible couple due to the one-electron reduction of 4-MNImOH in nonaqueous medium at different sweep rates. Arrow indicates the scan direction.", "texts": [ " This corresponds to the two-electron oxidation of the hydroxylamine derivative, which forms during the first negative-going sweep, to form the nitroso derivative R-NHOH \u2192 R-NO + 2e\u0304 + 4H+ f5g Adjusting the switching potential appropriately, we can study the nitro/nitro radical anion couple sRNO2/RNO2 \u2022\u2212d in isolation ~dashed line in Fig. 5!. To complete the electrochemical characterization to study the properties of the nitro radical anion in the absence of protonation reactions, we also studied an aprotic medium. In a totally nonaqueous medium containing 100% DMF with 0.1 M TBAP, we obtained a perfectly isolated couple ~Fig. 6! corresponding to the one-electron reduction of the nitroimidazole parent compound to produce the nitroimidazole radical anion derivative. The generation of the nitro radical anion derivative was also characterized by ESR. The radical was prepared in situ by controlled potential electrolysis at \u2212900 mV vs. Ag/AgCl in DMF. The nitroimidazole free radical displays a well-resolved ESR spectrum ~Fig. 7!. The interpretation of the ESR spectra led us to determine the coupling constants for all address" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000082_2004-01-2830-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000082_2004-01-2830-Figure2-1.png", "caption": "Figure 2: A framework design by means of the developed Box-joint system", "texts": [ " The preliminary name of this new design was \u201cBox-joint\u201d. The new Box-joint construction kit uses standard steel beams with square or rectangular cross-sections in a modular series of measures. This is to enable the required stability at the lowest possible cost. The joints of the system press the beams together by means of bolts and pressure plates to keep the beams using friction, in specified positions. Different sizes of beams can be combined and moved to feasible perpendicular positions relative to each other as is illustrated in figure 2. The pressure plates belonging to the system are also used to attach the dynamic modules in feasible positions along the beams of a framework as is indicated in figure 3. Thus, varied functions of the framework are enabled by means of a minimum number of different construction modules. By means of calculations and experimental tests the applicability and stability of the developed construction kit for static framework has been verified. Further more, the use of air-cushions for the transport of the static framework on a floor is recommended" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002738_9780470549148.ch6-Figure6.8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002738_9780470549148.ch6-Figure6.8-1.png", "caption": "Figure 6.8 Reaction forces and velocities at a joint center during dynamic activity. The dot product of the force and velocity vectors is the mechanical power (rate of mechanical energy transfer) across the joint.", "texts": [ "9 Mechanical Energy Transfer between Segments Each body segment exerts forces on its neighboring segments, and if there is a translational movement of the joints, there is a mechanical energy transfer between segments. In other words, one segment can do work on an adjacent segment by a force displacement through the joint center (Quanbury et al., 1975). This work is in addition to the muscular work described in Sections 6.0.4 to 6.0.7. Equations (6.5) and (6.6) can be used to calculate the rate of energy transfer (i.e., power) across the joint center. Consider the situation in Figure 6.8 at the joint between two adjacent segments. Fj1, the reaction force of segment 2 on segment 1, acts at an angle \u03b81 from the velocity vector Vj . The product of Fj1Vj cos \u03b81 is positive, indicating that energy is being transferred into segment 1. Conversely, Fj 2Vj cos \u03b82 is negative, denoting a rate of energy outflow from segment 2. Since Pj1 = \u2212Pj 2, the outflow from segment 2 equals the inflow to segment 1. In an n-joint system, there will be n power flows, but the algebraic sum of all those power flows will be zero, reinforcing the fact that these flows are passive and, therefore, do not add to or subtract from the total body energy", " (viii) Hip joint for frame 50 (ix) Hip joint for frame 70 (x) Hip joint for frame 4. (b) .(i) Scan the listings for muscle power in Table A.7 and identify where the major energy generation occurs during walking. When in the gait cycle does this occur and by what muscles? (ii) Do the knee extensors generate any significant energy during walking? If so, when during the walking cycle? (iii) What hip muscle group generates energy to assist the swinging of the lower limb? When is this energy generated? 5. Using Equation (6.10) (see Figure 6.8), calculate the passive rate of energy transfer across the following joints, and check your answers with Table A.7. From what segment to what segment is the energy flowing? (i) Ankle for frame 20. (ii) Ankle for frame 33. (iii) Ankle for frame 65. (iv) Knee for frame 2. (v) Knee for frame 20. (vi) Knee for frame 65. (vii) Hip for frame 2. (viii) Hip for frame 20. (ix) Hip for frame 67. 6. .(a) Using equations in Figure 6.19b, carry out a power balance for the foot segment for frame 20. (b) Repeat Problem 6(a) for the leg segment for frame 20" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.66-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.66-1.png", "caption": "Fig. 14.66 Single-cabin person and material lift with rackand-pinion hoisting gear", "texts": [], "surrounding_texts": [ "Material and equipment lifts with access to personnel are intended for the vertical transport of persons and materials during construction/assembly works and repairs of mainly high-rise buildings in housing and industrial construction. Their design is usually similar to that of material and equipment lifts with a rack-andpinion hoisting gear. A person and material lift consists of a cabin with a rack-and-pinion drive, moving on a mast secured at the bottom to the lift\u2019s base and anchored to the building\u2019s wall, and transport stages (stops) between which transport takes place. The mast has a segmental structure and can be extended by adding mast sections. It is anchored by means of a system of tubes, which makes it possible to adjust the mast\u2019s position relative to the building\u2019s wall. The lift can have two cabins, each with its own drive system, whereby the transport of persons and materials can be doubled. The lift\u2019s cabins move on a common mast independently of each other. Examples of person Part B 1 4 .4 Table 14.6 Specifications of selected person and material and equipment lifts Parameters Single-cabin lifts Two-cabin lifts Lifting capacity (kg) 1200\u20132000 2 \u00d7 1200\u20132 \u00d7 2000 Maximum lifting height (m) 150\u2013300 150\u2013300 Maximum number of persons 15\u201325 15\u201325 depending on lifting capacity depending on lifting capacity Lifting speed (m/min) \u2248 40 \u2248 40 Mast section (m) 1.5 1.5 Distance between anchors (m) 12.0\u201315 12.0\u201315 Lifting height without anchoring (m) 12.0\u201315 12.0\u201315 Davit\u2019s lifting capacity (kg) 150 150 Gripping device Yes Yes Cabin\u2019s dimensions \u2013 Height (m) 2.1\u20132.7 2.1\u20132.7 \u2013 Width (m) 1.3\u20131.5 1.3\u20131.5 \u2013 Length (m) \u2248 3.0 \u2248 3.0 Electric specifications \u2013 Supply voltage and frequency (V/Hz) 230\u2013400/50 230\u2013400/50 \u2013 Motor power (kW) 2 \u00d7 9.0\u20132 \u00d7 11 2 \u00d7 2 \u00d7 9.0\u20132 \u00d7 2 \u00d7 11 and material lifts with a rack-and-pinion, platform lifting gear are shown in Figs. 14.66, 14.67 and Table 14.6. The development of high-rise construction created a need for high-lifting-speed vertical transport equipment. For this purpose fast lifts with a lifting speed of up to 1.8 m/s are employed. These are used for the vertical transport of persons and materials and equipment in industrial construction, for building reinforced concrete chimneys, silos, television towers, and similar structures. One- and two-mast high-speed construction lifts are available. These lifts incorporate electrohydraulic drive systems whose basic unit is a hydrostatic gear. The hydraulic engine\u2019s output shaft is connected by a cou- Part B 1 4 .4 pling to the shaft of a worm gear assembly on the output shafts of which cylindrical gears mating with the mast\u2019s rack, and so making the cabin move up or down, are mounted. The weight of the loaded cabin is counterbalanced by a counterweight connected to the cabin by a steel cable passing through pulleys fixed to the top of the mast. The safety of persons in person and material lifts is ensured by appropriate guards around the traffic way and the stop platform access. Person and material lifts can be controlled from the moving cabin as well as from any stop level. The lift\u2019s base (the bottom stop) is fenced in to a height not less than 2.0 m. The access areas are protected by stop doors with a minimum height of 2.0 m, equipped with safety locks. Entrances to the person and material lifts\u2019 cabins are protected by doors. The cabin door is equipped with mechanical bolting devices and safety cutout switches, preventing the door from being opened as the cabin is moving when the cabin\u2019s floor is not within about \u00b10.25 m from a stop landing or when the door is not closed and the bolting device is not in the closed position. In addition, person and material lifts are fitted with similar safety devices as those used in rack-and-pinion material and equipment lifts, i. e.: \u2022 Gripping devices actuated at an excessive speed of descent\u2022 Protection against disengagement of the drive toothed wheel from the mast\u2019s rack\u2022 An emergency lowering system\u2022 Limit switches\u2022 Working platform skewness switches in two-mast lifts\u2022 Protection against electrical failures in the case of no voltage, voltage decay, or voltage drop\u2022 Electrical devices protecting: \u2013 Closed stop gate position \u2013 Stop gate bolting device position \u2013 Closed cabin or platform gate position \u2013 Bolted emergency hatch or door position Modern rack-and-pinion person and material and equipment lifts are characterized by: Multisectional mast Cabin with drive system Fencing Cabin with drive system Anchoring connectors Assembly davit Lift\u2019s base \u2022 A great lifting height: up to 300 m\u2022 Easy and quick assembly\u2022 A high lifting capacity: 2000 kg per cabin\u2022 Automatic stopping of the cabin at the terminal stops\u2022 The possibility of programming at which stops the cabin should stop\u2022 Completely safe operation owing to the use of appropriate protective measures\u2022 Easy operation and simple maintenance\u2022 An installation that enables audio communication between the cabin and the bottom stop\u2022 An overload control system Because of their advantages person and material lifts with a rack-and-pinion hoisting gear have gained a dominant position in the lift market." ] }, { "image_filename": "designv11_20_0002565_j.tws.2008.08.010-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002565_j.tws.2008.08.010-Figure2-1.png", "caption": "Fig. 2. Detailed design of the clutch disc:", "texts": [ " The clutch capability in reducing gear rattle noise has been theoretically investigated with reference to the influence on the phenomenon exerted by some clutch parameters, such as multivalued spring, dual mass flywheel and hysteresis rates [19]. Audible disturbance known as automotive squeal noise observed during the sliding phase of clutch engagement has been largely studied [20,21]. As seen above, the necessity of coupling or decoupling the engine and transmission during gearshift induced the development of optimized clutch components that aim to transmit the torque between the pressure plate and the flywheel. In this work, we investigate the behavior of the clutch disc (Fig. 2). It allows a soft gradual re-engagement of torque transmission. This progressive re-engagement obtained by the clutch disc characteristics in the axial direction preserves the driver\u2019s comfort and avoids mechanical shocks. It also plays the role of a damper through the springs disposed around the hub. They enable the clutch disc to filter the torque variations of the combustion engine (Fig. 2). The axial elastic stiffness of the clutch disc is obtained by a cushion disc (Fig. 3(a) and (b)) which is a thin waved sheet, located between the two facings and fixed by rivets. It acts like a spring allowing a soft gradual re-engagement. This nonlinear axial stiffness is obtained by cutting the cushion disc into paddles and forming them to get the wavy shape (Fig. 4). The nonlinear axial elastic stiffness of the cushion disc is described by the cushion curve (Fig. 5). This load\u2013deflection curve gives the axial load versus axial displacement obtained by compressing a cushion disc between two flat pressure plates" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003845_1.4003270-Figure11-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003845_1.4003270-Figure11-1.png", "caption": "Fig. 11 Planar metamorphic mechanism: state 2", "texts": [], "surrounding_texts": [ "p i\nL t\nT l j\nw R t l m o T\n0\nDownloaded Fr\nIf this mechanism state matrix were input into the computer rogram, the augmented mechanism state matrix would be given\n1\n1 1 4\n2 3\n5\n6\n87\nY\nX\nFig. 7 3RRR: state 2\n1\n1 1 4\n2 3\n5\n6\n8 7\nY\nX\nFig. 8 3RRR: state 3\nn Eq. 21 as\nhe mechanism state matrix is given in Eq. 24 as\n11012-6 / Vol. 3, FEBRUARY 2011\nom: http://mechanismsrobotics.asmedigitalcollection.asme.org/ on 02/03/2\nMASM = 2Z R,6Z R,8Z R 3Z R 4Z R 5Z R,7Z R 6Z R 8Z R 3 2Z R,6Z R,8V X 3Z R 4Z R 5Z R,7Z R 6Z R 8Z R 2\n2Z R,6Z R,8V X 3Z R 4Z R 5Z R,7Z R 6Z R 8V X 1\nThe augmented mechanism state matrix identifies how the DOF of this mechanism changes as it moves from one state to another. Additionally, it is easy to track how a link changes as the mechanism changes states. For instance, link 8 changes from a movable link in state 1 to a fixed link in states 2 and 3.\n6.2 Nine-Bar Linkage. Mechanism state matrices can be particularly useful when analyzing complex mechanisms in which the DOF of the mechanism is not obvious upon inspection. Such a mechanism is the nine-bar linkage in Fig. 9. In this example only one state is considered, and the mechanism state matrix is given in\n1 2\n3 9 8 5\n4 6\n7\nY\nX\nFig. 9 Nine-bar linkage\n1 1\na d b\n2\n4 53 P2\nP1\nc\n2\nY\nX\nEq. 22\nMSM = 2Z R,8Z R,9Z R 3Z R,5Z R 4Z R 5Z R,6Z R 7Z R 8Z R,9Z R 22\nink 1 is specified to be a fixed link. It is unknown which links and joints, if any, become fixed as a result of link 1 being chosen as he fixed link. The output augmented mechanism state matrix is given in Eq. 23 as\nMASM = 2Z R,8Z R,9Z R 3Z R,5Z R 4Z R 5Z R,6Z R 7Z R 8Z R,9Z R 2 23\nhe augmented mechanism state matrix specifies that the nine-bar inkage has 2 DOF. Link 1 is the only fixed link, and all of the oints are movable.\n6.3 Planar Metamorphic Mechanism. The third example ill be that of the planar metamorphic mechanism analyzed in ef. 16 and shown in Figs. 10\u201312. This is a five-link mechanism\nhat oscillates between pins P1 and P2. The spring embedded in ink 2 pushes link 3 along the slot in link 2. In every state the\nechanism is a five-bar linkage with one DOF. However, the state f the mechanism changes as it oscillates between pins P1 and P2.\nMSM = 2Z R,5V X 3V P 4Z R 5Z R 2Z R,5Z R 3V X 4Z R 5Z R\n2Z R,5V X 3V P 4Z R 5Z R\n24\nThe augmented mechanism state matrix for this example is given in Eq. 25 as\nMASM = 2Z R,5V X 3V P 4Z R 5Z R 1 2Z R,5Z R 3V X 4Z R 5Z R 1\n2Z R,5V X 3V P 4Z R 5Z R 1\n25\nIn this case, the DOF does not change as the mechanism changes states. However, the augmented mechanism state matrix shows how the links and joints change as the mechanism moves from one state to another.\nTransactions of the ASME\n016 Terms of Use: http://www.asme.org/about-asme/terms-of-use", "u c a e f m g\na n\nJ\nDownloaded Fr\n6.4 2R-2P Mechanism. Mechanism state matrices can be sed to identify topologically identical mechanisms with different onfigurations. Consider the 2R-2P mechanisms shown in Figs. 13 nd 14. These are topologically identical mechanisms with differnt configurations. The difference between them is evident by orming the augmented mechanism state matrices. The augmented echanism state matrix for the 2R-2P mechanism in Fig. 13 is iven by\nMASM = 2Z R,4X P 3Z R 4Y P 1 26 nd the augmented mechanism state matrix for the 2R-2P mecha-\n1 1 a d\nc\nb\n2\n4\n53\nP2 P1\nY\nX\nism in Fig. 14 is given as\nournal of Mechanisms and Robotics\nom: http://mechanismsrobotics.asmedigitalcollection.asme.org/ on 02/03/2\nMASM = 2V X,4X P 3V X 4X P 1 27 Equation 26 shows that either links 2, 3, or 4 could be an input link. This is not the case for the 2R-2P mechanism with parallel prismatic joints. As shown in Eq. 27 , only link 4 can be an input link. The other three links are fixed. It should be noted that the 2R-2P mechanism with parallel prismatic joints is a special case in which the configuration of the mechanism causes three of the four links to be fixed.\n7 Conclusions This paper introduced mechanism state matrices as a novel way to represent the topological characteristics of reconfigurable mechanisms. It was shown that mechanism state matrices can be used to calculate the DOF of planar mechanisms containing only one DOF joints. This includes mechanisms that contain partially locked kinematic chains. The DOF matrix can then be combined with the mechanism state matrix to form an augmented mechanism state matrix. Augmented mechanism state matrices can be used to represent all five topological characteristics of reconfigurable mechanisms. Future work will involve applying mechanism state matrices to aid in both systematic analysis and synthesis of reconfigurable mechanisms.\nReferences 1 Liu, N., 2001, \u201cConfiguration Synthesis of Mechanisms With Variable\nChains,\u201d Ph.D. thesis, National Cheng Kung University, Tainan, Taiwan. 2 Yan, H., and Liu, N., 2000, \u201cFinite-State-Machine Representations for Mecha-\nnisms and Chains With Variable Topologies,\u201d ASME Paper No. DETC2000/ MECH-14054. 3 Kuo, C., 2004, \u201cStructural Characteristics of Mechanisms With Variable Topologies Taking Into Account the Configuration Singularity,\u201d MS thesis, National Cheng Kung University, Tainan, Taiwan. 4 Kuo, C.-H., and Yan, H.-S., 2007, \u201cOn the Mobility and Configuration Singularity of Mechanisms With Variable Topologies,\u201d ASME J. Mech. Des., 129 6 , pp. 617\u2013624. 5 Yan, H.-S., and Kuo, C.-H., 2009, \u201cReconfiguration Principles and Strategies for Reconfigurable Mechanisms,\u201d ASME/IFToMM International Conference on Reconfigurable Mechanisms and Robots, pp. 1\u20137. 6 Dai, J., and Qixian, Z., 2000, \u201cMetamorphic Mechanisms and Their Configuration Models,\u201d Chin. J. Mech. Eng., 13 03 , pp. 212\u2013218. 7 Dai, J. S., and Jones, J. R., 2005, \u201cMatrix Representation of Topological Changes in Metamorphic Mechanisms,\u201d ASME J. Mech. Des., 127 4 , pp. 837\u2013840. 8 Zhang, L., Wang, D., and Dai, J. S., 2008. \u201cBiological Modeling and Evolution Based Synthesis of Metamorphic Mechanisms,\u201d ASME J. Mech. Des., 130, p. 072303. 9 Parise, J., Howell, L., and Magley, S., 2000. \u201cOrtho-Planar Mechanisms,\u201d Proceedings of the 26th Biennial Mechanisms and Robotics Conference. 10 Carroll, D., Magleby, S., Howell, L., Todd, R., and Lusk, C., \u201cSimplified Manufacturing Through a Metamorphic Process for Compliant Ortho-Planar Mechanisms,\u201d Proceedings of IMECE2005, Paper No. IMECE2005-82093. 11 Yan, H.-S., and Kuo, C.-H., 2006, \u201cTopological Representations and Characteristics of Variable Kinematic Joints,\u201d ASME J. Mech. Des., 128 2 , pp. 384\u2013391. 12 Dai, J., and Wang, D., 2006, \u201cDifferential Geometry Based Analysis of Synthesis of a Multifingered Robotic Hand With Metamorphic Palm,\u201d ASME Paper No. DETC2006-99532. 13 Ziesmer, J. A., and Voglewede, P. A., 2009, \u201cDesign, Analysis and Testing of a Metamorphic Gripper,\u201d ASME Paper No. DETC2009-87512. 14 Tsai, L.-W., 2001, Mechanism Design: Enumeration of Kinematic Structures According to Function, CRC, Boca Raton, FL. 15 Dai, J., and Jones, J. R., 1999, \u201cMobility in Metamorphic Mechanisms of Foldable/Erectable Kinds,\u201d J. Mech. Des., 121 3 , pp. 375\u2013382. 16 Lan, Z., and Du, R., 2008, \u201cRepresentation of Topological Changes in Metamorphic Mechanisms With Matrices of the Same Dimension,\u201d ASME J. Mech. Des., 130, p. 074501. 17 Yan, H.-S., and Kuo, C.-H., 2006, \u201cRepresentations and Identifications of Structural and Motion State Characteristics of Mechanisms With Variable Topologies,\u201d Trans. Can. Soc. Mech. Eng., 30, pp. 19\u201340. 18 Yan, H.-S., and Kuo, C.-H., 2009. \u201cStructural Analysis and Configuration Synthesis of Mechanisms With Variable Topologies,\u201d ASME/IFToMM Interna-\ntional Conference on Reconfigurable Mechanisms and Robots, pp. 23\u201331.\nFEBRUARY 2011, Vol. 3 / 011012-7\n016 Terms of Use: http://www.asme.org/about-asme/terms-of-use" ] }, { "image_filename": "designv11_20_0002840_iembs.2009.5334055-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002840_iembs.2009.5334055-Figure2-1.png", "caption": "Fig. 2. Prototypes of the tongue control system. (a) Sensor plates, electronics, battery, and charging coil. (b) Encapsulated intra-oral device. Arrangement for securing the device to the teeth is not shown on the picture. (c) Board with micro controller and radio chip used to implement the composite USB device.", "texts": [ " Radio communication Radio communication between devices in the system is performed on the 2.4 GHz Industrial Scientific and Medical (ISM) band. A proprietary radio protocol is used to keep complexity and power consumption as low as possible. The embedded controller receives radio packets, containing raw sensor readings, from the intra-oral device every 33 ms (30 Hz). B. Intra-oral device The intra-oral device consists of 18 inductive sensors connected to a Printed circuit Board which contains a micro controller, a radio chip, and a battery, Fig. 2a. The sensors are placed in two separate areas, a key-pad area in front of the mouth which houses 10 sensors and a mouse-pad area in the back with 8 sensors, Fig. 5. The system is encapsulated in a dental retainer made of standard dental materials, Fig. 2b. Details on the intra-oral device and its use as a tongue interface is presented in [5]. The framework for computer control is outlined in Fig. 3. The intra-oral device transmits raw sensor readings, and the computer only receives standard USB mouse and USB keyboard commands. All functionality is therefore defined by the embedded controller and the composite USB device attached directly to the computer. A. Functionality in the Framework 1) Embedded controller The embedded controller performs signal processing common to all external devices", " To ensure that the user can find the correct sensors only a few of the sensors in the key area are used, Fig. 5. Mouse functionality is also controlled by a finite state machine, which tracks the state of the mouse buttons. Activating the sensor for the left or right mouse button for more than 3 seconds makes the mouse button stick, and allows operations such as drag and drop. Fig. 5 shows sensors associated with the mouse functionality. The software prototype for the USB device was implemented on a USB dongle containing a 8-bit micro controller and 2.4GHz radio (Fig. 2c, [9].) The embedded controller was implemented on a development board also containing an 8-bit micro controller from ATMEL. A. Proof of concept An uncontrolled test showed that the system could be used to control both the keyboard and mouse in both Windows XP, Ubuntu (Linux), and OS X (Mac) using only the built-in drivers. It was possible to type simple sentences and delete characters by activating the sensors with the help of an \u201cactivation pin\u201d, a small stick glued to the ferromagnetic part of the tongue piercing" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002361_gt2007-27314-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002361_gt2007-27314-Figure7-1.png", "caption": "Figure 7. Rotor model of the turbogenerator with slip ring shaft between the NDE- and the end bearing", "texts": [ "asmedigitalcollection.asme.org/ on 10/18/2017 Te driven by a synchronous motor (Fig. 2). The generator is operated at its rated speed. The drive motor only needs to deliver the losses of the generator (around one per cent of the machine\u2019s rated power). During the running test of a hydrogen-cooled turbogenerator with three bearings spiral vibration with increasing magnitude occurred. In the text below the three bearings will be referred to as DE- (driven end), NDE- (non-driven end) and end bearing (Fig. 7). The spiral vibration with increasing magnitude would limit the duration of continuous running in the power plant. Immediately a root cause analysis was started to solve the problem. Various tests with the running generator were performed with the aim to identify the location of the hot spot. In parallel a hot spot calculation model was set up. (b) Horizontal DE-bearing pedestal vibration. The generator was run at speeds below rated speed in order to find out the speed range during which the spiral vibration occurs", " It is assumed that the heat input is proportional to the shaft velocity instead of the shaft displacement, since a friction force changes its direction with the direction of the velocity. Case c: Heat input proportional to shaft acceleration. For rub with a contact force varying due to inertia effects this would be the appropriate relation, e.g. in case the inertia forces of the brush masses themselves would play a role. The heat input is assumed to be proportional to the shaft acceleration. oaded From: http://proceedings.asmedigitalcollection.asme.org/ on 10/18/2017 Ter Figure 7 shows the rotor dynamic model of the turbogenerator. The model of the turbogenerator with brush gear unit consists of the generator rotor (GEN) and the slip ring shaft (SR), supported on three journal bearings in free-standing pedestals, which are modeled as equivalent mass-springdamper systems with representative parameters for the horizontal and vertical direction. In Fig. 7 the locations of the slip ring 1 and 2 and the three bearings (DE-, NDE- and endbearing) are indicated. Natural frequencies and damping ratios Table 2 summarizes the natural frequencies and corresponding damping ratios at rated speed. The generator rotor has four critical speeds nk up to rated speed of 3600 rpm. The 1st bending modes of the generator (GEN) and the slip ring shaft (SR) have U-shape. The 2nd bending modes of the GEN-rotor have S-shape. The natural frequencies of the generator (k=1 to 4) are well below rated speed and have minor influence on the hot spot stability" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.60-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.60-1.png", "caption": "Fig. 14.60 Five-hundred-kilo\u2013capacity lift with cable hoisting drive", "texts": [ " The cable\u2019s end is secured to the carriage, which is equipped with rollers. The carriage suspended on the cable moves along the mast\u2019s guides. Accessories for transferring different materials are fixed to the carriage (Fig. 14.59). The carriage can slew around the mast by an angle of about 90\u25e6, which improves operating safety during the unloading of materials. To facilitate the transport of the lift, its base is equipped with vehicle wheels. Another 500 kg-capacity lift with a cable working platform raising drive is shown in Fig. 14.60. Part B 1 4 .4 The lift consists of the following structural units (Fig. 14.60): \u2022 A base\u2022 A cable winch\u2022 A transport basket\u2022 A carriage\u2022 A multisectional mast\u2022 A head with cable pulleys The basket can slew around the mast by 90\u25e6. The lifts are self-assembled using the provided accessories. Lifts with a cable hoisting drive are equipped with the following control and safety systems: \u2022 Gripping devices activated when the limit speed is exceeded\u2022 Limit switches making it possible to stop the lift at set stop levels\u2022 A carriage (with the platform installed) upper position limit switch\u2022 A supply-failure emergency carriage lowering system\u2022 Stops in the lift\u2019s base Ladder Lifts Ladder lifts are cableways for transferring a load simultaneously in the vertical plane and in the horizontal plane" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002261_j.mechmachtheory.2009.09.006-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002261_j.mechmachtheory.2009.09.006-Figure9-1.png", "caption": "Fig. 9. The points taken into account on the gear.", "texts": [ " Seven points along the tooth width and five points along the tooth height are sufficient to obtain correct results. Finally, a total of thirty five points is necessarily to determine the bending definition matrix. Table 4 Maximal and mean errors for the calculation of displacements for the 2nd comparison. Maximal error 7.3% Mean error 1.3% Fig. 12. Load Sharing. All these points have been imposed as nodes in the meshing. The size of the mesh is small on the teeth having points, in order to have correct results in term of displacement computations. The global mesh can be coarser, as shown in Fig. 8 Fig. 9 presents the different points taken into account in the comparison. The figure corresponds to the points on the spiral bevel gear. All the different points have the same parameter L20. Only parameter L10 differs from one point to another. The principle is the same for studying the pinion. Applying a unit load successively on different points, the displacements are first calculated by FEM and then calculated with the functions. The different results are presented in Fig. 10. This figure shows the displacements of the different points for several load cases" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001911_iros.2007.4399209-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001911_iros.2007.4399209-Figure2-1.png", "caption": "Fig. 2. Model of HRP2 Humanoid Robot", "texts": [ " [4] proposed the multi-step planning method applicable to the rock-climbing humanoid robot. They [5] also proposed a method to guide the search by using the motion primitives. In this research, by adapting the primitive to the environment model, the PRM is performed within the period of time when the collision occurs. However, there is no research on motion planning of a humanoid robot walking on the flat/rough terrain with keeping the dynamical balance. Also, the feature of the proposed method is that we can plan all the DOF of the robot under constraint. Fig. 2 shows the model of a humanoid robot. Let p\u2217/\u03c6\u2217 be the three dimensional vectors of the position/orientation of the coordinate frame fixed to a link. The subscripts Fj, Hj, B and G denote the j-th foot, the j-th hand, the waist and the CoG, respectively. We assume that the 3D models of the robot and the environment are known. These models are used for collision checking. A configuration q \u2208 C of the humanoid robot is composed of the position/orientation of the waist (pB/\u03c6B) and all the joint angles (\u03b8)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000674_0954406042369080-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000674_0954406042369080-Figure1-1.png", "caption": "Fig. 1 Layout of experiment showing a ball projected at a force platform using an air cannon", "texts": [ " This model must be capable of calculating the force that acts on the ball and the motion of the ball centre of mass during impact and also the velocity of the ball after impact. The objectives of this study are: (a) to measure experimentally the dynamic response of a tennis ball for an impact with a rigid surface; (b) to derive a physically realistic model of a tennis ball impact on a rigid surface; (c) to assess the accuracy of the model by comparing experimentally obtained data with the data calculated by the model. The equipment used in this experiment is shown in Fig. 1. The balls were projected at a piezoelectric force platform using an air cannon. The speed gates were used to determine the inbound and rebound velocity of the ball, defined as VB and V 0 B respectively. These values were used to calculate the coefficient of restitution which is defined as the ratio of the rebound to inbound ball velocity. The ball was propelled at inbound speeds of between 14 and 30m/s (31.3 and 67.1mile/h), perpendicular to the surface of the force platform. The magnitude of the ball deformation that is associated with these inbound speeds is equivalent to that measured for typical ball\u2013racket impacts" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002348_1.2943299-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002348_1.2943299-Figure6-1.png", "caption": "Fig. 6 Two two-side JRS sheets", "texts": [ " For simplicity, the discussion in this section will be explained with the JRS of five-bar chains. The model is valid for any single-loop N-bar chains. JRS sheet. The JRS of a linkage branch is called a JRS sheet. It represents the configuration space of a linkage branch in the input domain. Thus, for single-loop chains, the JRS of a Class I linkage contains two sheets and that of a Class II linkage contains only one sheet. There is no motion continuity between sheets, i.e., a linkage cannot be transformed between configurations corresponding to points on different JRS sheets Fig. 6 . The formation of a branch or a JRS sheet is determined by the chain of the linkage and irrelevant to the input condition. Edge of a JRS sheet. The edge of a JRS sheet is the boundary curve of the JRS Fig. 6 . The edge of a sheet, which can be obtained by keeping the three passive joints collinear, may be a hole on a JRS sheet. Each point on the boundary curve of the JRS corresponds to an uncertainty singularity configuration of a linkage. The edge of a JRS sheet and, therefore, the shape of the JRS sheet depend on the choice of the input joints. Although the shape of a JRS sheet is affected by the choice of the input joints, the content of the JRS sheet i.e., the corresponding linkage configurations is not. Side of a JRS sheet. The edge of a JRS sheet separates the sheet into sides Fig. 6 . Each side of a JRS sheet represents the configuration space of a linkage sub-branch in the input domain. A point on one side of a JRS sheet corresponds to one and only one linkage configuration. Since a linkage can be programed within one side of a JRS sheet without reaching the boundary where an uncertainty singularity occurs, each side of a JRS sheet represents a uncertainty singularity-free configuration space. If a point is to reach the other side of the sheet, it has to reach a point on the edge, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001808_la8003728-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001808_la8003728-Figure1-1.png", "caption": "Figure 1. Schematic of the drifted motion of a water drop on a tilted (10\u00b0) hydrophobic surface subjected to a white noise vibration.", "texts": [ " However, in the presence of hysteresis, the vibration energy injected to the drop is used up in both viscous friction (mostly near the contact line) as well as in overcoming hysteresis, for which we assume the following simple form (see Appendix B): 1 2 K\u03c4)VA 2 +\u2206\u03c4VA (4) The drift velocity of the drop can then be obtained by combining eqs 3 and 4. When \u2206) 0, Vdrift ) g(sin \u03b8)\u03c4, which is independent of the power of noise K. On the other hand, for the case of finite hysteresis (\u2206 * 0 and K , \u22062\u03c4), the drift velocity as shown in equation 5 is a strong function of K, as was previously suggested by de Gennes.9 Vdrift ) g sin \u03b8\u03c4 1+\u22062\u03c4/K (5) Biased Motion of Drop under White Noise. We performed an experiment (Figure 1) in which small drops (10-5 kg) of water were placed on a hydrophobic silicon wafer at a tilt angle of 10\u00b0. In order to verify the effect of the power of vibration on drift velocity, bands of white noise of root-mean-square acceleration ranging from 12 to 250 m/s2 and \u03c4c \u223c 40 \u00b5s were used. A detailed description of the basic apparatus was published before in ref 13. Here we provide a brief description only. The silicon wafer silanized with decyltrichlorosilane was firmly attached to an aluminum plate that was connected to the stem of a mechanical oscillator (Pasco Scientific, model SF-9324)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000590_1.1844991-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000590_1.1844991-Figure4-1.png", "caption": "FIG. 4. Maxwell slip model: N massless elasto-slip elements in parallel approximate the hysteretic friction. Each element is characterized by a stiffness kis1\u00f8 i\u00f8Nd and a certain value for the force at which slip occurs.", "texts": [ " The advantage of this method is that the movement of the mass is described ana- 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: 130.160.4.77 On: Sat, 20 Dec 2014 10:50:57 lytically. A disadvantage of this method is that the size of the memory stacks has to be chosen in advance, resulting in possible overflow if too many open reversals are made.7 \u2022 Method based on the Maxwell slip model of the hysteresis force (see Fig. 4):6,23 This method models the hysteresis as N massless elasto-plastic elements in parallel. Each element is characterized by a stiffness kis1\u00f8 i\u00f8Nd and a certain value for the force at which slip occurs. The input x is common to all elements. This way only N reversal points can be remembered and the problem of overflow is solved at the expense of loss of accuracy. In subsequent analysis, it will be clear that this type of hysteretic behavior is characterized completely by the form of the virgin curve ysxd, given the reversal rules I and II" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000098_b:frac.0000021022.48417.a6-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000098_b:frac.0000021022.48417.a6-Figure3-1.png", "caption": "Figure 3. Two cavities of a surface crack: fully filled with lubricant and partially filled with lubricant.", "texts": [ " The volume of the cavity Ci cannot be smaller than the volume V i 0 of the lubricant trapped within it. This follows from the lubricant incompressibility. Thus, V i \u2265 V i 0 , where the cavity volume is V i \u2261 V i(x0) = \u222b Ci v(x)dx. If voids (lubricant tree volumes) occur in the cavity Ci , then the normal stress applied to the crack faces within this cavity is pi n = 0 (the lubricant vapor pressure in the void and the lubricant surface tension are neglected). Therefore, we arrive at the system of alternating equations and inequalities (see Figure 3) pi n = 0, V i > V i 0 ; pi n \u2264 0, V i = V i 0 ; i \u2208 I (x0) for y0 = \u2212l sin |\u03b1|. (9) The method for determining the lubricant volumes V i 0 is given below. It is necessary to mention that conditions (9) hold, as long as the neighboring cavities do not merge with Ci , i.e. Ci \u22c2 Cj = \u03c6, i = j ; i, j \u2208 I (x0). (10) Relations (6)\u2013(10) represent the necessary boundary and additional conditions from which one can obtain the unknown stresses pi n acting on the crack faces of the cavity Ci, i \u2208 I (x0), if the lubricant volume V i 0 is known" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003349_s12046-010-0005-1-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003349_s12046-010-0005-1-Figure1-1.png", "caption": "Figure 1. The coordinate systems used.", "texts": [ " Since, in presence of more than two bearings a rotor becomes statically indeterminate, the present work first computes the static equilibrium position of the rotor system using a nonlinear static finite element analysis. It then proposes a method to compute the quadratic nonlinear force from oil-film bearings. Using this information, the present work performs a dynamic analysis around the static equilibrium position using Newmark scheme specially adapted to tackle nonlinear forces. 2. Analysis 2.1 Coordinate systems The coordinate systems are shown in figure 1. For the three right-handed coordinate systems XYZ, X\u2032Y \u2032Z\u2032 and X\u0304Y\u0304 Z\u0304, the axes X, X\u2032 and X\u0304 are coincident. They are along the axis of the rotor. All the finite element equations are written with respect to the XYZ coordinate system. In the present work, the Z axis points vertically upward. At each oil-film bearing location coordinate systems X\u2032Y \u2032Z\u2032 and X\u0304Y\u0304 Z\u0304 are considered. In the X\u2032Y \u2032Z\u2032 coordinate system the axis Y \u2032 points towards the static radial displacement of the journal. In the X\u0304Y\u0304 Z\u0304 coordinate system the axis Z\u0304 points towards the static resultant force on the journal. The angles \u03b30, \u03b31, \u03b8 and \u03b10 between different axes are indicated in figure 1. 2.2 Modelling of coupling misalignment in a finite element model of a rotor The coupling is modelled using three constituents \u2014 coupling left half (CLH), coupling right half (CRH) and flexible part of coupling (CF). The CLH and CRH can be considered as short circular beams of large diameter so that they become almost rigid elements. The CF element is assumed to be composed of two shear springs and two rotational springs. The CF element has two nodes and just like beam elements two translation and two rotation degrees of freedom are attached with each node", " This model is, therefore, believed to be useful for development of techniques for detection of misalignment by post-processing of vibration response. Different methods based on frequency response functions can be explored for this purpose. In a rotor subjected on multiple oil-film bearings, it is also interesting to identify the bearing(s) operating in the sufficiently high nonlinear region. List of symbols p = p(X, \u03c6, t) Oil-film pressure X, Y, Z, Y \u2032, Z\u2032, Y\u0304 , Z\u0304, \u03c6, \u03b3, \u03b8 Coordinates and angles as described in figure 1 t Time variable \u03b7 Dynamic viscosity h = h(\u03c6, t) Oil film thickness when the journal and the bearing centers are not coincident h0(\u03c6) Oil film thickness when the journal and the bearing centers are coincident e, \u03b3 Polar coordinates of journal center with respect to Y\u0304 , Z\u0304 e0, \u03b30 Polar coordinates of journal center with respect to Y\u0304 , Z\u0304 for static equilibrium {D}, {D0}, { D} Vectors of total nodal displacements, static nodal displacements and dynamic nodal displacements respectively. {D} = {D0} + { D} {D\u03040} Inflated vector of static displacement degrees of freedom due to inclusion of Lagrange multipliers {F\u0304 } Inflated (due to inclusion of Lagrange multiplier) static force vector of weight and static part of bearing force d\u0304oj j th element of {D\u03040} dj , doj , dj Total displacement, static displacement and dynamic dis- placement at degree of freedom j respectively uiz\u0304 Dynamic displacement at node i and direction Z\u0304 FJy \u2032, FJz\u2032 Components of forces on journal in Y \u2032 and Z\u2032 directions F\u0302Jy \u2032, F\u0302Jz\u2032 Components of forces on journal in Y \u2032 and Z\u2032 directions obtained after appropriate data fitting F\u0302J y\u0304 , F\u0302J z\u0304 Components of forces on journal in Y\u0304 and Z\u0304 directions obtained after appropriate data fitting FJ0 Resultant static force on journal {Funb} The force vector with appropriate entries for unbalance {Fnl} The quadratic (in displacement) force obtained from Taylor series expansion of the bearing force Fnl iz\u0304 Nonlinear force at node i in the direction Z\u0304 {F st } Static force on the rotor due to weight of the discs {Fma} Misalignment force on the rotor {Fo} The constant part of bearing force obtained from Taylor series expansion of the bearing force {F lin} The linear (in displacement) force obtained from Taylor series expansion of the bearing force [K], [Kc] Stiffness of the rotor without and with the bearing stiffness respectively [K\u0304] Stiffness of the rotor where Lagrange multipliers are included as degrees of freedom [M] Mass matrix of the rotor [G] Gyroscopic matrix of the rotor [C] Damping matrix of the rotor" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000057_s0069-8040(08)70029-3-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000057_s0069-8040(08)70029-3-Figure5-1.png", "caption": "Fig. 5 . The wall jet. (a) Flow functions. (b) schematic streamlines. Here u = (15M/2ur3)\u201d\u2019 f\u2018(q); uz = 0.75 ( 4 0 M ~ / 3 r \u2019 ) \u2019 / ~ h ( q ) where q = ( 1 3 5 M / 3 2 v 3 r \u2019 ) I f 4 z . (From ref. 43.)", "texts": [ " In particular, there is the stagnation region (Region 111) and the wall-jet region (Region IV) which give rise to the wall-tube and wall-jet electrodes, respectively. The relative sizes of electrode and impinging jet are thus most important. References pp. 434-441 374 (a ) Wall-jet electrodes The laminar flow hydrodynamics for a radial wall-jet were first considered by Glauert [41] and subsequently by Scholtz and Trass [42] . A more complete evaluation for electrochemical purposes has recently appeared [ 4 3 ] . In Fig. 5 are shown the velocity profiles and schematic streamlines. Close to the wall we find ( 7 0 4 (70b) u, = cz2r -15 /4 u , = Czr-11\u20194 which satisfy the equation of continuity and where 3 114 c = with M the flux of exterior momentum flux given by Here, a is the jet diameter and h is a constant approximately equal to unity. If we rewrite the velocity components as CZ rm I u, = where m = l l j 4 , we seen that they are identical to the Levich approximations for the velocity components a t the RDE if we put m = - 1, and use C = 0", " 434-441 376 t P = (;I c - c, Y = - C , where q = 4 (5-rn) ac a 2 c to arrive a t the convective-diffusion equation *c=s with p = 1 for a disc electrode and p = 3 for a ring electrode. The value of q is 9/8 whereas, at the rotating disc electrode, it is 3. For the limiting current, identical boundary conditions [eqn. (19)] apply. We arrive directly at (75) \u2019 - 1.59k nFc a - 1 / 2 ~ 2 / 3 v - 5 / 1 2 V 3 / 4 ,.9/8 - 918 2 i 3 1L - m f ( p r p - 1 ) The ratio of limiting currents at ring and disc is once again ,62\u20193. In order for eqn. (75) to be correct, and as can be seen from the streamlines in Fig. 5, a minimum nozzle exit/electrode separation is necessary [44, 451. The wall-jet disc electrode is clearly not uniformly accessible (current density a r-5i4 ). Another important point is that i, depends on the threequarter power of the flow rate: it is more sensitive in this sense than rotating or tube/channel electrodes. Yamada and Matsuda [46] arrived at the same result for the limiting current by another method. They determined the constant k experimentally to be 0.86, a value which was confirmed by later observations [44]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.4-1.png", "caption": "Fig. 14.4 Patent drawing of excavator signed by W. S. Otis on 24 February 1839", "texts": [ "\u2022 The construction of a floating bucket-ladder dredger, called the Amphibious Digger, by Oliver Evans \u2013 the pioneer constructor of steam engines in America \u2013 for dredging the river and port of Philadelphia in the years 1800\u20131804.\u2022 The invention of the steel cable by Albert in Germany in 1834, the use of which greatly contributed to the development of cranes.\u2022 The building of the first single-bucket excavator, called the American Steam Excavator or the Yankee Geologist, by William Smith Otis in the USA in 1836 (Fig. 14.4). Otis\u2019s machine has been a symbol of construction mechanization to this day. The proof of its originality and technical excellence is the fact that it set the direction of the development of excavators for 140 years. Otis\u2019s first machines were used in the construction of the Baltimore\u2013Ohio railroad. Otis\u2019s excavator with a 1.15 m3-capacity bucket replaced the work of 80 diggers. Its high economic efficiency significantly contributed to the development of the excavator, grab-dredger, and crane building industry" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001443_09544054jem783-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001443_09544054jem783-Figure1-1.png", "caption": "Fig. 1 Meshed three-dimensional finite element model", "texts": [ " It may be concluded that the FE model can be used for reliable prediction and optimization of process parameters. The commercial FE software ANSYS was used to develop the model [24] which, in the present study, is used to investigate the effects of the deposition patterns on deformation and residual stresses of the substrate plate. A sequentially coupled, thermomechanical FE model was developed, in which the transient temperature distribution from thermal analysis was applied as body load to the structural model. The geometry of the model and the mesh are shown in Fig. 1, with the mesh refinement at the deposition and substrate interface shown in the enlarged view. The heat distribution at the instant of heat flux application is shown in Fig. 2. The basic geometry of Proc. IMechE Vol. 221 Part B: J. Engineering Manufacture JEM783 IMechE 2007 at University of Bath - The Library on June 25, 2015pib.sagepub.comDownloaded from the model comprises a rectangular 160 \u00b7 110 \u00b7 10mm substrate plate, bolted at the four corners to a 200 \u00b7 130 \u00b7 34mm support plate. The deposition in the original model [24] consists of nine successively deposited rows and in between rows the model is allowed to cool down for 180 s (i" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002031_icma.2008.4798802-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002031_icma.2008.4798802-Figure1-1.png", "caption": "Fig. 1 Reference frames", "texts": [ " The positive axis E points to the earth\u2019s core. The axis E and E are orthogonal. E points to anywhere. E is a right handed coordinate. The north-east-depth coordinate is adopted in this paper. The Onboard Coordinate OXYZ, which is not an inertial coordinate, is also defined. Axis OX is parallel to the datum plane of the vehicle and points to the head of AUV. Axis OY is parallel to the datum plane and points to starboard of AUV. Axis OZ points to the bottom of the vehicle. The relationship of these two reference frames is shown in Fig. 1. A B. Modeling Heading Control System for AUV In general, the motion of AUV that neglected of the rolling movement and the influence of coupling between horizontal plane and vertical plane, can be simplified as the two planes movement campaign, respectively. Consequently, it is useful and convenient for researching on the heading maintenance and control for AUV, as in [10]. The dynamic model can be obtained by Kinetic Research. According to fluid dynamics, Momentum Theorem and Moment of Momentum theorem, the dynamic equations for AUV heading control system are the equations (1)-(3): 3 2 2 2 21 1 1 2 2 2 4 2 3 2 2 ' 21 1 1 2 2 2 [ ] r r u uu vv rr vr r prop m u vr L X u L X u L X v L X r L X vr L u X X\u03b4 \u03b4 \u03c1 \u03c1 \u03c1 \u03c1 \u03c1 \u03c1 \u03b4 \u2212 \u2032 \u2032 \u2032= + + + \u2032 \u2032+ + + (1) 1 4 ' ' ' ' | |2 1 3 ' ' ' ' ' | |2 1 2 ' ' ' | |2 1 2 ' 2 2 [ ] [ | |] [ | |] [ | |] ( ) cos sin r r pq qr r r v ur wp vq v r uv vw v v r m v ur wp L Y r Y pq Y qr Y r r L Y v Y ur Y wp Y vq Y v r L Y uv Y vw Y v v W B L Y u\u03b4 \u03c1 \u03c1 \u03c1 \u03b8 \u03d5 \u03c1 \u03b4 + \u2212 = + + + + + + + + + + + + \u2212 + (2) 1 5 ' ' ' ' | |2 1 4 ' ' ' ' | |2 1 13 ' ' ' 3 ' 2 | |2 2 ( ) [ | | ] [ | | ] [ | | ] r z y x r pq r r qr v vq wq v r v v v vw r I r I I pq L N r N pq N r r N qr L N v N vq N wq N v r L N uv N v v N vw L N u\u03b4 \u03c1 \u03c1 \u03c1 \u03c1 \u03b4 + \u2212 = + + + + + + + + + + + (3) The attitudes of the vehicle in the inertial frame \u03c8 and \u03b8 are computed by (4): ( sin cos ) / cosq r\u03c8 \u03c8 \u03c8 \u03b8= + " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002904_1077546309104878-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002904_1077546309104878-Figure1-1.png", "caption": "Figure 1. Irregularity AB on road crossed by a mass load.", "texts": [ " The first is above which the touch between vehicle and road or bridge-deck surface is lost and therefore the vehicle flies like a launched missile, and the second is above which the flying vehicle lands beyond the end of the irregularity. A two degrees of freedom model is considered for the solution of the bridge, while the theoretical formulation is based on a continuum approach, which has been used in the literature to analyze such bridges. at SIMON FRASER LIBRARY on November 17, 2014jvc.sagepub.comDownloaded from 2. ANALYSIS 2.1. Irregularity on Road 2.1.1. Assumptions 1. We assume that a mass load, carried on a system of a spring of constant ko and of a damper of constant co, moves on the road with constant velocity x (Figure 1). At instant t 0, it meets the irregularity AB, the shape of which is given by the equation o x (1) 2. Because of the limited length of such an irregularity, compared to the bridge length, one can assume that the velocity x remains constant during crossing the irregularity. 3. Depending on the value x of the velocity, the mass load may either move in touch with the road (and the surface of the irregularity) or take off following an orbit like the one of a launched missile, and finally land on point (Figure 1). at SIMON FRASER LIBRARY on November 17, 2014jvc.sagepub.comDownloaded from 4. During flying the mass mo of the wheel will be vibrating free, being hung from the mass M, the orbit of which will be considered as the reference level for the above vibration of mo. 2.1.2. Mathematical Formulation The total force acting on the road is F M g z mo g o (2) Cutting at point G (Figure 1a), and taking into account the equilibrium of forces, we get M z ko z o co z o (3) Due to equation 3, equation 2 becomes F M mo g koz co z ko o co o mo o where o 0 for x xo or xo e x o o a for xo x xo e (4) On the other hand, because o x 0, equation 3 may be written as follows: z 2 o z 2 oz 2 o o where 2 o co M 2 o ko M (5) The solution of equation 5, with initial conditions z 0 z 0 0, is given by Duhamel\u2019s integral: z t 1 o t 0 e o t 2 o o sin o t d 2 o o o 2 o 2 o o e ot o cos ot o sin ot where o 2 o 2 o (6) From equation 6, one can easily find z t 2 o o o e ot sin ot (7) Finally, equation 4, using equations 6 and 7, becomes at SIMON FRASER LIBRARY on November 17, 2014jvc.sagepub.comDownloaded from F M mo g 2ko 2 o 2 o 2 o o 2 o o o e ot ko 2 o 2 o o cos ot o sin ot co sin ot (8) During the irregularity\u2019s crossing, the mass load takes the tangential speed o (Figure 1), that is given by the following relation: o x cos x 1 tan2 x 1 o2 (9) The irregularity\u2019s radius of curvature is Ro 1 o2 3 2 o , and thus the developed centrifugal acceleration will be c 2 o Ro 2 x 1 o2 o 1 o2 3 2 or, finally, c 2 x o 1 o2 5 2 (10) The developed centripetal force, which causes the deviation of the vehicle, is Fc M mo c or, finally, Fc M mo 2 x o 1 o2 5 2 (11) The restoring weight force FR is given by the following relation: FR F cos F 1 o2 1 2 (12) and, therefore, the condition for a safe crossing of the irregularity without loss of touch between wheel and road surface will be FR Fc (13) with F given in equation 8. From the above equation 13, we find the first critical speed 1cr, which determines the vehicle\u2019s behavior. On the other hand, from equation 13, for a known speed x 1cr, one can find the time tcr and the point of the vehicle\u2019s take off , through the relation x tcr x . At point x , the vehicle loses touch and thus follows an orbit like the one of a missile launched with initial speed o at initial angle (Figure 1). at SIMON FRASER LIBRARY on November 17, 2014jvc.sagepub.comDownloaded from The equations of the orbit, in parametric form, are the following: x x ot cos h o ot sin gt2 2 (14) A vehicle moving with a speed greater than a certain value (dependent on the irregularity\u2019s form) will land beyond the point B, the end point of the irregularity. This speed is an important parameter, that from now on we will call the \u201csecond critical speed\u201d 2cr. Eliminating the time t from equations 14, we get h o x x tan g x x 2 2 2 o cos2 (15) Putting x e and h 0, the above equation gives the second critical speed as follows: 2cr e x g 2 e x tan 2 o (16) The vehicle will arrive on point in time t , which is obtained by the solution of the following equations: h t ot sin gt2 2 for x 2cr ot sin gt2 2 0 for x 2cr (17) Thus, the point , on which the vehicle lands, is determined by the relations at SIMON FRASER LIBRARY on November 17, 2014jvc" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000844_iemdc.2005.195712-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000844_iemdc.2005.195712-Figure9-1.png", "caption": "Fig. 9. 12-slot, 10-pole permanent magnet brushless motor.", "texts": [ " Again, good agreement is achieved. -0.06 -0.05 -0.04 -0.03 -0.02 -0.01 0 0.01 0.02 0.03 0.04 0.05 1st 3rd 5th 7th 9th 11th 13th 15th Harmonic order C og gi ng T or qu e (N m ) 1-slot 2-slot (b) Harmonics -0.1 -0.08 -0.06 -0.04 -0.02 0 0.02 0.04 0.06 0.08 0.1 1st 2nd 3rd 4th 5th 6th 7th 8th 9th 10th 11th 12th 13th 14th 15th Harmonic order C og gi ng T or qu e (N m ) 1-slot 4-slot (b) Harmonics IV. 12-SLOT, 10-POLE MOTOR The synthesis technique has also been applied to a 12-slot, 10-pole motor, Fig. 9. Its outer stator diameter and axial length are 100mm and 50mm, respectively, while its airgap length and magnet thickness are 1mm and 3mm, respectively. The width of the stator slot openings is 2mm, whilst the magnets are again parallel magnetized, have a remanence of 1.2T and a relative recoil permeability of 1.05, a polearc/pole-pitch ratio of 1.0, Fig. 10 shows open-circuit field distributions for 1-slot and 12-slot motors, whilst Fig. 11 shows the associated cogging torque waveforms. Since the least common multiple, Nc, of the 1-slot, 10-pole motor is 10, the cogging torque periodicity is 36 degrees mechanical, while for the 12-slot 10-pole motor, Nc = 60, C = 2, and the cogging torque periodicity is only 6 degrees mechanical, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002288_j.mechmachtheory.2008.03.007-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002288_j.mechmachtheory.2008.03.007-Figure4-1.png", "caption": "Fig. 4. Generation with two fixed axes. Case with S \u00bc T\u00f00\u00de\u2013G\u00f00\u00de.", "texts": [ " The unit vectors xt1 and xg1 , and the points qt1 and qg1 chosen on the tool and gear axes, respectively, have components xt1 \u00bc \u00f00;0;1\u00de; qt1 \u00bc \u00f0 D0 cos c0; E0; B0 D0 sin c0\u00de; \u00f0105\u00de xg1 \u00bc \u00f0cos c0;0; sin c0\u00de; qg1 \u00bc \u00f00;0;0\u00de: \u00f0106\u00de The coordinates of the twists nt1 and ng1 are therefore nt1 \u00bc \u00f0E0;D0 cos c0;0;0;0;1\u00de; \u00f0107\u00de ng1 \u00bc \u00f00;0;0; cos c0;0; sin c0\u00de: \u00f0108\u00de The components n1 and n2 are simply n1 \u00bc ng1 and n2 \u00bc nt1 , since gsg\u00f00\u00de \u00bc I, and the transformation matrix gsg in (95) is given by the product of the two following matrices: en\u03021h1 \u00bc 1\u00fe \u00f0cos h1 1\u00de sin2 c0 sin c0 sin h1 sin c0 cos c0\u00f0cos h1 1\u00de 0 sin c0 sin h1 cos h1 cos c0 sin h1 0 sin c0 cos c0\u00f0cos h1 1\u00de cos c0 sin h1 1\u00fe \u00f0cos h1 1\u00de cos2 c0 0 0 0 0 1 266664 377775; \u00f0109\u00de en\u03022h2 \u00bc cos h2 sin h2 0 D0 cos c0\u00f0cos h2 1\u00de \u00fe E0 sin h2 sin h2 cos h2 0 E0\u00f0cos h2 1\u00de \u00fe D0 cos c0 sin h2 0 0 1 0 0 0 0 1 26664 37775: \u00f0110\u00de The proximal and distal Jacobians Jg gt and Jt gt , respectively, have the following expressions: Jg gt\u00f0h1\u00de \u00bc 0 E0 \u00fe E0\u00f0cos h1 1\u00de sin2 c0 \u00fe D0 cos c0 sin c0 sin h1 0 D0 cos c0 cos h1 E0 sin c0 sin h1 0 cos c0\u00f0E0\u00f0cos h1 1\u00de sin c0 \u00fe D0 cos c0 sin h1\u00de cos c0 cos c0\u00f0cos h1 1\u00de sin c0 0 cos c0 sin h1 sin c0 1\u00fe cos2 c0\u00f0cos h1 1\u00de 26666666664 37777777775 ; \u00f0111\u00de Jt gt\u00f0h2\u00de \u00bc sin c0\u00f0E0\u00f0cos h2 1\u00de \u00fe D0 cos c0 sin h2\u00de E0 sin c0\u00f0D0 cos c0\u00f0cos h2 1\u00de E0 sin h2\u00de D0 cos c0 cos c0\u00f0E0\u00f0cos h2 1\u00de D0 cos c0 sin h2\u00de 0 cos c0 cos h2 0 cos c0 sin h2 0 sin c0 1 2666666664 3777777775 : \u00f0112\u00de The computation of the remaining expressions such as, e.g., the enveloping family pg and the equation of meshing f \u00bc 0 follows immediately from definitions (97) and (98) once the joint functions h1\u00f0u\u00de and h2\u00f0u\u00de as well as the parametric equations pt\u00f0u; v\u00de of the tool are given. Explicit expressions are here omitted for brevity. 6.3. Example calculations with S \u00bc T\u00f00\u00de\u2013G\u00f00\u00de and Ot \u00bc qt1 , Og \u00bc qg1 In the case depicted in Fig. 4, the tool frame T coincides with the fixed frame S in the initial condition, that is T\u00f00\u00de \u00bc S, and their origins are taken on the reference point qt1 on the tool axis. The gear blank frame G is defined such that G\u00f00\u00de has the same orientation of S but its origin is displaced on the reference point qg1 on the gear axis. It is worth noting that the posture of G is identical to the case presented in the preceding section: therefore quantities expressed in G like n1, n2 and Jg gt will be necessarily identical to the ones previously calculated" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure9-1.png", "caption": "Fig. 9. Hyperboloid of one sheet.", "texts": [ " For the first scheme, because the spiral angle of the directrix is constant, the freedom of machine tool, machining the tooth flank of spiral bevel gear with finger-cutter, is minimum; for the second, because the directrix is the geodesic, it is convenient to manufacture with disc-cutter and favorable to improve efficiency and precision. The possibility of normal circular-arc gear for crossed-axis drive has been mentioned above, but there are also some difficulties to determine datum surface and the directrix. Here the example of hyperboloidal-type normal circular-arc gear for the orthogonal crossed-axis drive is discussed. The known datum surface R\u00f01\u00dep is a hyperboloid of one sheet generated by the rotation of straight generatrix Cm that is nonparallel to z1-axis. In the coordinate system {o1,x1y1z1} (shown in Fig. 9), the equation of straight generatrix Cm is represented as Cm: q \u00bc r1i1 u sin d1j1 \u00fe u cos d1k1 \u00f037\u00de Here, r1 and d1 are the shortest distance and the crossing angle between Cm and z1-axis, respectively. u is the parameter of the straight generatrix, i1, j1, k1 are unit vectors of the coordinate system. Substituting Eq. (37) into Eq. (1) and referring to Appendix A, the hyperboloid of one sheet can be represented as R\u00f01\u00dep : P\u00f01\u00de \u00bc B1\u00f0k1\u00deq\u00f0u\u00de \u00bc r1e\u00f0k1\u00de \u00fe u\u00bd sin d1e1\u00f0k1\u00de \u00fe cos d1k1 \u00f038\u00de Hence, normal vector of datum surface is represented as N\u00f01\u00de \u00bc P \u00f01\u00de k1 P\u00f01\u00deu \u00bc r1 cos d1e\u00f0k1\u00de u sin2 d1k1 u sin d1 cos d1e1\u00f0k1\u00de \u00f039\u00de In Fig. 9, the circle in the middle of the hyperboloid of one sheet is called as the gorge circle, whose radius is r1. According to the property of hyperboloid of one sheet, the surface can be formed by two symmetrical straight generatrixes, as Cm, C\u00f0 \u00dem . Comparing Fig. 9 with Fig. 2, related vectors in Eq. (3) can be translated as a \u00bc i1 b \u00bc j1 X\u00f01\u00de \u00bc k1 X\u00f02\u00de \u00bc I j1 X\u00f021\u00de \u00bc \u00f0I j1 \u00fe k1\u00de \u00f040\u00de Substituting Eqs. (37) and (40) into Eq. (3) and noticing the orthogonal condition (i.e. R = p/2), the conditional expression of the line of action of hyperboloid of one sheet can be obtained: r1 cos k1 \u00fe u sin d1 sin k1 \u00bc A sin2 d1 \u00f041\u00de This expression can be represented as k1 = k1(u). Substituting it into Eq. (38), the vector expression of line of action is described as P(0) = P(1)(k1(u))" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000082_2004-01-2830-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000082_2004-01-2830-Figure3-1.png", "caption": "Figure 3: The Dynamic Modules", "texts": [ " This is to enable the required stability at the lowest possible cost. The joints of the system press the beams together by means of bolts and pressure plates to keep the beams using friction, in specified positions. Different sizes of beams can be combined and moved to feasible perpendicular positions relative to each other as is illustrated in figure 2. The pressure plates belonging to the system are also used to attach the dynamic modules in feasible positions along the beams of a framework as is indicated in figure 3. Thus, varied functions of the framework are enabled by means of a minimum number of different construction modules. By means of calculations and experimental tests the applicability and stability of the developed construction kit for static framework has been verified. Further more, the use of air-cushions for the transport of the static framework on a floor is recommended. This method of enabling larger workspace of the robot relative to the product, has successfully been tested as part of the experimental work with the system. THE DYNAMIC MODULES This section presents the Dynamic Modules. Some parts of the sub-systems presented in this section have previously been presented in [1]. The Dynamic Modules were designed so that an operator could attach them to the static framework manually. Seven different Dynamic Modules were developed, see figure 3. In a normal case there would probably be only one or two Dynamic Module solutions chosen in an ART system. In this case, there was seven Dynamic Modules developed for being able to evaluate as many different solutions as was possible with respect to project time and budget. In general a Dynamic Module consists of a base plate and a top plate. In between there can be a parallel mechanical structure [11] or a serial linked structure. The Octapod, Tripod and Hexapod are examples of parallel mechanical solutions, and the Cradles are examples of the serial linked solution" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000839_j.euromechsol.2004.08.003-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000839_j.euromechsol.2004.08.003-Figure2-1.png", "caption": "Fig. 2. A cantilever beam attached to a rotating rigid cylinder.", "texts": [ " (25) After dividing by \u03c1Al, one obtains, for arbitrary \u03b4q\u0304 and x = l\u03c2 M\u22022q\u0304 \u2202\u03c42 + G\u2202q\u0304 \u2202\u03c4 + ( Kf + K0 + K1)q\u0304 \u2212 F = 0, (26) with matrices M = 1\u222b \u03a6T\u03a6 d\u03c2, G = 2 1\u222b \u03a6T\u03b7\u0303\u03a6 dx, F = \u2212 1\u222b \u03a6T [ \u00b5 + ( \u2202\u03b7\u0303 + \u03b7\u0303\u03b7\u0303 ) \u03c2e ] d\u03c2, (27) 0 0 0 \u2202\u03c4 1 K0 = 1\u222b 0 \u03a6T ( \u2202\u03b7\u0303 \u2202\u03c4 + \u03b7\u0303\u03b7\u0303 ) \u03a6 d\u03c2, (28) K1 = 1\u222b 0 ( \u2202\u03a6T 2 \u2202\u03c2 \u00b7 \u2202\u03a62 \u2202\u03c2 + \u2202\u03a6T 3 \u2202\u03c2 \u00b7 \u2202\u03a63 \u2202\u03c2 )[ 1 2 (\u03b72 2 + \u03b72 3)(1 \u2212 \u03c22) \u2212 \u00b51(1 \u2212 \u03c2) ] d\u03c2, (29) Kf = 1\u222b 0 ( \u03bb \u2202\u03a6T 1 \u2202\u03c2 \u00b7 \u2202\u03a61 \u2202\u03c2 + \u03b2 \u22022\u03a6T 2 \u2202\u03c22 \u00b7 \u22022\u03a62 \u2202\u03c22 + \u22022\u03a6T 3 \u2202\u03c22 \u00b7 \u22022\u03a63 \u2202\u03c22 ) d\u03c2, (30) \u03bb = Al2 Iy , \u03b2 = Iz Iy , (31) \u03b7 = [\u03b71 \u03b72 \u03b73]T = T \u03c9, \u00b5 = [\u00b51 \u00b52 \u00b53]T = T 2 l a0. (32) In this section, a cantilever beam attached to a rotating rigid cylinder is simulated in order to show the stiffening effect and to clarify the application range of the linear modeling method. For 0 \u03b1 \u03c0 , the time history of the angular velocity of the beam is given by \u03c91 = \u03c93 = 0, \u03c92 = \u03c9 = \u03b8\u0307 = \u2126t T \u2212 \u2126 2\u03c0 sin 2\u03c0t T (rad/s), 0 < t T, \u2126 (rad/s), t > T . (33) Let n be the unit normal vector as shown in Fig. 2, the base acceleration vector is written as a0 = \u03c92Rn \u2212 \u03c9\u0307Re3 = \u03c92R(\u2212 sin\u03b1e1 + cos \u03b1e2) \u2212 \u03c9\u0307Re3 = [\u2212\u03c92R sin\u03b1 \u03c92R cos \u03b1 \u2212 \u03c9\u0307R]T, (34) \u03c9 = [\u03c9 cos\u03b1 \u03c9 sin\u03b1 0]T, (35) where e2 = [0 1 0]T, e3 = [0 0 1]T, and R is the radius of the rigid cylinder. The non-dimensional variables are given by \u03b7 = [\u03b7 cos \u03b1 \u03b7 sin\u03b1 0]T, \u00b5 = [ \u2212\u03b72\u03b5 sin\u03b1 \u03b72\u03b5 cos \u03b1 \u2212 \u2202\u03b7 \u2202\u03c4 \u03b5 ]T , (36) where \u03b7 = \u03c9T, \u03b5 = R l . (37) Neglecting the axial stretch and choosing the first mode for the lateral deformation in y and z direction, respectively, the modal matrices are given by \u03a61(\u03c2) = [0 0], \u03a62(\u03c2) = [ \u03c7(\u03c2) 0 ] , \u03a63(\u03c2) = [ 0 \u03c7(\u03c2) ] , (38) \u03c7(\u03c2) = cos(1" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.79-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.79-1.png", "caption": "Fig. 14.79 Portable two-mast climbing platform with rack-andpinion lifting gear", "texts": [ " The combination of the mast-climbing platform\u2019s sectional structure and the system of protractible struts makes it possible to obtain work access on walls of any shape (straight, curved, and with slants and bevels) and architectonic form (balcony, Part B 1 4 .5 loggia, niche, bay). The whole mast-climbing platform is fenced in with railings to protect persons working on it from falling out. It is also possible to combine two single-mast climbing platforms to form one platform (up to 40 m long) climbing two masts (Fig. 14.79 and Table 14.10). In many cases, mast-climbing platforms may replace stationary construction-assembly scaffolds. Similarly to material hoists and person and material hoists with a rack-and-pinion drive, mast-climbing platforms are equipped with the following safety devices: \u2022 An emergency lowering system\u2022 A braking device\u2022 A safety device preventing the driving gear wheel from disengaging from the mast\u2019s gear rack\u2022 Electric-shock protection\u2022 Overload protection for the electric motors\u2022 Work-platform-slanting cutouts (in two-mast climbing platforms)\u2022 Work platform terminal position cutouts\u2022 Sensors signalling a platform loading which may result in overturning of the mast-climbing platform or its damage Mobile Elevating Work Platforms Mobile elevating work platforms have a similar range of applications (elevating persons and equipment) as the portable mast-climbing platforms described above, except that their use in one work place is short" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001710_aim.2007.4412505-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001710_aim.2007.4412505-Figure5-1.png", "caption": "Fig. 5. Movement of a circular wavefront, as a problem of boundary conditions", "texts": [ " 1 F = |\u2207T | For multiple dimensions, the same concept is valid because the gradient is orthogonal to the level sets of the arrival function T (x). In this way, the movement of the front can be characterized as the solution of a boundary conditions problem. The speed F depends only on the position, then the equation 1 F = |\u2207T | or the Eikonal equation: |\u2207T |F = 1. (1) As a simple example we define a circular front \u03b3t = {(x, y)/T (x, y) = t} for two dimensions that advance with unitary speed (see fig.5) . The evolution of the value of the arrival function T (\u03b8) can be seen as the time increases (i.e. T = 0, T = 1, T = 2, ...) and the arrival function comes to points of the plane in more external regions of the surface as can be seen in fig. 7. The boundary condition is that the value of the wave front is zero in the initial curve. The direct use of the Fast Marching method does not guarantee a smooth and safe trajectory. Due to the way the front wave is propagated the shortest geometrical path is determined" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001583_1077546307078829-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001583_1077546307078829-Figure2-1.png", "caption": "Figure 2. Sound radiation from an asperity modeled as a cantilever beam (clamped to the gear tooth surface). (a) Schematic for sound radiation calculation. (b) Baffle configuration.", "texts": [ " Then the modal superposition method is adopted to describe the temporal behavior of asperity structures and the resulting sound radiation. It is assumed that the asperities are acoustically baffled along their axes (normal to the gear tooth surface) for the prediction of far-field sound. Further, only the free field is considered here and thus sound interactions among the asperity elements are not included in our model. at Bibliothekssystem der Universitaet Giessen on May 31, 2015jvc.sagepub.comDownloaded from Consider an asperity beam on the gear tooth surface, as shown in Figure 2(a) along with the gear contact coordinates which are denoted by the Line of Action (LOA) and Off-Line of Action (OLOA). The cantilever beam of Figure 2(a), describing an asperity, is fixed to the gear tooth surface. It is assumed to be acoustically baffled, along the LOA that is normal to the gear tooth surface as illustrated in Figure 2(b). This baffle assumption is made to develop a tractable problem though it somewhat deviates from reality. The acoustic wave equation with sources distributed over the cantilever beam is as follows where and are the velocity potential and Laplace operator respectively: 2 x t t2 2 x t x 0 t0 bdx0 (1) Here, x 0 t0 is the velocity distribution over the asperity beam with the width b and acoustic field is considered in the half space where y 0. Refer to Appendix A for the identification of symbols" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002188_bfb0110288-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002188_bfb0110288-Figure1-1.png", "caption": "Fig . 1. The reduced (1:80) size US Navy crane in the authors ' lab.", "texts": [ " The solution of the o p e n - l o o p motion planning problem is presented in Section 4 based on the flatness property of the model. Asymptotic global stability of equilibria in c l o s e d - l o o p using output feedback regulators of proportional-derivative type is studied in Section 5. Simulation results of an extension of the same controller with open-loop trajectory planning for tracking are presented in Section 6. 2 G e n e r a l d e s c r i p t i o n o f t h e e x p e r i m e n t a l s e t u p The reduced scale (1:80 size) model 1 of the US-Navy crane is depicted in Figure 1. Four DC motors (three of them winching ropes) are mounted on the structure allowing to manipulate the load in a three dimensional workspace. The control objective is to move the load swiftly from an initial position to a desired final position without sway and avoiding obstacles. Since the accelerations of the motors tend to create oscillations of the load, simultaneously fast and swayless displacements are hard to realize. The reduced size model comprises: 9 A load (maximal nominal mass: 800g) 9 A mobile pulley guiding the rope which hoists the load" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000319_095440605x8478-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000319_095440605x8478-Figure1-1.png", "caption": "Fig. 1 A force and a moment applied to a free end of a flexure hinge", "texts": [ " Then, a force and a moment applied to a point on a rigid body are derived. Next, an equation of motion for a flexure hinge-based compliant mechanism is proposed. Subsequently, it is shown that the equivalent stiffness matrix of the proposed equation of motion differs from that of the conventional equation of motion. Finally, by means of a comparison with FEM analyses, it is shown that the proposed equation of motion describes the behaviours of rigid bodies better than the conventional equation of motion. Figure 1 shows a force and a moment applied to a free end of a flexure hinge. In this figure, the flexure hinge coordinates are xij, yij, zij and the reference coordinates are X, Y, Z. The flexure hinge coordinates are rotated with respect to the reference coordinate by the angles g, b, a, expressed by the Z, Y, X Euler angles, for which g, b, a are rotations of the Z, Y and X axes respectively. If a force F h and a moment T h are applied to the free end of the flexure hinge, the displacement of the free end of the flexure hinge is a translation r h and a rotation u h, and the stiffness matrix of the flexure hinge is K ij h" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002213_13506501jet575-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002213_13506501jet575-Figure1-1.png", "caption": "Fig. 1 Geometrical configuration of the journal bearing", "texts": [ " Among the existing lowRe \u03ba\u2013\u03b5 turbulence models, the model developed by Abe et al. [20], the AKN low-Re \u03ba\u2013\u03b5 turbulence model, is regarded as one of the most reliable ones employed in the present analysis. The model can reproduce the near-wall limiting behaviour and provides accurate predictions for the boundary layer of turbulent flows with favourable or adverse pressure gradients such as for the lubricant flow in journal bearings [21]. Geometrical configuration of the related journal bearing is shown in Fig. 1. The governing equations for lubricant flow which are considered to be twodimensional, steady, incompressible, and turbulent are the equation of continuity, the time-averaged Navier\u2013Stokes equation, the equations of turbulent energy \u03ba and its dissipation rate \u03b5, the energy equation for the flow field, and Laplace\u2019s equation for the thermal conduction in the bearing. The non-dimensional forms of these equations in the Cartesian coordinate system (x, y) (Fig. 1) may be written as follows: (a) continuity equation \u2202u \u2217 \u2202x\u2217 + \u2202v \u2217 \u2202y\u2217 = 0 (1) (b) momentum equation in the x-direction \u2202 \u2202x\u2217 [ u\u0304\u22172 \u2212 ( 1 + \u00b5\u2217 t Re ) \u2202u\u0304\u2217 \u2202x\u2217 ] + \u2202 \u2202y\u2217 [ u\u0304\u2217v\u0304\u2217 \u2212 ( 1 + \u00b5\u2217 t Re ) \u2202u\u0304\u2217 \u2202y\u2217 ] = \u2212\u2202P\u0304\u2217 \u2202x\u2217 + \u2202 \u2202x\u2217 ( \u00b5\u2217 t Re \u2202u\u0304\u2217 \u2202x\u2217 ) + \u2202 \u2202y\u2217 ( \u00b5\u2217 t Re \u2202 v\u0304\u2217 \u2202x\u2217 ) (2) (c) momentum equation in the y-direction \u2202 \u2202x\u2217 [ u\u0304\u2217v\u0304\u2217 \u2212 ( 1 + \u00b5\u2217 t Re ) \u2202 v\u0304\u2217 \u2202x\u2217 ] + \u2202 \u2202y\u2217 [ v\u0304\u22172 \u2212 ( 1 + \u00b5\u2217 t Re ) \u2202 v\u0304\u2217 \u2202y\u2217 ] = \u2212\u2202P\u0304\u2217 \u2202y\u2217 + \u2202 \u2202x\u2217 ( \u00b5\u2217 t Re \u2202u\u0304\u2217 \u2202y\u2217 ) + \u2202 \u2202y\u2217 ( \u00b5\u2217 t Re \u2202 v\u0304\u2217 \u2202y\u2217 ) (3) (d) equation of turbulent kinetic energy \u2202 \u2202x\u2217 [ u\u0304\u2217k\u2217 \u2212 1 Re ( 1 + \u00b5\u2217 t \u03c3k ) \u2202k\u2217 \u2202x\u2217 ] + \u2202 \u2202y\u2217 [ v\u0304\u2217k\u2217 \u2212 1 Re ( 1 + \u00b5\u2217 t \u03c3k ) \u2202k\u2217 \u2202y\u2217 ] = 1 Re \u00b5\u2217 t \u03d5 \u2217 1 \u2212 \u03b5\u2217 (4) JET575 \u00a9 IMechE 2009 Proc" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001679_gt2008-50305-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001679_gt2008-50305-Figure7-1.png", "caption": "Figure 7. Enlarged view of turbine rotor to show relevant features: pilot diameter (A), blade profile (B) and hubline.", "texts": [ "org/ on 01/27/2016 T manufacturing process and the end user, but cost prohibitive due to a high scrap rate of incoming parts. The root cause had to be identified and corrected now that the problem was quarantined. It was theorized that the casting was not well centered in the workholding tool during the machining of the critical datum features used in subsequent machining steps, such as blade profile grinding and location of the pilot fit. This would create an effect where the cast aerodynamic features, as seen in Figure 7, were not symmetric about the rotating center of the part. As the casting, namely the massive hub, was eccentric, the initial unbalance of the worst components was quite high, generally proportional to the centering error. But it was still possible, even with badly unbalanced parts, to measure the machined dimensions to be in print relative to the machining datums. This was because the datums themselves were mis-aligned with the casting center of mass. It was clear at this point that there was a deficiency in the component definition that allowed this to occur" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000975_6.2004-5308-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000975_6.2004-5308-Figure1-1.png", "caption": "Figure 1. A N-body Hub-And-Spoke Tethered Formation.", "texts": [ " In the present paper, the dynamics of spinning multi-tethered satellite formations in low Earth orbit, in a hub-and-spoke configuration, having a spinning plane which is either the orbital plane or a plane orthogonal to the orbital plane, are examined. The two types of formations under consideration are hereby named, for ease of discussion, hub-andspoke configuration and closed-hub-and-spoke configuration. The hub-and-spoke configuration is a system consisting of a central (or parent) body of mass mP in the hub and N tethers of length li connecting N bodies of mass mi to the hub (Fig. 1). The closed-hub-and-spoke configuration is the same system but to which are added N external tethers connecting the N peripheral bodies in pairs in a ring-type setting (Fig. 2 of 23 American Institute of Aeronautics and Astronautics 2) , i.e m1 is joined to m2, mN is joined to m1, etc. The tethers are considered straight but elastic and are subject to structural damping, and their mass is assumed to be negligible. The dimensions of the end-bodies are much smaller than the length of the tethers so that the former can be approximated as point masses" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002825_ijmpt.2010.031891-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002825_ijmpt.2010.031891-Figure2-1.png", "caption": "Figure 2 Contact model for dry contact of two rough surfaces", "texts": [ " We used two random variables for each surface. That are the relative height of the asperities \u03be1,2 and the relative radius of the asperities \u03c11,2 (we model the asperities as spherical segments). The contact probability can be obtained with (Kuhn, 2001) 2 1 2 1 2 1 2 1 2 1 1 2 20 0 ( , ) ( ) ( ) d d ( ) d ( ) d .g g k k z z u ur u u a jF z u f f f f j \u03be \u03be \u03be \u03be \u03be \u03c1 \u03c1 \u03c1 \u03c1 \u2212 = = \u00f1 \u00f1 \u00f1 \u00f1 (3) Here ja is the sum of the number of asperities and jr is the number of contacts. For that simulated dry contact we obtain the Figure 2. The investigations of Holweger (1998) deliver the grease topography as shown in Figure 3. He used the IR-microscopy to observe the density distribution within the grease film. Now we add the lubricating grease to the contact model and in the first step, we only look into the minimum gap within the contact area (which are the areas where we find a direct contact in the dry contact model). In addition to this, we select some density areas to observe the shear process. To get an idea we developed Figure 4 (Kuhn, 2000)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000718_1.1829068-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000718_1.1829068-Figure3-1.png", "caption": "Fig. 3 Ratcheting indexer2", "texts": [ " Treating linkage E-H-J as the input and considering gear B the output would result in a kinematically equivalent submechanism ~C-B! to links 2 and 4 of the wobble gear having 48 and 46 teeth. The motion of member J at ball joint K is analogous to the cam pair of the wobble gear. It will be shown later that the wobble gear, the mower-bar actuator just described, and a host of other mechanisms may all be described as a mechanism family, or in other words as either \u2018\u2018similar\u2019\u2019 or \u2018\u2018equivalent\u2019\u2019 mechanisms using the terminology outlined in the next section. Figure 3 is a diagram of a mechanism which uses a wobbling motion to create discrete, indexed displacement @1#. Connecting rods 6 and 68 are connected to the wobble plate by spherical joints and to a pair of concentric ratcheting indexers by revolute joints, effectively creating a pair of four-bar mechanisms. The wobble plate is kept from rotating by pins A and A8, which ride in slots in the housing. Thus one rotation of the input shaft 1 causes link 2 to wobble, the four-bar is actuated, and the ratchet is indexed", " Although these two mechanisms are very interesting applications of wobbling motion, probably the most widespread implementation of such motion is found in piston-cylinder or slidercrank mechanisms which, like the mowing machine of Fig. 2, are of the rotational-to-linear motion conversion type. Wobble-plate compressors and engines have an extensive history @3#. Figure 4 shows one of many examples of wobbling slider-crank mechanisms given by Artobolevsky @1#. Note the similarity of the piston rods here to members 6 and 68 in Fig. 3. Wobble plate 2 is attached to connecting rods 4 and 48 at spherical joints B and B8. Since rollers b and b8 are guided in slots in the housing, constrain- 005 by ASME MARCH 2005, Vol. 127 \u00d5 269 shx?url=/data/journals/jmdedb/27802/ on 03/07/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F ing the wobble plate from rotating out of the plane of the paper, revolution of the input shaft 1 results in actuation of the pistons 5 and 58. While the slider-crank devices are among the most widely used of the wobbling mechanisms, there exist more than a few cases of wobbling speed reducers like the wobble gear of Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002592_1.2772634-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002592_1.2772634-Figure1-1.png", "caption": "Fig. 1 An example of a structure with friction contact interfaces: \u201ea\u2026 a bladed-disk assembly, \u201eb\u2026 friction contact between blade shrouds, and \u201ec\u2026 friction underplatform dampers", "texts": [ " This vector depends on displacements at the nterface nodes, q t , and on a vector of parameters, b, which haracterize all friction contact interfaces included in the strucure. The contact interface parameters can be gap, friction coeffiient, contact stiffness coefficients, normal stresses which are due o the action of the centrifugal forces and temperature fields, and thers. In modern industrial problems large-scale finite element odels containing 105\u2013107 DOFs are routinely used. An example f a practical bladed disk with friction contact interfaces is shown n Fig. 1. 22503-2 / Vol. 130, MARCH 2008 om: http://gasturbinespower.asmedigitalcollection.asme.org/ on 01/29/201 values, and by other statistical characteristics. In other cases only estimates of ranges within which the parameters can take values can be provided. In all these cases, there is a need to have an efficient method allowing calculation of the stochastic characteristics and ranges of uncertainty in the forced response as a function of characteristics of design parameter uncertainty see Fig. 2 " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003320_gt2010-22058-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003320_gt2010-22058-Figure1-1.png", "caption": "FIG. 1 SCHEMATIC OF FORWARD WHIRL", "texts": [ " The calculated stiffness coefficients were in agreement with the measurements for different preswirl values. The effect of upstream and downstream cavities on the predictions were studied. In this work, two experimental methods as well as CFD modelling to obtain stiffness and damping coefficients of labyrinth gas seals are analyzed. The aim is to explain the observed difference between static and dynamic experimental methods and to show accuracy of CFD predictions. Shaft precession motion within the seal clearance produces the reaction forces in the seal. Figure 1 shows a forward whirl when the shaft whirls around the seal center with the amplitude e and the whirl speed \u03a9, and simultaneously rotates on its axis with the rotational speed \u03c9. Under the assumptions of circular whirl orbit and linear dependence of seal reactions on shaft eccentricity, the rotordynamic model of the seal becomes [16]: {Fr /e=\u2212K\u2212c F t /e=k\u2212C (1) 2 Copyright \u00a9 2010 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/19/2016 Terms of Use: http://www.asme" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000645_j.dental.2005.04.025-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000645_j.dental.2005.04.025-Figure1-1.png", "caption": "Figure 1 Three types of finish line curvatures\u2019 abutments (distal).", "texts": [ "6 wt% and gold content of 98.4 wt%. These were alloyed in an argon arcmelting furnace (TAM4-s, Tachibana Riko, Sendai, Japan) to manufacture the Au-1.6 wt% Ti alloy. KIK, a commercially available gold alloy for metal ceramic crowns (Au 85.5%, Ag 0.5%, Pt 4.0%, Pd 8.0%ite ISHIFUKU METAL INDUSTRY CO. LTD, Tokyo, Japan) was used as an appropriate control as it is used extensively for metal ceramic crown production in Japan. Three types of abutment finish line curvatures (1, 3, and 5 mm-curvature) were prepared (Fig. 1). A maxillary right central incisor Ivorine tooth (A1-500, NISSIN DENTAL CORPORATION, Tokyo, Japan) was prepared by an experienced prosthodontist with a circular 1 mm 90-degree shoulder finish line, a uniform 1.5 mm two-plane facial reduction, a 2 mm incisal reduction and 1 mm remaining axial reduction with 6-degree taper. The prepared toothwith a 1 mm vertical distance between proximal margin level and buccolingual margin level was defined as the 1 mmcurvature abutment, namely the abutment with a 1 mm-curvature finish line. The prepared tooth was reproduced in a gold-silver-palladium alloy (12% CASTWELL M.C., GC Corporation, Tokyo, Japan), and 1 mm-curvature metal abutment was made (Fig. 1, left). The additional 1 mm-curvature metal abutments were produced, then 3, 5 mm-curvature metal abutments were respectively fabricated by further apical reducing buccolingualmargins of these 1 mm-curvature metal abutments and keeping the proximal margins untouched (Fig. 1, middle, right). The dimensions of 5 mm-curvature metal abutments are shown in Fig. 2. Wax patterns for metal ceramic crowns were fabricated directly on metal abutments with blue inlay wax (CROWN WAX, GC Corporation, Tokyo, Japan). Thirty wax patterns were fabricated (ten for each type of abutment). Wax patterns were examined under a stereoscopic zoom microscope (SMZ-1, Nikon Corporation, Tokyo, Japan) to ensure that there was no gap between wax pattern and metal abutment margins. After storage at room temperature for 24 h, the wax patterns were vacuum invested in phosphate-bonded investment (CERAVEST G, GC Corporation, Tokyo, Japan) according to the manufacturer\u2019s recommendation for the liquid: power ratio" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001838_0094-114x(75)90072-5-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001838_0094-114x(75)90072-5-Figure1-1.png", "caption": "Figure 1. Elliptic wheels derived from the antiparallelogram.", "texts": [ " For finishing purposes for instance there might be used a grinding wheel with a plane face normal to its axis; if it is moved in such a way that a line parallel to this axis is constrained to roll on the base ellipse g, the flat side of the tool would grind the involute face of a tooth. A similar concept would serve for the motion of cutters or hobs generating the tooth spaces in the gear blank. Corresponding developments could be performed for elliptic bevel gears by means of operations on the sphere analogous to those used here in the plane. Elliptic gear wheels are based upon the motion of an antiparallelogram E,F2E2F, and its pole curves which consist of two equal ellipses pl and p2 with foci E,, FI and E2, F2 respectively (Fig. 1). Any two smaller equal ellipses gl and g2, confocal with the rolling ellipses pl and p: and rigidly connected with them, have the property that a crossed transmission belt slung around them does not disturb the original motion (Fig. 5). This fact gaves the possibility for an apparently new gear system for elliptic wheels which represents a natural generalization of the well-known involute system for circular wheels: The tooth profiles are throughout involutes of the \"base\" ellipses g, and g~ (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000367_pesc.2005.1581876-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000367_pesc.2005.1581876-Figure3-1.png", "caption": "Fig. 3. Back-EMF waveforms of the simulated machine.", "texts": [ " vdq = Rsidq + (Ls \u2212Ms) d dt idq+ j (Ls \u2212Ms) \u03c9ridq + j \u221a 3 2 \u03a6m\u03c9r (13) Tel = zp \u221a 3 2 \u03a6miq (14) In this well known case the electromagnetic torque is directly proportional to the quadracture stator current. This is a characteristic of the sinusoidal machines that provide good dynamic performance to these machines. C. Vectorial model for non-sinusoidal SM-PMSMs using the extended DQ transformation Non-sinusoidal SM-PMSMs present back EMF waveforms varying from ideally trapezoidal to very complex behavior or waveform depending on the magnets, stator coils distribution and other constructive factors [8]. As an example, the figure 3 shows a typical back EMF waveform of a real SM-PMSM. The classical dq transformation procedure applied to the nonsinusoidal back EMF machines shows that the electromagnetic torque will present high ripple content. In the sequence, the extended dq transformation, that renders possible to express the electromagnetic torque by a simple equation will be developed. By using this equation it is possible to drive the SM-PMSM machine in the open loop with very smooth torque ripple. After the application of \u03b1\u03b20 transformation (4), the proposed dq transformation given by (15) is used, where ax is a variable dq coefficient as indicated in the figure 4", " vdx = ( Rskix + (Ls \u2212Ms)\u03c9r ( 1 ax dax d\u03b8r kix\u2212( 1 + d\u03b8x d\u03b8r ))) iqxref (31) vqx = ( Rs + (Ls \u2212Ms) \u03c9r ( 1 ax dax d\u03b8r + ( 1 + d\u03b8x d\u03b8r ) kix )) iqxref + \u221a 3 2 \u03a6m 1 a2 x \u03c9r (32) Commercially, the association between a SM-PMSM and its electrical converter is named \u201cbrushless-DC machine\u201d. A brushless DC motor drive system was simulated in three different conditions presented below. The machine used in the system is a brushless DC machine of Siemens, model 1FT5, whose back EMF waveforms are given in the figure 3. Considering that back-EMF waveform, the angle \u03b8x and the coefficient ax, respectively given in (21) and (23) are obtained as depicted in the figures 5 and 6. Rotor angle 60 electrical degrees/div 0 thetax (rd) 0.04 \u22120.04 Rotor angle Fig. 6. The solution for the dqx coefficient ax. The first condition of machine simulation is the machine fed by a conventional six-step converter. In this case, the motor is started from rest with no load using the six-step converter commanded conventionally by the hall sensors signals" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.95-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.95-1.png", "caption": "Fig. 14.95 Scaffold crane", "texts": [ "91 Diagram representing hoisting capacity of crane with maximum capacity of 130 t. Note: The characteristic is determined for the main jib and the auxiliary jib. The numbers above the curves specify the allowable hoisting capacity for a given jib and hoisting height A range of portable machines, based on winches and other accessories, for handling materials and transferring light equipment on construction sites has been developed. This machinery includes: \u2022 Scaffold cranes mounted on scaffolds (Fig. 14.92 items 5 and 6, and Fig. 14.95)\u2022 Portable cranes (Fig. 14.92 item 4, and Figs. 14.96, 14.98) fixed to steel supports installed between floors, in window openings or on the roof (Fig. 14.97, basic parameters are shown in Table 14.17)\u2022 Gantries mounted on the roof (Fig. 14.92 item 3), in an opening in the building\u2019s elevation (Fig. 14.92 item 2) or on a scaffold (Fig. 14.89 item 1) Part B 1 4 .6 The main component of the above machines is a universal winch that can work in tandem with various accessories. Figure 14.92 shows the use of scaffold cranes, portable cranes, and small-capacity gantries during building erection" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000613_s1526-6125(06)70096-2-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000613_s1526-6125(06)70096-2-Figure1-1.png", "caption": "Figure 1 Schematic of LDM for Thin Wall (Vasinonta, Beuth, and Griffith 2001)", "texts": [ " In this study, to achieve the thin wall with thickness less than 1.0 mm, the minimum laser power of a focused Gaussian beam for manufacturing thin-wall metallic parts and thickness of the formed wall are investigated, and a CO2 laser with low power and small light spot on the interacting area is applied in single-pass cladding. A thin metallic wall is usually fabricated with laser scanning along a single row, and the wall thickness, , is equal to the width of the molten pool created in laser cladding. As shown in Figure 1, the laser beam moves in the positive x-direction at a velocity, v. The origin is on the center line of the laser beam. In this study, the following assumptions can be made: 1. The laser beam diameter is defined as the diameter in which the Gaussian power density is reduced from the peak value by a factor of e2. 2. All thermophysical properties are considered to be temperature independent. 3. Because the temperature of the melt pool is usually below 2000\u00b0C (for ferrous alloys), the radiant heat of its surface and convective flow of heat from its surface to the surroundings are neglected" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001710_aim.2007.4412505-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001710_aim.2007.4412505-Figure4-1.png", "caption": "Fig. 4. Propagation of a wave and the corresponding minimum time path when there are two media of different slowness (diffraction) index. (a), the same with an vertical gradient (b), Wavefront propagating with velocity F (c), and Formulation of the arrival function T (x), for an unidimensional wavefront.(d).", "texts": [ " In the proposed method, the first potential (the repulsive one) is used as the refraction index of the wave emitted from the goal point. This way a unique field is obtained and its associated vector field is attractive to the goal point and repulsive from the obstacles. This method inherits the good 1-4244-1264-1/07/$25.00 \u00a92007 IEEE 2 properties of the electromagnetic field. III. INTUITIVE INTRODUCTION OF THE EIKONAL EQUATION AND THE FAST MARCHING PLANNING METHOD Intuitively, Fast Marching Method gives the propagation of a front wave in an inhomogeneous media as shown in fig 4a and 1b. Let us imagine that the curve or surface moves in its normal direction with a known speed F (see fig. 4c). The objective would be to follow the movement of the interface while this one evolves. A large part of the challenge, in the problems modelled as fronts in evolution, consists of defining a suitable speed, that faithfully represents the physical system. A way to characterize the position of a front in expansion is to compute the time of arrival T, in which the front reaches each point of the underlying mathematical space of the interface. It is evident that for one dimension (see fig. 4d) the equation for the arrival function T can be obtained in an easy way, simply considering the fact that the distance x is the product of the speed F by the time T : x = F \u2219 T . The spatial derivative of the solution function becomes the gradient: 1 = F dT dx and therefore we have that the magnitude of the gradient of the arrival function T (\u03b8) is inversely proportional to the speed. 1 F = |\u2207T | For multiple dimensions, the same concept is valid because the gradient is orthogonal to the level sets of the arrival function T (x)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000602_s1526-6125(05)70092-x-Figure11-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000602_s1526-6125(05)70092-x-Figure11-1.png", "caption": "Figure 11 Miniature Pump Prototype", "texts": [ " Preliminary work in this direction (as seen in Figure 10) has produced evidence that the modest temperature rise in combination with intense strain rates in the metal produces a concentration of lattice vacancies that is several orders of magnitude higher than the equilibrium one. This mechanism promotes bonding of the foil materials by interdiffusion of vacancies. Further, quantitative description of these metallurgical aspects of ultrasonic welding is currently under investigation. The miniature pump in Figure 11 is one example of a functional and testable model. This 3-D pump was constructed out of 60 layers of aluminum foils (127 \u00b5m/layer) in less than 20 minutes. It has a flexible top diaphragm, solid bottom part, and internal chamber cavity, with an inner hole on the bottom and an outlet hole on the side. Amon, C.H.; Beuth, J.L.; Weiss, L.E.; Merz, R.; and Prinz, F.B. (1998). \u201cShape deposition manufacturing with microcasting: processing, thermal and mechanical issues.\u201d Journal of Mfg. Science and Engg" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000395_fie.2003.1265943-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000395_fie.2003.1265943-Figure2-1.png", "caption": "FIGURE 2. S O L l D M O D E L ~ F T H E A U T O M D A D H E S l V E SYSTEM", "texts": [ " The students performed a complete design analysis including velocity, acceleration, and torque calculations for the linear motion system, electrical power calculations for motors and other electrical components, economic analysis for cost savings to the sponsoring company, programmable logic controls and servomotor programming, and selection of an appropriate spray nozzle. The final prototype was successful and installed at the sponsoring manufacturing facility. The student design team was very impressed that the sponsoring company built a new enclosed work area for the machine to be located. The automated application system has been in operation for the past two years. A 3D solid model of the design is shown in Figure 2. Additional projects have involved design and build of various electro-mechanical measurement systems for component testing, variable capacity storage devices for HVAC equipment, automated production equipment, and hardwareisoftware design of circuit board and component test stations. ASSESSMENT RESULTS Sponsors and students are surveyed at the conclusion of each project. The overall response has been positive from both parties. Students are surveyed as part of the exit interview process prior to graduation" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002780_cae.20393-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002780_cae.20393-Figure4-1.png", "caption": "Figure 4 (a) Beginning of maintenance animation of superior trunnion plugs. (b) Removing the screw of crank axis. (c) Removal of the crank covering. (d) Removal of wad. [Color figure can be viewed in the online issue, which is available at www.interscience.wiley.com.]", "texts": [ " According to the type of maintenance chosen by the trainee, the animation is carried out, and the procedural instructions are shown in the text area interaction with the virtual environment is the least possible in this mode. The maintenance procedure is developed through an animation that shows the assembly and disassembly movements of the piece, and the trainee controls the entire process. This learning aims to present the pieces that are involved in the maintenance and to show the correct position of the technician in the HUE structure and the correct sequence in which the pieces must be manipulated. An animation example of the disassembly maintenance procedure of the superior trunnion plug is provided in Figure 4a,b. The Figure 4a shows the beginning of the assembly procedure. In the bottom part of the text screen, the system asks the user to click on the \u2018\u2018Animation\u2019\u2019 button to start the procedure, and the image in the center of the screen shows the removal of the eccentric pine (inside the vertical circle) and the screws from group A (inside the horizontal circle). The remaining consecutive steps of the maintenance procedure are shown in the Figure 4b d. The VGU system offers two types of visualization for this maintenance step. The first is the external superior vision of the Cone, which allows the user to visualize the withdrawal of the lid of the cone floor from the outside, and, from the opening in the lid of the floor, the user can manipulate the pieces in the maintenance area (see Fig. 4b). Figure 4a shows the second type, namely, the inferior lateral view inside the Cone. Guided Mode. In this learning mode, the system guides the user by commands that detail the maintenance procedures. In contrast to the automatic mode, here the trainee has to manually select the maintenance tasks using the mouse, which allows the selection and movement of pieces through picking events (virtual choice of objects by clicking on the mouse). When a piece, which is part of the maintenance context, is clicked on, a message informs the name of the object in the scene and, simultaneously, the instructions related to the object are shown in the text inferior area" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001255_bf03266535-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001255_bf03266535-Figure7-1.png", "caption": "Figure 7 \u2013 Schematic representation of the laser brazing", "texts": [ " Due to the low heat input a narrow evaporating zone of a few tenths of a millimetre only is formed right next to the brazing seam, which remains protected against corrosion due to the distant cathodic effect of the zinc without any problems. Based on the properties and advantages shown the laser brazing process is certainly an interesting alternative jointing technology especially compared to the conventional welding process, even if the filler material is still significantly more expensive than a conventional welding wire. Figure 7 shows the schematic representation of the laser brazing process. With the new brazing process it is now possible to joint this type of coated material by meeting the high requirements. Usually, a copper-based brazing filler metal is used as filler material. Different alloys are available. In line with the Cu-Si two-component phase plot there is a melting range between 950 oC and 1 050 oC for SGCuSi 3. Other filler metals on the basis of low-meltingpoint copper-based alloys are also SG-CuAl 8 and SGCuSn" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.94-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.94-1.png", "caption": "Fig. 14.94 Winch mounted on boom", "texts": [ " Machinery of this type is intended for lifting and transferring loads of up to 200 kg to a height of 80 m. The design and technical specifications of these winches make them a highly effective means of vertical transport in construction work involving scaffolds as well as the assembly and disassembly of scaffolds. Winches in scaffold cranes can be mounted in two ways: \u2022 Outside the crane, to the lowest (from the ground) scaffold upright (Fig. 14.93)\u2022 On the crane\u2019s boom (Fig. 14.92 item 5, and Fig. 14.94) In the case of winches mounted using the former method, a limit switch, functioning also as a load limiter and a block upper position switch, is incorporated into the winch\u2019s housing. The way in which a winch is mounted onto the boom is shown in Fig. 14.94 and Fig. 14.92 item 6. The working radius of the boom with a mounted winch can be changed by protruding the load-bearing tube. There is a series of holes in the inner tube for a blocking pin. The boom with the winch can be attached in a slewing mode to all kinds of support elements (Figs. 14.95\u201314.98). The advantage of winches mountable on booms is their simple design and assembly owing to the elimination of intermediate cable pulleys. Their disadvantage is the unfavorable weight and load distribution along the boom\u2019s end, resulting in the increase in the forces needed to slew the loaded boom and in heavier loading of the load-bearing structure" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002657_6.2009-6274-Figure13-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002657_6.2009-6274-Figure13-1.png", "caption": "Fig. 13", "texts": [ " x 2 = \"x 1 = sin \"1 L 1 2R # $ % & ' ( x 3 = V )R In a manner similar to the approach developed in Section C the stability of this three dimensional state space system can be studied by construction of D sets. However the two dimensional method must now be generalized to three dimensions. As described earlier, each two-dimensional D set is a closed curve so it encloses an open set of points that constitute the interior of D. The union of D with its interior, defined as ! D, is a closed set, defined as ! D. ! D = DUD The desired three-dimensional invariant sets are created by the intersections of extrusions of two-dimensional ! D sets, which are illustrated in Fig. 13 and defined as follows: Definition of D Set Extrusions-Given a set ! D 12 , that is the boundary of a closed set ! D12 , which is a subset of the two dimensional space ! \"2 with coordinates ! x 1 , x 2 , then the extrusion of ! D12 into the three space ! \"3 , with coordinates ! x 1 , x 2 , x 3 , is defined as ! D12\"123 = {x 1 , x 2 , x 3 # $3 : x 1 , x 2 # D12} ! In similar fashion given a set ! D 23 , that is the boundary of a closed set ! D23 , which is a subset of the two dimensional space ! \"2 with coordinates " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003284_s12239-010-0071-8-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003284_s12239-010-0071-8-Figure3-1.png", "caption": "Figure 3. Pump module integrated with the motor and vane pump.", "texts": [ " This paper will describe the development of the electric motor-driven pump unit and IPMSM for EHPS system, introduce the equivalent magnetic circuit method to analyze and design the IPMSM, and finally provide experimental results of the EHPS system. 2. DESCRIPTIONS OF EHPS COMPONENTS 2.1. Hydraulic Pump The vane type pump, which has a complicated structure, is used for the EHPS system. This pump has several advantages such as high efficiency, low acoustic noise and small pulsation. The pump has 100 bar of pressure at the relief condition. A suction port and a discharge port are integrated into the pump. The suction port opens when the pressure is too high and all of the oil comes out through the discharge port as shown in Figure 2. Figure 3 displays the pump module integrated with the developed pump motor and its vane pump. The ECU commands a suitable hydraulic pressure based on the actual motor speed, which is obtained from the hall sense in the pump motor, the car speed and the turn speed of the steering wheel. Figure 4 shows the block diagram of the ECU. The pump motor is an IPMSM with a six poled rotor. The speed of the motor is controlled by a current vector control module using a maximum torque-per-current (MTPC) control (Hur, 2008)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003409_elan.200900548-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003409_elan.200900548-Figure1-1.png", "caption": "Fig. 1. (A) Flexible graphite foil. (B) Assembled graphite electrode (for use in batch cells): a) Active electrode area; b) Insulated area (toner mask covered with varnish); c) electrical contact (done by wrapping wire and graphite tightly with PTFE tape); d) electrical wire. (C) Individual parts of the FIA cell: a) inlet (polyethylene tube); b) outlet (stainless steel tube which is also the counter electrode); c) polyethylene spacer (central hole defines the working electrode area); d) graphite foil (4 2 cm) as the working electrode; e) miniaturized Ag/AgCl reference electrode; and f) screws. (D) Fully-assembled FIA cell.", "texts": [ " A miniaturized batch cell (5 mL inner volume) was constructed in the lab utilizing 10 mL polyethylene measuring cylinders (which were cut in their middle). A polyethylene cover was built and four holes were drilled into it, where the three electrodes \u2013 a miniaturized reference Ag/ AgCl electrode (saturated KCl, constructed in the laboratory [22]), a platinum wire and the working electrode \u2013 were positioned. The fourth hole allowed the addition of solutions without opening the cell. For flow injection studies utilizing graphite foil electrodes, a flow cell was designed. Figure 1 depicts (A) the graphite foils utilized, (B) a graphite electrode used in the batch measurements, (C) the components of a flow cell and (D) the flow cell assembled. An Autolab PGSTAT20 potentiostat from EcoChemie (Ultrech, The Netherlands) was used for all voltammetric studies. A Microsonic SX-20 (Eurosonic, Brazil) ultrasonic bath operating at 20 kHz was used to accelerate the acidic decomposition of the organic matrices prior to zinc determination by stripping analysis. A Hewlett Packard 8452A UV-VIS spectrophotometer with a quartz cell (optical path of 1", "5 V (for 20 s); frequency\u00bc 20 Hz; amplitude\u00bc 190 mV; step\u00bc 5 mV. To demonstrate the potential of this new source of electrodes for electroanalytical applications coupled to flow 1292 www.electroanalysis.wiley-vch.de 2010 Wiley-VCH Verlag GmbH & Co. KGaA, Weinheim Electroanalysis 2010, 22, No. 12, 1290 \u2013 1296 injection analysis, a very simple flow system was assembled. It consisted of a reservoir containing a KCl solution (0.1 mol L 1), fixed on the bench and connected to the flow cell (depicted in Figure 1D) by a polyethylene tube. The reservoir height was used to control the flow rate and a manual injector was connected in front of the electrochemical cell. With this system, a frequency of 60 samples/h can be attained. An even higher frequency can be reached if higher flow rates are utilized. The oxidation of ascorbic acid is demonstrated by cyclic voltammetry in 0.1 mol L 1 KCl electrolyte solution. Preliminary experiments showed an oxidation peak starting at 0.25 V and reaching a maximum value at 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000980_1-4020-3559-4-Figure26-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000980_1-4020-3559-4-Figure26-1.png", "caption": "Figure 26. The European satellite with the folded and unfolded configurations of the antenna.", "texts": [ " The reduced equations of motion for a linear flexible body are [5][ Mr MrfX XTMfr I ]{ q\u0308r w\u0308 } = { gr XTgf } \u2212 { sr XT sf } \u2212 { 0 \u039bw } (31) where \u039b is a diagonal matrix with the squares of the natural frequencies associated with the modes of vibration selected. For a more detailed discussion on the selection of the modes used the interested reader is referred to [5]. The methodology is demonstrated through the application to the simulation of the unfolding of a satellite antenna, the Synthetic Aperture Radar (SAR) antenna that is a part of the European research satellite ERS-1, represented in Fig. 26. During the transportation the antenna is folded, in order to occupy as little space as possible. After unfolding, the mechanical components take the configuration shown in Fig. 26(a). The SAR antenna consists in two identical subsystems, each with three coupled four-bar links that unfold two panels on each side. The central panel is attached to the main body of the satellite. Each unfolding system has two degrees of freedom, driven individually by actuators located in joints A and B. In the first phase of the unfolding process the panel 3 is rolled out, around an axis normal to the main body, by a rotational spring-damper-actuator in joint A, while the panel 2 is held down by blocking the joints D and E" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000946_j.snb.2004.12.008-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000946_j.snb.2004.12.008-Figure1-1.png", "caption": "Fig. 1. Exploded view of the disposable glucose biosensor used in this work.", "texts": [ " lumina nanoparticles, with particle size ranging from 10 o 1000 nm, and PVPAC were synthesized in house. The hosphate-buffer saline solution (PBS pH 7.4) was preared from phosphate salts (0.020 M) and sodium chloride 0.15 M). The PVFcAA, glucose and GOX solutions were repared with the PBS buffer. Glucose stock solution was llowed to mutarotate for at least 24 h before use. All other hemicals were of certified analytical grade and all solutions ere prepared with Millipore deionized water. .2. Preparation of the disposable glucose biosensor Fig. 1 illustrates the fabrication of the disposable glucose iosensor. The nanocomposite membrane was screen-printed nto carbon stripusing an aqueous slurry \u2018ink\u2019 of PVFcAA, GOX, a PVPAC binder and alumina nanoparticles. A series of patterns were used to screen print layers of carbon and the nanoparticulate membrane onto a 300 m polyester film. The first step was printing the carbon substrate electrode. An array of 5 mm \u00d7 30 mm carbon electrodes was printed on the polyester film using Electrodag 423SS carbon (Acheson Colloids Co" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000759_imece2004-60714-Figure13-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000759_imece2004-60714-Figure13-1.png", "caption": "Figure 13. Bending of long vesicle beam by different pressure configuration: (A) UP-UP-UP-UP-UP-UP; (B) UP-DOWN-DOWN-DOWN-UPUP; (C) UP-UP-UP-DOWN-DOWN-DOWN; (D) DOWN-DOWN-UP-UPUP-UP", "texts": [ " With Equation (12), we can calculate the work and energy density for the bending actuator to be 5.56kJ/m3, which is significantly lower than the energy density of a stack actuator of a similar configuration (case (C) in Figure 10) which is 118 kJ/m3. The bending actuator is still interesting because the de- Copyright c\u00a9 2004 by ASME url=/data/conferences/imece2004/71571/ on 05/09/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use formation will be amplified significantly when a long beam of actuator is considered. Figure 13 shows deformation of two rows of 30 vesicles under different pressure configurations. In the simulation, a two dimensional model is used and the actuator is modeled as a plain stress problem in order to reduce the computation effort for the complex 3-D model. First the 2-D model and the 3-D model are both applied to the configuration in Figure 11 with five columns of vesicles in order to verify the accuracy of the 2-D model. The results shows both two models consists well and the error of deformation is less than 2%. Then we apply the 2-D model to the long beam which is stacked with six of the five column blocks for a total of thirty columns of vesicles. In all the simulations in Figure 13, the five columns of vesicles in each single block are applied the same pressure pattern . One block can be chosen to either have pressure in the upper row or in the lower row. In Figure 13, the block which is bending up is denoted as \u201dUP\u201d and bending down is denoted \u201dDOWN\u201d. The shape of the long beam can be easily controlled by the pressure configurations applied to each vesicles. The geometry of the configuration of thirty columns of vesicles is quite small, only 3.6\u00b5m long. It is possible that the shape of bulk material can be controlled very accurately at a larger scale. This paper presents a coupled analysis of the use of biological transport mechanisms for producing controllable 8 Downloaded From: http://proceedings" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002239_j.euromechsol.2008.06.008-Figure21-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002239_j.euromechsol.2008.06.008-Figure21-1.png", "caption": "Fig. 21. Spatial 3-link WMM.", "texts": [ " The solution shown in Fig. 20 has been calculated by RPA using Nm = 3 and N p = Na = 32. It has required a runtime of 22 minutes. We note that the arm manipulator is generally oriented towards the center turn of the path; this is in order to compensate, by the weight of the arm manipulator, the influence of the centrifugal force on the dynamic stability of the robot. We consider the minimum-time trajectory problem with constraints on maximum torque and on stability for the spatial 3-link WMM (Fig. 21) whose characteristics are listed in Table 3. The plat- form is the same as that of the planar WMM shown in Fig. 9. The only difference concerns the maximum torques (which now are fixed as follows: \u03c4max p1 = \u03c4max p2 = 20 N m). All inertia moments are given with respect to the center of mass of the corresponding bodies. The moving frames attached to the rigid bodies of the manipulator are defined using the Khalil\u2013Kleinfinger method (Khalil and Dombre, 2002). The workspace is a (6 m \u00d7 6 m) flat floor with one obstacle (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003288_amr.308-310.1513-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003288_amr.308-310.1513-Figure1-1.png", "caption": "Fig. 1. Safety gear brake block.", "texts": [ " The stress and deformation distribution of a cylindrical type safety gear\u2019s brake block was investigated using finite elements and experimental methods. Results found with FEM by using ABAQUS/CAE software were compared to experimental results. It is clearly seen that the element type and boundary conditions used in finite element modelling give satisfactory results. High compressive and friction forces occur during safety gear brake operation of an elevator. These forces occur between guide rail and cylinder (roller) in figure 1, cylinder (roller) and brake block in figure 2. The generated friction forces between these elements provide a secure and safe way to stop the elevator car and cabin. Safety gear operation. During instantaneous brake operation guide rails also deflect. This deflection occur a contact area between rail and block that can lead to bending, compression and torsion stresses. That\u2019s why different guide rail support cases were taken into account. When the distance between the brackets (supports) change; the deflection quantities also change, therefore the effects on the brake block change dramatically" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000939_cdc.2005.1582304-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000939_cdc.2005.1582304-Figure2-1.png", "caption": "Fig. 2. Schematic of the triple inverted pendulum with the mechanical parameters in Table I.", "texts": [ " Treuer, and M. Zeitz are with Institut fu\u0308r Systemdynamik und Regelungstechnik, Universita\u0308t Stuttgart, D\u201370550 Stuttgart, Germany {graichen,treuer,zeitz}@isr.uni-stuttgart.de straightforward manner with the standard MATLAB function bvp4c [17]. Due to the accuracy of the nonlinear feedforward control, a simple linear LQR controller is sufficient as the stabilizing feedback part \u03a3FB in Fig. 1. Experimental results reveal the accuracy and robustness of the control scheme. The triple pendulum in Fig. 2 consists of three links with different lengths li and the angles \u03c6i(t), i = 1, 2, 3 to the vertical. The first link is attached to a cart moving on a rail with the position coordinate xc(t). The meaning and values of the mechanical parameters are listed in Table I. The acceleration x\u0308c(t) of the cart serves as input u(t) = x\u0308c(t) to the system. In the experimental set\u2013up (see Section V), x\u0308c(t) is controlled by an underlying fast PI\u2013controller. Furthermore, the cart is subject to the following constraints |xc| \u2264 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003961_detc2013-12361-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003961_detc2013-12361-Figure10-1.png", "caption": "FIGURE 10. Sun Gear: Spalled Sun Gear Case", "texts": [ " The accelerometers had a bandwidth of 50kHz. The outputs of the accelerometers and tachometer pulses were routed to antialiasing filters and a PC-based data acquisition system. For all test cases and test conditions, data were acquired at 50 kHz sampling rate with a 25 kHz aliasing filter cut-off frequency. The data were acquired for 40 sec. per set. In total, there are two sets of baseline cases, two sets of single spall cases, one test for the multiple spall, and one test for the cracked sun gear. The single spall case is shown in Figure 10. This spall covers around 75% of the toothface. No other ap- preciable damage was noticed via visual inspection on any of the other teeth. Another component tested was the sun gear shown in Figure 11. This gear had four teeth with severe damage. On ST10, a spall about one-third the facewidth exists. On tooth STID12, there is a chip at the tip which extends about 20% of the facewidth. ST14 has spall covering about 80% of the facewidth and, on ST15, almost the full facewidth is spalled. A simulated crack was machined using Electrical Discharge Machining (EDM) and is shown in Figure 12" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001512_1.3453401-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001512_1.3453401-Figure1-1.png", "caption": "Fig. 1 Lubrlcallng film", "texts": [ " 101 I 497 Copyright \u00a9 1979 by ASME Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use relationship of a pseudoplastic (non-Newtonian) lubricant may be rz = yAz + az (10) expressed in the form [13], ^ =r + 0, the inflated membrane belongs to the semi-space z < 0; conversely, if R < 0, the membrane belongs to the semi-space z > 0. By using the above solution, we can obtain the equilibrium shape of a closed membrane made up of two strips of the same width L which are joined along their edges and whose common end-sides, initially placed at s \u00bc 0 and s \u00bc L, are forced to assume the new positions \u00f00;0\u00de and \u00f0nL; 0\u00de, with n < 1. A typical inhomogeneous solution is represented in Fig. 7 for n \u00bc 1=4. In this case, we obtain g \u00bc L=R ffi 5 and R ffi 0:20L. Varying the shortening ratio n changes both the enclosed volume V and the magnitude H of the horizontal load needed to maintain the inflated membrane in the assigned configuration. For n \u00bc 2=p 0:637, the directrix of the inflated membrane is a circle. Here, the volume reaches a maximum, while the horizontal load is zero. When n < 2=p, the volume increases almost proportionally with n, and H is directed inward. In the limit case of n \u00bc 0, the directrix is 8-shaped, the volume is half the maximum and h \u00bc H=pL \u00bc 1=p 0:3183" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002821_j.sna.2010.07.002-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002821_j.sna.2010.07.002-Figure9-1.png", "caption": "Fig. 9. Tangential forces acting on a mass element of the frictional layer in side portion of roller.", "texts": [ " (29) can be used to calculated the normal pressure p(z) in side portion of roller: p(z) = A b [ a ( 2 ) (m\u03c92 \u2212 cN) sin ( 2 L\u2032 ) + cN A \u221a r2 \u2212 (z \u2212 H)2 + dNa\u03c9 ( 2 ) cos ( 2 L\u2032 )] (29) The unknown boundary (c, c\u2032) can be found iteratively by using Eq. (30): FNside = b \u222b c c\u2032 p(z)dz, c = H + r sin \u2032, c\u2032 = H \u2212 r sin \u2032 (30) In the slip zone, the tangential stress in side portion of roller can be obtained as follows: (z) = \u2212sign(vr \u2212 U\u0307z) p(z) (31) In stick zone, Newton\u2019s second low can be applied for tangential element of frictional layer in side portion of roller (Fig. 9) mdzU\u0308k z (L\u2032) = (z)bdz \u2212 cT dzUk z (L\u2032) \u2212 dT dzU\u0307k z (L\u2032) (32) (z) = A b [ (cT \u2212 m\u03c92) cos ( 2 L\u2032 ) + dT \u03c9 sin ( 2 L\u2032 )] (33) 3. Torque and power of rotor The motor torque can be calculated by integrating the tangential stress acting onto the down and side portion of roller. Total tangential stress acting onto a roller consists of two parts, one tangential stress acting onto the down portion rollerdown and the other tangential stress acting onto the side portion rollerside roller = rollerdown + rollerside (34) roller= (\u2212dT vw b Uk\u2032 x (x)+ cT b Uk x (x) ) + (\u2212dT vw b Uk\u2032 z (L\u2032)+ cT b Uk z (L\u2032) ) (35) The torque of each roller can be calculated as follows: Mroller = rb [\u222b L \u2212L rollerdown dx + \u222b c c\u2032 rollerside dz ] (36) The roller-rotor contact is reflected in Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002627_acc.2008.4586712-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002627_acc.2008.4586712-Figure3-1.png", "caption": "Fig. 3. Experimental 3 DoF ViSHaRD3 system.", "texts": [ " (22) Proof: According to Theorem 1, the switching controller (3) stabilizes the closed-loop system (4) if the inequality (9) is satisfied. Subsitute (19), (20) and B\u0304T = [0 BT ] into (9) and let Y (i) = K(i)X11(i). The nonlinear terms in (9) are eliminated and the LMI (21) is derived. In order to validate the proposed switching control approach, experiments of the position control for a 3 DoF robotic manipulator ViSHaRD3 [18] are conducted. The device is equipped with a fixed end-effector and three revolute joints as shown in Fig. 3. Each joint is actuated by a Maxon RE40 DC motor coupled with a harmonic drive gear (gear ratio 1:100). The DC-motor current, resulting in torque, is provided by the PWM amplifier operated under current control. The reference signal is given by voltage from D/A converter and is an output of the I/O board. The ViSHaRD3 device is connected to a PC running RT Linux. The control loop and the communication network, i.e. the transmission delay, are implemented in MATLAB/SIMULINK blocksets. Standalone realtime code is generated directly from the SIMULINK models" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001620_bf00382707-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001620_bf00382707-Figure1-1.png", "caption": "Fig. 1: Coordinate systems used in deriving squeeze film equations.", "texts": [ " The scope of an earlier paper [7] is broadened to include \"flat strip\" squeeze films in addition to circular disc type squeeze films; effects due to fluid inertia will increase in importance as the speed of approach of the surfaces becomes large. It is assumed that the magnitude of the extra stresses in the fluid may be related to the deformation rate by a \"power law\" type of model, as suggested by experiment [2, 3, 14]. Further assumptions involve the sign of the extra stresses and the relationships between the extra stresses. Figure 1 illustrates the co-ordinate systems used in deriving equations. The two surfaces approach each other with a normal velocity v and are separated by a distance h. (i) VISCOUS CONTRIBUTION TO LOAD BEARING CAPACITY-NEWTONIAN FLUID (a) Discs 3 ~ r / x ( - v ) R 4 (1) F 1 = 2 h 3 (b) Flat Strips For a liquid flowing from between flat strips, the volumetric flowrate in each direction caused by the relative motion of the surfaces is q~ = ( - v ) w x (2) THE INFLUENCE OF FLUID INERTIA 221 Now, using a standard equation for the flow of a Newtonian liquid in a parallel-sided slot with w >> L (see, for example, ref 15, with n = 1) wh 3 Op qx - 12ix 3x (3) From equations (2) and (3) Op_ 12tx(-v)x Ox h 3 (4) The solution of this equation under the boundary condition p = 0 at x = L/2 yields the pressure distribution, which is integrated to give the mean pressure in the liquid as / z ( -v )L 2 e - h3 (5) This is balanced by the applied load /x ( -v)wL 3 F2 = h3 (6) This is the equivalent of the Stefan equation for the case of flat strips" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000900_wcica.2006.1713769-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000900_wcica.2006.1713769-Figure4-1.png", "caption": "Fig. 4 Coordinate frames on the robot head.", "texts": [ " A fitted circle region is detected through the Hough transform (Fig .3). The centroid of the circle is selected as the target point in image plane. As the target is a spherical ball, two target points in image planes represent the same point in three-dimensional space. Then we make two eyes converge at the fixation point using PD control. To maintain fixation on the same point in the space, eyes should make appropriate movements to compensate for the head motion. We define seven coordinate frames on the robot head using Denavit-Hartenburg (D-H) notation [12] (Fig. 4). Coordinate frame {B} is the base frame fixed in the robot head. Joint angles 1 and 2 are neck tilt and pan angles, respectively. Joint angel 3 is the two eyes common tilt angle. Joint angles 4 and 5 are left eye and right eye pan angles, respectively. At time t , the head gazes at the red ball and its two eyes converge at the same point in the space. Therefore the fixation point is 0 0 1 T L lZ z in the left camera frame and 0 0 1 T R rZ z in the right camera frame. Joint angles 1 2 3 4, , , and 5 are known" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002061_6.2007-6525-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002061_6.2007-6525-Figure2-1.png", "caption": "Figure 2. Rail Based \u201cPitch And Plunge Apparatus (PAPA)\u201d - 2 DOF: Pitch and Plunge.", "texts": [ " 2 except for the way it is mounted in the wind-tunnel. The model consists of a rigid inboard spar with fiberglass skin and a flexible main aluminum spar shaped to reflect the structural bending characteristic of a full-scale concept wing. The wing, pictured in figure 1, has four evenly spaced trailing edge control surfaces and one leading edge surface near the tip. The model has an eleven foot span and is attached to the wall of the TDT test section using the rail based \u201cPitch And Plunge Apparatus (PAPA),\u201d depicted in figure 2, that affords 2 DOF motion: pitch and plunge. The wind-tunnel model is instrumented with accelerometers along the spar, strain gauges at the root and mid-spar, a rate gyro at the wing tip, a gust sensor vane in front of the wing, and a rate gyro and accelerometers at the tunnel attachment point. The layout of the associated instrumentation on the model is presented in figure 3. The accelerometers, strain gauges and rate gyro allow the control system to sense the bending modes and the structural stresses" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001215_s00707-005-0298-z-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001215_s00707-005-0298-z-Figure1-1.png", "caption": "Fig. 1. Functional schematic of a robotic bevel-gear train", "texts": [ " The theory is demonstrated by the kinematic and static moment analyses of two articulated robotic mechanisms. Acta Mechanica 182, 265\u2013277 (2006) DOI 10.1007/s00707-005-0298-z Acta Mechanica Printed in Austria Usually a robot manipulator is an open-loop kinematic chain since it is simple and easy to construct. However, it requires the actuators to be located along the joint axes increasing the inertia of the manipulator system. In practice many manipulators are constructed in a partially closed-loop configuration to reduce the inertia loads on the actuators. For example, the Cincinnati T3 shown in Fig. 1 uses a three roll wrist mechanism which is made of a closed-loop bevel-gear train [13]. Gear trains are commonly used to transmit power or motion between shafts with small offset distance. It is necessary to add intermediate shafts and idler gears in order to keep the size of the gears reasonably small when the center distance between two offset shafts becomes large. An alternative method for reducing the inertia of power transmission is to use tendons or belts and pulleys. This type of mechanisms is called tendon-driven mechanisms" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002131_2009-01-2097-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002131_2009-01-2097-Figure2-1.png", "caption": "Figure 2 Factors determining dynamic characteristics of subframe", "texts": [ " In the light of NVH, merits and demerits of a solid mounted subframe for various powertrain types were examined, and several countermeasures such as structural reinforcement, dynamic dampers and dynamic characteristics of A and G bushes were applied and suggested. The front subframe has an important role in NVH, ride, handling and crash performance especially of FF cars. In view of NVH, Figure 1 shows that the subframe is an engine force path from roll mount bush to body and is a road force path from tire to body. Thus the dynamic characteristics of subframe is an important factor in the engine and road noise of cars. As shown in Figure 2, main factors determining dynamic characteristics of the subframe are structural stiffness of its members and stiffness and damping of its mount bush and A,G bushes. In this study, the relationship between these properties and NVH performance of cars is identified and optimization methods were proposed. VIBRATION ISOLATION OF SUBFRAME SYSTEM From the modelling of the subframe system shown in Figure 3, the effective stiffness of subframe system consists of the serial connection of the member stiffness and the mount bush stiffness" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002490_wcica.2008.4592891-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002490_wcica.2008.4592891-Figure1-1.png", "caption": "Fig. 1 The power circuit of brushless DC motors", "texts": [ " In fact, because of brushless DC motors commutate every 60 degrees the magnetic field circumrotate unequably but jump every 60 degree, so the torque is not even. We analysized this kind torque ripple theoretically through analytical formla and numerically by finite element method, and pionted out that the magnetic field jump caused by control pattern is another important reason of torque ripple. A new PWM control algorithm was given out to eliminate this kind torque ripple. .THE ANALYSIS OF TORQUE RIPPLE CAUSED BY STATOR The power circuit of brushless DC motor is as Fig.1, the control pattern is as Fig.2. The commutation rule of brush -less DC motors is: two power IGBTs, one upper and one lower, conduct at any time, and commutation occur once every 60 degrees electric angle. For instance, T1 and T2 conduct between t1 and t2 , T1 is turned off and T3 is turned on at the instant of t2, while T2 retains on for 120 degrees electric angle. The the jump of stator magnetic field is as Fig3, for the convenience of analysis, we take a two pole, concentrated winding brushless DC motor as an example" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001312_3-540-45118-8_2-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001312_3-540-45118-8_2-Figure8-1.png", "caption": "Figure 8. Robot system on ETS-VII.", "texts": [ " Due to the limitation of the on-board arm controller specification of ETS-VII, however, we could not apply the original PD-type controller. Instead, we slightly modified the control scheme so as to match the on-board arm controller specification. Unfortunately the stability condition of the modified controller is not simple like eq.(3), but one can guarantee the system stability with appropriate gains[7]. Figure 7(a) shows the configuration of the experimental system. Figure 7(b) shows the overview of the control station. A 2-DOF force feedback joystick, which is shown in Fig.7(c), was used for the master handle. Figure 8 shows the robot experimental system on the ETS-VII and the task board used in the experiment. Unfortunately, we could not conduct the benchmark test using LEGO in space. Instead, several tasks, such as pushing task and slope tracing task, pegin-hole task, and slide handle task, were carried out by using the experimental facilities on the task board shown in Fig.8(b). In pushing task, accuracy of force command was evaluated. In slope tracing task, we did not notify the operator the starting point on the slope and examined how accurately the operator can recognize the constraint surface shape. In peg-in-hole task, the accuracy of recognizing contact state transitions by the operator was checked. Finally, in slide handle task, we made the sliding direction unknown to the operator by inserting an arbitrary rotational coordinate transformation and evaluated the accuracy of recognizing this unknown constraint direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002918_iecon.2010.5675069-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002918_iecon.2010.5675069-Figure2-1.png", "caption": "Fig. 2 Variation of IPMSM PM flux linkage", "texts": [ " The second step (innovation step) corrects the predicted state estimate and its covariance matrix through a feedback correction scheme that makes use of the actual measured quantities. This is realized by the following recursive relations. | | 1 | 1 | | 1 | 1 [ ( )]ek k ek k k k ek k k k k k k k k k x x K y h x P P K H P \u2212 \u2212 \u2212 \u2212 = + \u2212\u23a7\u23aa \u23a8 = \u2212\u23aa\u23a9 (17) Where | 1 ' ' 1 | 1 | 1( ) ( ) / ek k k k k k k k k k k x x K P H H P H R H h x x \u2212 \u2212 \u2212 \u2212 = \u23a7 = +\u23aa \u23a8 = \u2202 \u2202\u23aa\u23a9 . B. IPMSM Model Based on EKF In this paper, sensorless control system of IPMSM is considered. The amplitude and phase angle of permanent magnet flux vector may change with different operation condition as shown in Fig. 2, so that equation (3) can be rewritten as below\uff1a q q rqd d d d d d d q q qd d rd q q q q \u02c6d \u02c6 d d \u02c6 \u02c6 d L ii Ri u t L L L L i Ri uL i t L L L L \u03c9 \u03c8 \u03c9 \u03c9 \u03c9\u03c8 \u23a7 = \u2212 + + +\u23aa \u23aa \u23a8 \u23aa = \u2212 \u2212 \u2212 +\u23aa\u23a9 (18) In order to observe the rotor flux linkage, the rotor flux linkage is chosen as one state variable. Because the rotor flux linkage cannot change sharply, the deviation of the rotor flux linkage is set to zero. rd rq d 0 d d 0 d t t \u03c8 \u03c8 \u23a7 =\u23aa\u23aa \u23a8 \u23aa =\u23aa\u23a9 (19) By choosing the new set of variables and rotor flux linkage as the state variables, a four-order Kalman filter is constructed" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002541_s12206-009-0326-3-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002541_s12206-009-0326-3-Figure2-1.png", "caption": "Fig. 2. Coordinate system and notations of MUUTV.", "texts": [ " The related research on the modeling is as follows. Gertler and Hagen announced standard equations of motion of submarine simulation in an NSRDC report [8]. Based on the standard equations of motion, Abkowitz and Feldman presented equations of motion which are more close to a real situation [4,7]. Healey and Lienard presented equations of motion about NPS AUV II [1]. The motion of underwater vehicles is analyzed by considering two coordinate systems which are bodyfixed frame and earth-fixed frame as shown in Fig. 2. The 6DOF equations of motion are represented in [6] where the external forces and moments are based on [4] and the hydrodynamic coefficients are estimated from the PMM test. A simulation program was developed with the pro- posed mathematical model by using MATLAB/ SIMULINK as shown in Fig. 3. The controller block is composed of a sliding mode controller and a PID controller for the depth and heading control. Using the developed simulation program we analyzed the dynamic performance. The MUUTV has an initial speed of 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003104_cdc.2009.5400803-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003104_cdc.2009.5400803-Figure2-1.png", "caption": "Fig. 2. Surface plot of the mean square distortion for different average sojourn times. For these plots, a = 1 and h = 0.1 have been used.", "texts": [ " Obviously, J is increasing and convex in TB in Equation (11), and increasing and concave in TB in Equation (12). For the case a < 0, limTB\u2192\u221e J(TB) = A2 and thus we can identify the maximum mean square distortion for a stable process to be: A2 = \u03b1 4a2h(1 + \u03b1) ( e2ah \u2212 1 ) \u2212 1 2a (13) Simulations to see how the mean square distortion varies with the average sojourn times of the two Markov states were carried out. The result from computations for an unstable system (a = 1) with a step size of h = 0.1 is shown in figure 2. As can be seen in the figure, it is when the average sojourn time for the good state is short and the average sojourn time for bad state is long that the performance starts to deteriorate significantly. Translated into parameters of the Gilbert-Elliot model this means that p and q are both large. It is also obvious that distortion becomes unbounded for small TG when TB grows beyond 5.5, which corresponds to the critical value for convergence that can be derived analytically. For this simulation p was varied between 0.01 and 0.5 and q between 0 and 0.81. The same plot as in figure 2, but for a stable system (a = \u22121) is shown in figure 4. For these computations p was varied between 0.01 and 0.5 and q between 0 and 0.95. As can be seen in the figure, J is never unbounded for the stable system in contrast to for the unstable system. To illustrate the conclusions drawn in section IV, computations with an unstable system (a = 1) and a stable one (a = \u22121) were carried out. In the computations, the loss probability was set constant to ploss = 1/10, ie \u03b1 = 9, while TB was varied. The results are shown in figure 6" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001822_1.2825392-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001822_1.2825392-Figure1-1.png", "caption": "Fig. 1. View of ball and cue looking down onto the table. Sidespin is generated when the ball is struck off-center. The ball is deflected by an angle away from the line of approach of the cue.", "texts": [ " In this paper, experimental and theoretical results are presented showing how the friction force varies with the impact parameter in such a way that the squirt angle remains relatively small when the ball is struck not too far off center and when the tip is chalked. A large fraction of the elastic energy 205 Am. J. Phys. 76 3 , March 2008 http://aapt.org/ajp Downloaded 30 Sep 2012 to 136.159.235.223. Redistribution subject to AAPT stored in the tip and the shaft is recovered during the latter stages of the impact, with the result that a cue tip has elastic properties similar to that of a superball. Figure 1 shows a situation where a cue stick approaches a billiard ball of radius R with impact parameter b, where b is the perpendicular distance from the center of the ball to the line of approach of the contact point. The tip of a cue has a radius of about 5 mm, so the contact point on the cue tip does not generally coincide with the central axis of the cue. The normal reaction force N acts along a line from the contact point to the center of the ball, while the friction force F acts at right angles to N", " The resultant force T acts at an angle to the radius vector, where tan =F /N= defines an effective coefficient of friction between the cue tip and the ball. If the tip slides on the ball, then = k, the coefficient of sliding friction. If the tip grips the ball, then k. If the ball is initially stationary, it will exit from the cue in a direction parallel to T at an angle to the line of approach of the contact point. The angle , commonly known as the squirt angle, describes the undesirable deflection of the ball from the intended path. From the geometry of Fig. 1 b it can be seen that + + =90\u00b0, where cos =b /R. The angle is therefore a relatively simple function of the impact parameter and the coefficient of friction and is nominally independent of the mass of the ball and the mass or length of the cue stick. Figure 2 shows a graph of versus for several values of b /R. The main physics question of interest is how the coefficient of friction varies in such a way that the squirt angle remains small regardless of the impact parameter. A related question is why some billiard cues generate smaller squirt angles than others" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002180_rcs.166-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002180_rcs.166-Figure1-1.png", "caption": "Figure 1. (a) Scheme showing polariscope visualization of the birefringence in the vascular model. (b) Arrows indicate the birefringence produced on the vascular model wall by a guide wire", "texts": [ " Transparent membranes of isotropic materials, such as polyurethane elastomer, show a temporal birefringence when an external stress is applied. This birefringence is caused by the change on the refraction indices of the material produced by the applied stress. Therefore, when plane polarized light passes through a patientspecific vascular model, a phase difference ocurs between the wave components in each principal stress direction. This phase difference produces a clear rainbow-coloured pattern and allows us to study the stress on the vascular model membrane (Figure 1a). The observed photo-elastic pattern is caused by the sum of intensity of each wavelength. A segment of a common carotid artery was modelled with polyurethane elastomer, using the technique described in (5), to produce a birefringence when an external stress is applied with a catheter to the vascular model wall. The polyurethane elastomer patient-specific vascular model was fixed, as shown on Figure 2, in an acrylic case filled with glycerin, to give to the vascular model the necessary rigidity for preserving the shape of the vascular model during several catheter insertions. This transparent system was then placed inside a polariscope and a video camera was placed in order to register the rainbowcoloured pattern seen in the polariscope output produced when stress is applied to the vascular model membrane (Figure 1b). Copyright 2008 John Wiley & Sons, Ltd. Int J Med Robotics Comput Assist Surg 2007; 3: 349\u2013354. DOI: 10.1002/rcs The LSM proposed by F. Arai (6) was selected as the catheter-driving mechanism because it was shown in (10) to have the capability to be commanded by the computer to reproduce a planned catheter insertion path and to drive the catheter with a discrete motion. The LSM unit has a linear degree of freedom (DOF) to insert and extract the catheter at variable speed and a rotational DOF to twist the catheter in one direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002859_20090630-4-es-2003.00151-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002859_20090630-4-es-2003.00151-Figure2-1.png", "caption": "Fig. 2. Influence of tire pressure on a) wheel radius; b) torsional wheel oscillations", "texts": [ " Secondly, parametric spectral analysis methods will be applied to the wheel speed oscillations to estimate changes in tire pressure from changes in the wheel speed\u2019s spectrum. As a third method, the vertical wheel acceleration will be investigated via spectral analysis to detect changes in tire pressure. Then, the results from all three methods will be compared and a strategy for the fusion of different indirect methods is proposed. As already mentioned, one option for indirect tire pressure monitoring is to compare the wheel speeds. As shown in Fig. 2a) the vertical force Fz of a tire is supported by the tire wall with its stiffness cw and by the pressure p inside the tire. A loss in tire pressure will be compensated by an increase in the contact area Ac and an increasing compression of the tire wall. Both result in a decrease of the dynamic tire radius rd. Therefore, the wheel speed increases to achieve the same velocity (Persson et al. 2002, Mayer 1996). However, with increasing speed the tire radius increases again because of centrifugal and speed dependant stiffness effects", " In (Mayer 1996) wheel speed ratios of two neighbouring wheels { rrrl, rr;fr, rl;fl, fr;fl,,with1,,rat \u2208\u2212= jir j i ji \u03c9 }\u03c9 (4) are introduced to detect changes among the wheel speeds. At the mantle of a tire a disturbance torque T is generated by variations in the road-height and the friction coefficient. This and the elasto-kinematic bearings cause oscillations which are transferred from the wheel speed at the mantle \u03c9 to the wheel speed \u03c9 of the wheel rim (Persson 2002, Prokhorov 2005). As shown in Fig. 2b) the tire possesses a torsional stiffness c\u2019 and a damping factor d\u2019 . D M t t Following first principles, equations for the balance of torques at the mantle \u03b8\u03b8\u03c9 && ttDM1 '' dcTJ \u2212\u2212= (5) and at the wheel rim \u03b8\u03b8\u03c9 && ttM2 '' dcJ += (6) can be derived where J1 and J2 are the moments of inertia of the mantle and the rim, respectively. The twist angle \u03b8 between mantle and rim is given by . Applying Laplace Transformation and inserting equation (6) in (5) results in the transfer function \u03c9\u03c9\u03b8 \u2212= M & scJJsdJJsJJ csd sT ssG t21 2 t21 3 21 tt D ')(')( '' )( )()( ++++ + == \u03c9 (7) wb, vertical body acceleration or vertical wheel acceleration " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001980_sisy.2008.4664900-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001980_sisy.2008.4664900-Figure3-1.png", "caption": "Figure 3. Rotation of the supporting foot about its edge", "texts": [ " The ZMP position inside the support area can be easily determined with the aid of force sensors on the sole. All the time the ZMP is within the support area there will be no rotation about the foot edge and the humanoid robot will preserve its dynamic balance. Warning means that the ZMP is coming closer to the foot edge. Let us suppose that a critical situation has evolved and that the humanoid is falling down by rotating about one of the foot axes. Without loss of generality, let us assume that it is the foot front edge, along which we will place the rotation axis (x-axis), as presented in Fig. 3. Based on the D'Alembert pricnicple one can write for a rotating foot equations (2), which contain equations (1) along with some additional terms related to the foot, which in this case moves (there is no link immobile with respect to the ground): ( ) ( ) ( ) ( ) ( ) ( ) ( 0+ + + = + + gravitational gravitationalinertial inertial without foot without foot of rotating foot of rotating foot x x x x gravitational gravitationalinertial without foot without foot of rotating foot y y y M M M M M M M ) 0= (2) Apart from the introduction of new terms, the existing ones also change their values compared with terms of eq", " Hence, in the case the humanoid's overturning, the point at which is formally fulfilled 0=\u03a3 xM and 0=\u03a3 yM , does not represent the ZMP, because dynamic balance has not been preserved. Let us consider the case of regular gait of a humanoid robot with two-link foot (Fig. 4). In that case the link representing toes is fixed with respect to the ground, whereas the link of the rear part of the foot moves. In view of the fact that the gait is regular the ZMP is inside the area covered by the toes link, and the system retains its dynamic balance. We should especially emphasize the difference between the situations illustrated in Figs. 3 and 4. In contrast to that shown in Fig. 3, the humanoid robot whose foot is sketched in Fig. 4, preserved its desired position with respect to the environment. In view of the fact that in the case shown in Fig. 4 dynamic balance has not been lost, the point inside the support area (and this is the area covered by the toes link) for which 0=\u03a3 xM and 0=\u03a3 yM , represents the ZMP. In all above examples, only rigid foot was considered. In the literature one can find examples of imprecise definition of ZMP as the point on the ground for which the ZMP conditions are fulfilled, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000769_1.1631018-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000769_1.1631018-Figure4-1.png", "caption": "Fig. 4 Typical circumferential pressure distribution for the optimal and circular bearings \u201e\u00ab\u00c40.5, a\u00c40.1, and L\u00d5D\u00c40.1\u2026", "texts": [ " The clearance configuration of the designed bearing is compared with that of the full circular bearing in Fig. 3; the radius of the inscribed circle is Cr . A load is placed vertically and downward on the bearing, and the journal rotates counter-clockwise. It is shown in Fig. 3 that the optimal configuration slightly depends on the length-to-diameter ratios. Table 1 shows the values of the Fourier coefficients for three length-to-diameter ratios (a50.1). A typical circumferential pressure distribution of the designed bearing (L/D50.1) is compared at \u00ab50.5 with that of a full circular bearing in Fig. 4. Only the positive values of the pressure are normalized by the maximum pressure value of the optimal bearing for the two bearings. As shown in the figure, the journal Transactions of the ASME s of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F eccentricity develops larger pressure values than those of the circular bearing. Attitude angles of the designed bearings are plotted over a region of Sommerfeld number for the five length-todiameter ratios ~Fig. 5!. Figure 5 shows that the attitude angles increase with increasing length-to-diameter ratios" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000227_ic00198a035-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000227_ic00198a035-Figure4-1.png", "caption": "Figure 4. Variation of log K?I1 for Fe(TPC)+ as a function of the pK, of the pyridine or amine used. The straight line is the least-squares fit of the data.", "texts": [ " The least-squares analysis of the data consistently showed that the spectral changes corresponded to the addition of one pyridine molecule and the loss of one C1- from the complex. In other words, Fe(TPC)(py)Cl was completely formed before Fe(TPC)(py)2+ became the dominant species. The spectral changes for the formation of the mono(pyridine) complex are slight, indicating a weak interaction between the pyridine and the iron. This interaction may be more akin to a solvation effect, rather than a strong covalent bond. log vs. the pKa of the pyridine used is plotted in Figure 4. The slope of the line is 0.63 f 0.08. The only significant deviation from this line (not included in the slope calculation) is benzylamine. Further purification of benzylamine did not affect the results. Discussion A comparison between Fe(TPP)C1 and Fe(TPC)Cl can be seen by examining Tables I and I11 and Figures 2 and 3. The potentials 616 Inorganic Chemistry, Vol. 24, No. 4, I985 Feng, Ting, and Ryan bonding between the ferrous porphine and the pyridine ligand has significant contributions from both u- and a-effects" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001386_2013.35993-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001386_2013.35993-Figure1-1.png", "caption": "FIG. 1 Line diagram of a single front wheeled row-crop tractor showing the vehicle fixed and space fixed coordinate systems.", "texts": [ " However, comparatively little work has been done in the tractor handling area. TRANSLATIONAL EQUATIONS OF MOTION The assumptions used in deriving the mathematical model of the tri cycle-type wheel tractor discussed in this article are: 1 The tractor can be considered a rigid body. 2 The ground surface is nondeformable. 3 The tractor pulls or operates no external implements. 4 The x-z plane is one of tractor symmetry. 5 Aerodynamic forces are ignored. The two coordinate systems used in describing the motion of the tractor are shown in Fig. 1. The xs z ^ 1976\u2014TRANSACTIONS of the ASAE 195 system is fixed in space while the (x, y, z) system is fixed in the tractor with origin at the center of gravity of the tractor. The equations describing the translational as well as the rotational motions of the tractor are written in terms of the coordinate system which is fixed with respect to the tractor frame. Using a bar over a variable to indicate its vectorial nature, the translational equations of motion are succinctly expressed by the vector equation F = mv [1] where F is the vector representing the total external force acting on the tractor, m is the mass of the tractor, and v is the absolute acceleration of the center of gravity of the tractor", " 4 shows the successful simulation of this turn; the direction of tractor travel changing from parallel to the xs axis to a new heading. 1.00 1.00 TiP ANGLE, RAD FIG. 3 Separation of stable and unstable or potential overturning tractor operating condi tions. The simulation of a steering maneuver also clearly showed the action of the differential and the effect of changes in the tire forces, and thus moments, on tractor motion. For example, as the steering maneuver was initiated, the negative roll mo ment resulting from force FV2 (Fig. 1) caused the tractor to rotate counter clockwise about the roll axis (Fig. 5). As the roll angle became greater, the vertical reaction on the left rear tire increased and the load on the right rear tire decreased, resulting in a mo ment resisting further roll of the tractor. Eventually the sum of these two roll moments became positive, the roll velocity became positive and the tractor rotated back toward the equilibrium position. The equilibrium state will be reached after some cycling about this position with the time required to reach equilibrium determined by the tire-soil damping" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000611_j.matchar.2004.08.007-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000611_j.matchar.2004.08.007-Figure1-1.png", "caption": "Fig. 1. Single ball/(V-groove) fatigue tester [8].", "texts": [ " These tests assessed the life in hours of the balls as a function of the following: (1) the original amount of retained austenite; (2) the percentage of retained austenite that transformed prior to spalling; and (3) changes in residual stress. Ten F28.575-mm bearing balls made of 52100 steel were used for this research. The bearing balls were fatigued to spall failure prior to the start of the study with a V-groove single ball tester. In the tester, the bballs laid in between two V-groove raceways which were slanted to 1308. For the purpose of indexing the locations at the surface of the balls where measurements were to be taken, Fig. 1 identifies as 1 and 2 the areas where the ball\u2019s surface was in touch with the slanted surface of the V-groove rings (or raceways)Q [6c]. Furthermore, the areas that were in contact with the V-groove rings and that then exhibited spall failure were identified as failed tracks, areas in contact with the V-groove rings that exhibited no spall failure were identified as undamaged tracks, and areas that were not in contact with the V-groove rings (middle of the bearing balls) were identified as remaining in the unrun condition of the bearing balls" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001698_ac8019619-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001698_ac8019619-Figure1-1.png", "caption": "Figure 1. (A) Fabrication of the injection molding plastic strip with two inserted Au-BPEs. (B) Schematic representation of the electric connector for testing the sensor strips.", "texts": [ " A twoelectrode configuration was used in which one Au-BPE served as the working electrode and the other Au-BPE as the reference/counter electrode. The Au-BPE with a diameter of 1.145 mm was dried at room temperature (25 \u00b1 2 \u00b0C) to 50% relative humidity before use. All applied potentials were versus the reference/counter electrode and carried out at room temperature. Reference values measured by using the YSI 2300 STAT Blood Glucose Analyzer (YSI, Yellow Springs, OH) were used to estimate the accuracy of capillary blood glucose. Electrode Strip. Figure 1 clearly demonstrates the feasibility of fabricating such sensors with the stepwise fabrication process of the proposed electrode strips. Compared to the multiple-step procedure required by screen printing technology, the fabrication steps are largely reduced. Barrel plating of gold was performed by immersing the electrodes into a 2 g/L KAu(CN)2 solution in (27) Lee, C. H.; Wang, S. C.; Yuan, C. J.; Wen, M. F.; Chang, K. S. Biosens. Bioelectron. 2007, 22, 877\u2013884. (28) Cabaniss, G. E.; Diamantis, A" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002090_9780470264003-Figure12.9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002090_9780470264003-Figure12.9-1.png", "caption": "Figure 12.9 Reconfigured view of pressure recorder assembly.", "texts": [ "6) and with Nm = 1, the time-based efficiency is given by \u03b7time min= \u00d7 = \u00d7 \u00d7 = t N T m s 100 9 1 235 2 100 3 8 % . % . % (12.7) The efficiency value is very low, leading us to conclude that the service procedure needs to be simplified, possibly using efficient disassembly methods and reconfigure the assembly so that the items needing frequent service are DESIGN FOR SERVICEABILITY 289 290 MEDICAL DEVICE DESIGN FOR X accessed conveniently. In our example, considering PCB as a primary service structure, the assembly is reconfigured so that the board is on the outermost layer (Figure 12.9). Using the same database values, the estimated time for both disassembly and reassembly, Ts, is 16.5 s. Hence, the new efficiency is \u03b7time min= \u00d7 = \u00d7 \u00d7 = t N T m s 100 9 1 16 5 100 54 5 % . % . % (12.8) This DFS calculation approach can be extended to multiservice procedures, say i = 1, 2, . . . , k. The overall time-based efficiency is a weighted average of the procedures of interest by the failure frequencies, fi. This is given by \u03b7overall = = =\u2211 \u22111 1 1 i k i i k i i f f \u03b7 (12.9) The DFX family provides a systematic approach to analyzing and improving device design from a spectrum of perspectives" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000951_sice.2006.314746-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000951_sice.2006.314746-Figure1-1.png", "caption": "Fig. 1 Mmf distribution in Induction Motor.", "texts": [ " The basic induction motor model is initially presented and then the fundamentals of basic DTC are explained. The concept of new torque ripple reduction method is then discussed. A new switching table for optimum selection of voltage vector is presented. The simulated results are then given to compare the performance of the basic scheme and the new scheme. The dynamic analysis and description of revolving field machines is supported by well established theories [10], [11]. An Induction Motor of uniform air gap, with sinusoidal distribution of mmf is considered as shown in Fig. 1. The saturation effect and parameter changes are neglected. 89-950038-5-5 98560/06/$10 \u00a9 2006 ICASE 3587 The dynamic model of the induction motor is derived by transforming the three phase quantities into two phase direct and quadrature axes quantities. The equivalence between the three-phase and two-phase machine models is derived from the concept of power invariance: the power must be equal in the three phase machine and its equivalent two-phase model [12]. The d and q axes mmfs are found by resolving the mmfs of the three phases along the d and q axes", " The torque comparator conditions are given as dte = I for |te < teref - Ate (13) dte = -I for Ite >2 teref + Ate (14) dte = Ofor teref - Ate < te < teref + Ate (15) 3.4 Optimum Switching Table The voltage vector selection table is given in Table 1. The voltage vector that is most suitable for the obtained flux and torque errors in all the sectors is given. The number in each block of the table indicates the corresponding voltage vectors designated with the same number as the subscript as shown in Fig. 3. 4.12-SECTOR METHODOLOGY The 12 sector method uses the same block diagram as shown in Fig. 1 but the switching table now consists of 12 non zero voltage vectors to select. The flux angle now lies in one of the 12 sectors as shown in Fig. 6. The switching logic is similar to that of basic DTC. For example if the flux angle is in the first sector, then for an increase in flux the vectors U65 and U32 are applied instead of u2 and U6, enabling quick response. For an increase in torque the vectors u65 and U5 are applied instead of u6 and U5. Thus a suitable voltage vector is selected according to the flux and torque requirements" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000601_polb.20802-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000601_polb.20802-Figure2-1.png", "caption": "Figure 2. Molecular orientations of the aligned acrylic films (ca. 200 lm thick). The arrows show the rubbing directions.", "texts": [ " A glass cell with a cell gap of about 200 lm (220\u2013240 lm), in which two glass plates were coated with poly(vinyl alcohol) (PVA) and previously rubbed, was filled with the curable raw materials. The two glass plates were successfully removed after the polymerization to yield the desired freestanding films (ca. 200 lm thick), which were used for polarized optical microscopy (POM) observations and wide-angle X-ray diffraction (WAXD) measurements (with Ni-filtered Cu Ka radiation; R-Axis IV Rigaku Co., Ltd.) to investigate their alignments. The molecular orientations of the three types of films are schematically shown in Figure 2. The 1808-twisted film was obtained by the addition of 0.11 wt % compound 2 to compound 1. The thermal conductivity (k) was obtained with eq 1: k \u00bc a q c \u00f01\u00de where a, q, and c represent the thermal diffusivity, density, and specific heat at 25 8C, respectively. The thermal diffusivity was measured by laser flash analysis with the half-time method with a TC-7000 laser flash thermal constant analyzer (Ulbac Co., Ltd.): a laser pulse was flashed onto a sample surface, and the rate of the temperature increase at the rear surface of the sample was measured" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000759_imece2004-60714-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000759_imece2004-60714-Figure7-1.png", "caption": "Figure 7. Mesh of a finite element model for stack actuator.", "texts": [ " As a stack actuator, the maximum actuation capability is achieved when each vesicles are subjected to maximum osmotic pressure simultaneously. The purpose of the model is to simulate the deformation of the bulk actuator under osmotic pressure. A line of five vesicles embedded in matrix is chosen as prototype of the stack actuator. The reason to choose five vesicle is that it will give more accurate results than only one vesicles while it does not require much computation cost. The finite element model is shown in Figure 7. Only half of the structure is analyzed due to the symmetry of geometry and loads. One end of the stack is fixed and the other end is subjected the compressive external loads. We simulate the deformation of the loaded end under osmotic pressure which is applied to each spherical inner surface. A typical deformation of a stack actuator is shown in Figure 8. In the simulation an osmotic pressure 10MPa and elastic modulus E = 100MPa is used. The maximum deformation is 41nm which is approximately 7% of the total length of 600nm" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002981_s10846-010-9401-3-Figure13-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002981_s10846-010-9401-3-Figure13-1.png", "caption": "Fig. 13 Wheel velocities", "texts": [ " Appendix 2: The Kinematic Model of the Mobile Robot The kinematic model for a differentially driven wheeled mobile robot (as the ones used in our experiments) is selected as the basis for this work: \u23a1 \u23a2 \u23a3 x\u0307c y\u0307c \u03bb\u0307 \u23a4 \u23a5 \u23a6 = \u23a1 \u23a3 cos \u03bb 0 sin \u03bb 0 0 1 \u23a4 \u23a6 [ v \u03c9 ] , y = [ xc yc \u03bb ] u = [v \u03c9]T , (47) where y is the system state; the robot is located at (xc, yc) turning to the right, \u03bb is the robot heading angle with respect to the x-axis, and the control u consists of the linear velocity v and the angular velocity \u03c9. Although in Eq. 47, the controls of the mobile robot are its linear and angular velocities, the actual commands provided to the vehicle are the right and left wheel velocities, Fig. 13. Let vl, vr, and vR represent the velocities of the left wheel, the right wheel, and the robot, respectively. Also, let d be the distance between the two wheels and D be the distance between the right wheel and the instantaneous center of curvature, ICC. The commands generated by the navigation-guidance/obstacle-avoidance algorithm set the linear velocity, vP, and the angular velocity, \u03c9P, of the pursuer: \u03c9P = vP d/2 + D . (48) The motion commands are executed by specifying vl and vr. With d known, it is possible to determine R, the turning radius of curvature of the robot, as the distance between the center of the robot and the ICC" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002131_2009-01-2097-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002131_2009-01-2097-Figure3-1.png", "caption": "Figure 3 Modelling of subframe system", "texts": [ " Thus the dynamic characteristics of subframe is an important factor in the engine and road noise of cars. As shown in Figure 2, main factors determining dynamic characteristics of the subframe are structural stiffness of its members and stiffness and damping of its mount bush and A,G bushes. In this study, the relationship between these properties and NVH performance of cars is identified and optimization methods were proposed. VIBRATION ISOLATION OF SUBFRAME SYSTEM From the modelling of the subframe system shown in Figure 3, the effective stiffness of subframe system consists of the serial connection of the member stiffness and the mount bush stiffness. Thus, the both of them should be considered to increase the total stiffness of the subframe. The force effectiveness(E), the ratio of the transmitted force with bush to that without bush, can be expressed with the stiffness(K) of each part as the equation (1).[5] To get the better isolation, the stiffness of the member and the body should be increased and that of bush should be decreased" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.96-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.96-1.png", "caption": "Fig. 14.96 Portable crane mounted in window opening", "texts": [], "surrounding_texts": [ "A range of portable machines, based on winches and other accessories, for handling materials and transferring light equipment on construction sites has been developed. This machinery includes: \u2022 Scaffold cranes mounted on scaffolds (Fig. 14.92 items 5 and 6, and Fig. 14.95)\u2022 Portable cranes (Fig. 14.92 item 4, and Figs. 14.96, 14.98) fixed to steel supports installed between floors, in window openings or on the roof (Fig. 14.97, basic parameters are shown in Table 14.17)\u2022 Gantries mounted on the roof (Fig. 14.92 item 3), in an opening in the building\u2019s elevation (Fig. 14.92 item 2) or on a scaffold (Fig. 14.89 item 1) Part B 1 4 .6 The main component of the above machines is a universal winch that can work in tandem with various accessories. Figure 14.92 shows the use of scaffold cranes, portable cranes, and small-capacity gantries during building erection. Machinery of this type is intended for lifting and transferring loads of up to 200 kg to a height of 80 m. The design and technical specifications of these winches make them a highly effective means of vertical transport in construction work involving scaffolds as well as the assembly and disassembly of scaffolds. Winches in scaffold cranes can be mounted in two ways: \u2022 Outside the crane, to the lowest (from the ground) scaffold upright (Fig. 14.93)\u2022 On the crane\u2019s boom (Fig. 14.92 item 5, and Fig. 14.94) In the case of winches mounted using the former method, a limit switch, functioning also as a load limiter and a block upper position switch, is incorporated into the winch\u2019s housing. The way in which a winch is mounted onto the boom is shown in Fig. 14.94 and Fig. 14.92 item 6. The working radius of the boom with a mounted winch can be changed by protruding the load-bearing tube. There is a series of holes in the inner tube for a blocking pin. The boom with the winch can be attached in a slewing mode to all kinds of support elements (Figs. 14.95\u201314.98). The advantage of winches mountable on booms is their simple design and assembly owing to the elimination of intermediate cable pulleys. Their disadvantage is the unfavorable weight and load distribution along the boom\u2019s end, resulting in the increase in the forces needed to slew the loaded boom and in heavier loading of the load-bearing structure. The structure of winches with a hoisting capacity of 60\u2013200 kg, employed in portable cranes and gantries, is shown in Fig. 14.99. The characteristic feature of such winches is the use of an electric motor with a built-in brake and the integration of all the units, i. e., the electric motor, the toothed gear, the drum, and the electric control system. The drive units of modern winches commonly incorporate: Part B 1 4 .6 \u2022 Clutchless connection between the motor and the gear transmission \u2013 a gear wheel interacting with the transmission gear\u2019s toothed wheel is mounted in the rotor shaft\u2019s end.\u2022 The cable drum is equipped with bearings internally whereby the transmission and the cable drum are compact. Modern winches are intended for vertical transport for a wide range of construction works. Hence the range of handled construction materials is highly diverse as regards kind, shape, dimensions, and so on. For this reason the manufacturers of light cranes offer a wide range of accessories for securing the load. Examples of accessories for handling different kinds of materials are shown in Fig. 14.100. The use of such elements greatly increases work effectiveness and improves operational safety." ] }, { "image_filename": "designv11_20_0002669_iecon.2008.4758452-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002669_iecon.2008.4758452-Figure2-1.png", "caption": "Fig. 2. Schematic view of a roller bearing.", "texts": [ " This leads to the concept of a (frequency/frequency resolution) dyad, that is usually defined kurtogram [9], fig. 1. The number of levels k is upperly constrained by log2(L) \u2212 7, where L is the number of signal samples. In [10] an optimal algorithm for the computation of the kurtogram is presented that shows how the computational cost is similar to that of FFT. Bearings in general consist of two concentric rings, outer and inner, with balls or rollers between them. Rolling elements are bound by a cage which ensures a uniform distance between them and prevents any contact, fig. 2. Bearing defects under normal operational conditions occur because of material fatigue. At first, cracks will appear on the tracks and on the rolling elements. Then, pitting and tearing off of material can quickly accelerate the wear of a bearing and intense vibrations are generated as a result of the repetitive impacts of the moving components on the defect. Impacts or impulsive forces which rise in the vibration signal of the support in operating conditions when a surface defect appears on a bearing rolling elements are periodic in case of constant speed of the shaft" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002982_10402000802105448-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002982_10402000802105448-Figure2-1.png", "caption": "Fig. 2\u2014Macroscopic side view of the contact area.", "texts": [ " device was used during the second part of the 1960s, the Lingard and Tevaarwerk apparatus was operated in the mid-1970s, and the Gadallah and Dalmaz one in the mid-1980s. Tribogyr is T ri bo lo gy T ra ns ac tio ns 2 00 9. 52 :1 71 -1 79 . the result of a modern design and provides wider operating conditions together with improved state-of-the-art equipment (electronics, sensors, etc.) allowing an independent measurement of forces and moments on each specimen. A more detailed view of the specimens and the contact geometry is shown in Fig. 2. The forces experienced by the upper and lower assemblies are given by four 3-D piezoelectric sensors mounted in each assembly. Thus, in this device, normal contact load, and longitudinal and traverse friction forces on each specimen are directly and independently measured. Moreover, two thermocouples are implemented as close as possible to the contact area. One is placed in the lubricant flow of the contact feeding system (temperature T1, thermocouple n\u25e61 in Fig. 3). The other one slides on the disc specimen, on the same radius RD as the contact one (temperature T2, thermocouple n\u25e62)", " The spin-induced shear is bi-directional and is composed of a longitudinal component (along the main entrainment direction) and also a transverse one. Both longitudinal and transverse shears are caused by the speed difference between two parts (inner\u2013outer, inlet\u2013outlet) of the contact area on the spherical end specimen as shown in Fig. 4. This difference increases when the rotational speed increases and is equal to SP. This additional effect is enhanced because the distance between the center of rotation T ri bo lo gy T ra ns ac tio ns 2 00 9. 52 :1 71 -1 79 . of the spherical end specimen and the contact center (= RBsin\u03bb, see Fig. 2) is of the same order of magnitude as the contact diameter. In classical contact configuration the former distance is much longer and the latter is smaller, thus minimizing this effect. The speed gradients along two lines belonging to the two bounding surfaces in the y-z-plane (i.e., at x = 0) are schematically drawn in Fig. 12. Moreover, the speed vectors outside the contact transverse central plane and therefore at x = 0, gain a new component in the z-direction. This effect is clearly described in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003731_tmag.2011.2151845-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003731_tmag.2011.2151845-Figure1-1.png", "caption": "Fig. 1. Configuration of the proposed FSPMLM.", "texts": [ " Based on the analysis of thrust force of the proposed machine, the thrust ripple caused by reluctance thrust increases with the increasing armature current and cannot be reduced using only assistant teeth. Therefore, a current fault-tolerant control strategy based on current harmonic injection is implemented to ensure that the average thrust force is kept invariant under fault. The harmonic currents are injected into healthy phases to compensate for the thrust ripple caused by cogging and reluctance forces. The finite element analysis (FEA) results under normal and fault conditions are compared and validated by the ones from the experiment. Theconfigurationof theproposedFSPMLMisshowninFig.1. The existing FSPMLM presented in [4] is shown in Fig. 2. Manuscript received February 21, 2011; accepted April 29, 2011. Date of current version September 23, 2011. Corresponding author: H. Yu (e-mail: htyu@seu.edu.cn). Color versions of one or more of the figures in this paper are available online at http://ieeexplore.ieee.org. Digital Object Identifier 10.1109/TMAG.2011.2151845 Fig. 3 shows the differences between two FSPMLMs. The primary teeth of existing FSPMLM are uniformly arranged, which means that the values of tooth pitches of the primary are identical, " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000144_s11249-004-8095-8-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000144_s11249-004-8095-8-Figure1-1.png", "caption": "Figure 1. Physical configuration of a journal bearing.", "texts": [ " The closed form of oil film pressure is theoretically solved from the generalized stochastic non-Newtonian Reynolds-type equation with Gu\u0308mbel\u2019s boundary conditions. Then, the equilibrium solution of the steady state can be obtained. By a first-order perturbation scheme, the eight oil-film stiffness and damping coefficients are calculated. Finally, by the Routh\u2013Hurwitz stability criterion, the stability threshold speed of a rotor bearing is investigated under various steady eccentricity ratios, couple stress parameters and roughness parameters. Considering the physical configuration of the journal bearing as shown in figure 1, the lubricant used is an incompressible Stokes\u2019 couple stress fluid, while the body forces and the body couples are assumed to be negligible. Using the assumptions of hydrodynamic lubrication applicable to the thin-film theory, the nonNewtonian Reynolds-type equation is extended from Christensen [10] and Lin [20]. o ox g\u00f0h; l\u00de @p @x \u00fe @ @z g\u00f0h; l\u00de @p @z \u00bc 6lU @h @x \u00fe 6lh @U @x \u00fe 12lV \u00f01\u00de where g\u00f0h; l\u00de \u00bc h3 12l2h\u00fe 24l3 tan h\u00f0h=2l\u00de and l \u00bc ffiffiffiffiffiffiffiffi g=l p . U and V denote the velocity components of the journal surface in tangential and normal directions and can be expressed in terms of x, de=dt and du=dt follows [16]: U \u00bc Rx\u00fe C de dt sin h Ce du dt cos h \u00f02a\u00de V \u00bc C de dt cos h Ce\u00f0du dt x\u00de sin h \u00f02b\u00de where e is the dimensionless eccentricity of the journal centre, u the attitude angle and t denotes time" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001149_s002211207800261x-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001149_s002211207800261x-Figure6-1.png", "caption": "FIGURE 6. Coriolis contribution to the instability mechanism in a nematic: an inhomogeneous radial fluctuation n, induces a viscous force FV which creates a velocity fluctuation v. When wl > wg, the Coriolis coupling induces a radial velocity u. The corresponding shear rate au/az tends to increase the initial orientation fluctuation.", "texts": [ "18) as in the planar case. Now, owing to the Coriolis force, this fluctuation creates a radial velocity fluctuation (absent in the planar case) given by (5.16) : (5.19) This radial flow is not uniform in space and exerts a tangential viscous torque component on the molecules: r; = a2 e) (20,7,) SN,. 72 (5.20) This contribution is destabilizing for usual .nematics with a2 < 0 and v2 > q3 when oms < 0. For w, > 0 this corresponds to the inner cylinder rotating faster than the outer one (6w < 0 * w1 > w2, figure 6). A parallel analysis could be performed in the case of a purely tangential orientation fluctuation (N+ =+ 0, N, = 0) and would lead to the same conclusion. Indeed one would (5.21) This also corresponds to destabilizing torque when 0,s < 0, since a\u2019 = *(a, +a2) < 0 in general. In order to get an estimate of the threshold one has now to consider the coupling between N, and N+ as described by the effective torque equations. 5.2. Threshold Let us write the effective torque equations (5.10) and (5.11) in terms of dimensionless (5" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000623_1.2118732-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000623_1.2118732-Figure3-1.png", "caption": "Fig. 3 The example system", "texts": [ "org/about-asme/terms-of-use Downloaded F eters used to define this example system and the numbers presented here as results could assume any units as long as they have the correct dimensions corresponding to its type force or length . Position vectors for the three foot contact points C1, C2, and C3 expressed in, and from the origin of the body coordinate frame are given as r\u0304C1 = 4.0,4.0,\u2212 3.9 r\u0304C2 = \u2212 4.0,4.1,\u2212 4.0 11 r\u0304C3 = 0.0,\u2212 4.0,\u2212 4.1 and at each foot contact point C1, C2, and C3, the foot contact coordinate frames are set following the rules defined using the four unit contact force component vectors e , e , e , and e as shown to scale in Fig. 3. The unit surface normal direction vectors uN1, uN2, and uN3 are given as u\u0304N1 = \u2212 0.577,\u2212 0.577,0.577 u\u0304N2 = 0.667,\u2212 0.333,0.667 12 u\u0304N3 = 0.000,0.707,0.707 and their corresponding friction coefficients C1, C2, and C3 are C1 = 0.2, C2 = 0.32, C3 = 0.15 13 The known external force F\u0304O and the known external moment M\u0304O expressed in the body coordinate frame and acting on the origin O are F\u0304O = 0.0,0.1,\u2212 8.0 , M\u0304O = 2.0,0.1,0.0 14 This example system is represented to scale in Fig. 3 with the friction cones shown at each contact point. Now we use this example system to illustrate the process of applying the static equilibrium constraints. First, we can directly solve for the three force components using the strategy illustrated above. The FC1 , FC2 , and FC3 component forces by summing the moments about e , e , and e , and their magnitudes are then 298 / Vol. 128, JANUARY 2006 rom: http://mechanicaldesign.asmedigitalcollection.asme.org/pdfaccess.as FC1 = 1.901, FC2 = 1.876, FC3 = 4" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003427_c1ay05328k-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003427_c1ay05328k-Figure1-1.png", "caption": "Fig. 1 Chemical structures of flavonoids: (a) general structure, (b) apigenin.", "texts": [ " Flavonoids constitute one of the most characteristic classes of compounds containing hydroxyl groups attached to ring structures.26 Many flavonoids are easily recognized as flower pigments in most angiosperm families. However, their occurrence is not restricted to flowers but includes all parts of the plants. They constitute most of the yellow, red and blue colors in flowers and fruits.27 Flavonoids are broken down into categories of isoflavones, anthocyanidins, flavans, flavonols, flavones, and flavanones.32 The molecule structure of apigenin, a derivative of flavonoids, is given in Fig. 1. The main purpose of this study was to demonstrate an electrochemically modified PolyApi/GC electrode in aqueous media by CV, characterize the PolyApi/GC electrode by CV and EIS, propose the structure of the complex formed between the PolyApi/GC electrode with Cu(II), investigate the interference effects and apply the PolyApi/GC sensor electrode for Cu(II) determination at trace levels in soil samples for the first time. Apigenin and other chemicals were of analytical-reagent grade supplied from Sigma-Aldrich" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001149_s002211207800261x-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001149_s002211207800261x-Figure2-1.png", "caption": "FIGURE 2. Pieranski-Guyon instability mechanism: owing to the fluctuation n,, the shear flow exerts a torque r, on the director (a). This induces a fluctuation n1 as indicated on (c ) . Owing to this fluctuation, a viscous torque F, appears, the sign of which depends on that of viscosity coefficient a3 ( b ) . For a8 < 0, this tends to increase the initial fluctuation n, ( c ) .", "texts": [], "surrounding_texts": [ "When n is perpendicular to the shear plane [geometry (3) of figure 11 the nematic looks like an isotropic liquid. However, the study of the flow stability does not reduce to that for isotropic liquids. Indeed any fluctuation away from the direction of the unperturbed orientation feels a part of the viscous torques exerted in positions (1) and (Z), proportional to its amplitude. These torques tend to make the orientation rotate as indicated in figures 2 (a, b). Now as discovered by Pieranski & Guyon (1973), an instability may follow from the coupling between orientation fluctuation components through these viscous torques (figure Zc). Indeed, assume a fluctuation n2 > 0. It induces a torque r2 > 0 which tends to create a fluctuation n, > 0. Now this fluctuation induces a torque rl which reacts on n2. The coupling turns out to be destabilizing when u3 is negative and stabilizing in the opposite case. When u3 is negative, the instability takes place only when the destabilizing mechanism is strong enough to overcome the stabilizing effect of the Frank orientational elasticity. Let us define the characteristic evolution time of orientation fluctuations: where y is the orientational viscosity, K a typical Frank modulus and q the wave vector of the fluctuation. 7 0 = y/Kqz, (2.1) 10-2 From a purely dimensional point of view, one can parametrize the flow by the Ericksen number Er = srO where 70 is evaluated for q N n / d , d being the cell thickness, and one can infer that the instability threshold s, will be given by s, 70 N 1 (where in our case 70 is typically of order 1 8). (2.2) The basic mechanism just described leads to a distortion which is uniform in the plane of the flow (homogeneous instability: Pieranski & Guyon 1973). However a second instability mode is possible with a distortion periodic in the direction of the unperturbed orientation. It isassociatedwith a secondary flow which takes the form of rolls pardlel to the flow direction (roll instability: Pieranski & Guyon 1974a). In order to understand this kind of instability one has to take into account the contribution of velocity fluctuations to the viscous torques exerted on the molecules. In fact, spontaneous velocity fluctuations are much more rapid than orientation fluctuations. Indeed the characteristic time of velocity fluctuations is where p is the densiky ( N 1 g/cm3) and 7 the viscosity. 7, = P/W29 (2.3) If one compares this with the orientation fluctuation time 70, one gets 7,/70 = pK/rq N (2.4) for typical values: y N 7 N 0.1 to 1 and K N 10\" cgs. The evaluation could suggest that velocity fluctuations do not contribute to the instability mechanisms since they are not coherent over a sufficiently long time to be coupled with orientation. However, owing to the special form of the Leslie viscous stress tensor, in a shear flow, a non-uniform orientation induces a visww force P speci$c to nematics. The motion equation for velocity fluctuations may then be simplified as dv at p- = ' I ]Av+P, where Fv is very slowly varying (rate 7c1 < 7 ~ ~ ) . Then one may consider that the viscous force Fv induces the slowly varying secondary flow v roughly given by y A v + P N 0. Couette flow in nematic liquid crystals 277 Now this flow is non-uniform and contributes to torques exerted on the molecules. This contribution appears as intercalated in the basic sequence of fluctuation amplification. Orientation fluctuation -+ (viscous force +velocity fluctuation +) --f viscous torque -+ orientation fluctuation. This implies the notion of effective torques taking into account the effect of secondary flows induced by an orientation fluctuation via the viscous forces; accordingly, one can define effective viscosity coeficients a: and a:. Then, the existence of a roll instability specific to nematics will be subject to the condition a: < 0, where . secondary flow effect af = a3-a1az/r2. (2.5) The instability remains specific to nematics and as a consequence the threshold is again given by a condition of the form (2.2). Now let us summarize the situation in the planar case. (a ) When a3 is negative, the homogeneous instability as well as the roll instability can take place since a: is also negative but the intensity of the mechanisms is different. The type of the instability which occurs can be monitored by a magnetic field applied along the direction of the unperturbed orientation which adds its stabilizing contribution to the elastic restoring torques. In the absence of an external field and under weak fields, the elastic stabilizing contribution is dominant and since the periodic distortion associated with rolls implies a greater elastic expense, the corresponding threshold is higher than \u20acor the homogeneous instability. Conversely, under high fields, the magnetic stabilizing contribution is dominant and rolls which correspond to the strongest destabilizing mechanism have the lowest threshold. Experiments have been performed using the well-known nematic compound MBBA (Pieranski & Guyon 1974a; Dubois-Violette et al. 1977) and the cross-over from the homogeneous instability to the rolls has been found to be in agreement with the sketchy description given above. (a) When a3 is positive, the homogeneous instability disappears and the rolls can take place as long as a3 is small enough [see (2.5)] : 0 < a3 < ara2/v2. Experimental evidence has been given by Pieranski & Guyon (1976) using CBOOA, a nematic compound for which a3 + + 00 close to a nematic-smectic A phase transition. In the following we shall examine the particular contribution of rotation to instabilities which are specific to nematics, that is to say instabilities which result from a coupling between orientation fluctuations and occur at a threshold roughly given by ScTo - 1 (for the homogeneous as well as for the roll instability), where T~ is the time constant characteristic of the evolution of orientation fluctuations. Before we discuss the effect of rotation, let us specify the geometry of the problem, discuss some approximations and present the linearized hydrodynamic equations. 278 E. Dubois- Violette and P. Manneville" ] }, { "image_filename": "designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure8-1.png", "caption": "Fig. 8. The geodesic on the plane.", "texts": [ " (26) yields the equation of the directrix C\u00f01\u00dep : C\u00f01\u00dep : R\u00f01\u00dep \u00bc h sin\u00f0b0 k1 sin d1\u00de \u00bdsin d1e\u00f0k1\u00de \u00fe cos d1k1 \u00f033\u00de Repeating the process of Eq. (29), the frame expression of the directrix C\u00f01\u00dep can be obtained: e \u00f01\u00de 1 \u00bc cos\u00f0b0 k1 sin d1\u00de\u00bdsin d1e\u00f0k1\u00de \u00fe cos d1k1 \u00fe sin\u00f0b0 k1 sin d1\u00dee1\u00f0k1\u00de e \u00f01\u00de 2 \u00bc sin\u00f0b0 k1 sin d1\u00de\u00bdsin d1e\u00f0k1\u00de \u00fe cos d1k1 \u00fe cos\u00f0b0 k1 sin d1\u00dee1\u00f0k1\u00de e \u00f01\u00de 3 \u00bc cos d1e\u00f0k1\u00de \u00fe sin d1k1 9>= >; \u00f034\u00de Similarly, the expressions of C\u00f02\u00dep and C\u00f03\u00dep can be acquired. For example, in coordinate system of gear 3 (shown in Fig. 8), the equation of the directrix C\u00f03\u00dep can be represented as C\u00f03\u00dep : R\u00f03\u00dep \u00bc h sin\u00f0b0 k3\u00de e\u00f0k3\u00de \u00f035\u00de Here e(k3) is circle vector function defined in the coordinate system of gear 3. It can be found that the directrix C\u00f03\u00dep is a straight line on the pitch plane of gear 3. Substituting Eq. (33) into Eq. (15) yields R\u00f01\u00deb : R\u00f01\u00de \u00bc h sin\u00f0b0 k1 sin d1\u00de \u00bdsin d1e\u00f0k1\u00de \u00fe cos d1k1 \u00fe r\u00f0e\u00f01\u00de2 cos h\u00fe e \u00f01\u00de 3 sin h\u00de \u00f036\u00de In short, two proposed schemes have the important practical value and respective characteristics" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002171_j.triboint.2009.05.009-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002171_j.triboint.2009.05.009-Figure1-1.png", "caption": "Fig. 1. Schematic draw and general view of a hermetic reciprocating compressor.", "texts": [ " The feasibility of applying active lubrication to the main bearings of a hermetic reciprocating compressor is studied in this work, with the help of multibody dynamics and fluid film theory. Small-scale reciprocating compressors are of common use to compress coolant gas in household refrigerators and air conditioners. This type of compressors have pistons that are driven directly through a slider-crank mechanism, converting the rotating movement of the rotor to an oscillating motion, as illustrated in Fig. 1. The performance of the bearings affects key functions of the compressor, such as durability, noise, and vibrations. Therefore the study and optimization of the dynamic behaviour of reciprocating compressors, taking into account the hydrodynamics of bearings, can be of significant importance for the development of new prototypes. Several studies related to the modelling of small reciprocating compressors can be found in the literature [3,4], however, only a few of them have incorporated in their models the dynamics of the fluid films" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002900_s10544-008-9283-3-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002900_s10544-008-9283-3-Figure2-1.png", "caption": "Fig. 2 Illustration of (a) hexagonal closed packed microdisc array layout, (b) domain wall approximation of the microdisc geometry for the finite element simulation", "texts": [ " The concentration\u2013overpotential relationship given by the Butler\u2013Volmer equation (Bard and Faulkner 2001) was used to solve for the electrode surface reactions kf \u00bc k0 exp aF E E0\u00f0 \u00de RT \u00f03\u00de kb \u00bc k0 exp 1 a\u00f0 \u00de E E0\u00f0 \u00de RT \u00f04\u00de where kf and kb are the forward and reverse reaction rate constants, respectively, and k0 is the heterogeneous rate constant, taken as 1.4\u00d710\u22122 m/s for the FcCO2H redox reaction. The value of \u03b1 was assumed to be 0.5. The standard potential, E0, was fixed at 300 mV for the purpose of this simulation. R is the universal gas constant (8.314 J/K), T (temperature) was set at room temperature (298 K) and F is the Faraday\u2019s constant (96,487 C/mol). The simulation geometry used is shown in Fig. 2. Hexagonal close packed arrays of recessed microdiscs were simulated using the domain wall approximation approach (Davies and Compton 2005). The domain wall approximation approach dramatically reduces the computation time by reducing the 3D problem to a 2D one (Lavacchi et al. 2006). The radius of the circle inscribed in the hexagon in Fig. 3(a) is the simulation geometry, which is illustrated in Fig. 3(b). The Nernst\u2013Plank relationship without electroneutrality application mode with Lagrange triangular quadratic meshing elements was used to solve the problem", " 2006): h \u00bc 6 ffiffiffiffiffiffiffi Dtc p \u00f05\u00de where, h is the simulation cell height or distance above the electrode surface, D is the diffusion coefficient of the electroactive species and tc is the maximum time required to perform one half of a CV scan. The recess depth was taken as 0.5 \u03bcm; dictated by the thickness of the nitride layer of the MDEA. All boundaries were assigned the insulation/symmetry boundary condition (BC) except for the boundary corresponding to the microdisc electrode\u2013 solution interface (recessed region in Fig. 2), which was assigned a flux BC, with the flux given by the following relationship: M \u00bc Cox kf Cred kb \u00f06\u00de where Cox and Cred are the surface concentrations of the oxidized and reduced species, and kf and kb are the forward and reverse reaction rate constants, respectively. Geometric singularities were particularly meshed at a higher density than other areas. Simulations were performed for all scan rates corresponding to the experimental scans. 3.1 Cyclic voltammetry of MDEA 050, 100 and 250 Multiple scan rate cyclic voltammetry (MSRCV) measurements were performed on the three MDEAs of differing geometries (variable microdisc diameter and density), but of similar electroactive surface area (A=0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003269_13506501jet675-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003269_13506501jet675-Figure3-1.png", "caption": "Fig. 3 Position of the k-thermocouple in the test gears and the surface of the polished gear", "texts": [ " The load-dependent losses of gears Ps and bearings Pbl, that is the total power loss PL minus the no-load loss Pnl, were determined directly by measurement. Using equations (3) to (5), the mean friction coefficient of the gear contact can be calculated. The material for the test gears is case-hardening steel 20 NiCrMo2\u20132. The gears are case hardened to a depth of 0.8\u20131 mm, with a specific surface hardness of 60\u201362 HRC. Test gears were case hardened, ground, and polished, which gives the gear surfaces a mirrorlike finish, as shown in Fig. 3. Polishing has been done using a trough vibrator with chips and compounds, and the final surface roughness Ra-values were close to 0.05 \u03bcm. The main features of the gears are shown in Table 1. The test lubricant was mineral base oil ISO VG 220 with kinematic viscosity at 40 and 100 \u25e6C of 220 and 19 mm2/s, respectively. The density of the lubricant at 15 \u25e6C was 892 kg/m3. The tests were carried out at steady oil inlet temperatures of 40, 60, and 80 \u25e6C and at a lubricant flowrate of 2.0 l/min. Tests were carried out according to Table 2", " The nature of the surface, and thus the bulk temperatures, is the dominant parameter when gear contact scuffing load-carrying capacity is evaluated [27]. The increase in temperature decreases the oil film thickness, which increases the risk of a scuffing failure. The measurement of the surface temperature in gear contacts is a very difficult task and requires a lot of effort to implement it as a routine measurement. Consequently bulk temperature measurement was more convenient for use in this study. In Fig. 6, the bulk tooth temperatures are shown for each of the four locations indicated in Fig. 3. The oil inlet temperature was kept at 60 \u25e6C, but the measured bulk temperatures were higher in every test case. This is due to friction between contact surfaces causing heat, which is partially transferred to the gear. The bulk temperature increases with increasing pitch line velocity due to increasing power losses, that is heat generated in the gear contact. It seems that the bulk temperature is fairly independent of the location of measurement with the pitch line velocities used in this study" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000114_1.1759343-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000114_1.1759343-Figure1-1.png", "caption": "Fig. 1 Hole-entry journal bearing configuration \u201ea\u2026 symmetric, \u201eb\u2026 asymmetric, and \u201ec\u2026 geometric details, coordinate system and thermoelastic deformation", "texts": [ " Ever increasing demand for the hybrid bearings to operate at high speed and the conditions of limit design makes is necessary to design the bearing carefully and precisely. In such cases, stability of the bearing in terms of the threshold speed becomes the most vital design parameter. The analysis presented in the following subsection uses finite element method to model hole entry journal bearing. The mathematical model involves simultaneous solution of Reynold\u2019s and Energy equations in flow domain and Heat conduction and Elasticity equations in solid domain ~Journal and bearing bush!. The bearing configurations considered for the analysis are shown in the Fig. 1~a! and ~b!. The Reynold\u2019s equation governing the flow of an incompressible lubricant in the clearance space of a journal bearing is defined in nondimensional form as 004 by ASME Transactions of the ASME s of Use: http://asme.org/terms Downloaded F ] ]a S h\u03043F\u03042 ] p\u0304 ]a D 1 ] ]b S h\u03043F\u03042 ] p\u0304 ]b D 5VF ] ]a H S 12 F\u03041 F\u03040 D h\u0304J G1 ] h\u0304 ] t\u0304 (1) where, the nondimensional viscosity function F\u03040 , F\u03041 , and F\u03042 are defined as F\u030405E 0 1 ~1/m\u0304 !dz\u0304 , F\u030415E 0 1 ~ z\u0304/m\u0304 !dz\u0304 , and F\u030425E 0 1 ~ z\u0304/m\u0304 !$ z\u03042~ F\u03041 /F\u03040", " It is quite reasonable to consider the bearing shell in this manner as usually the bearing shell is made of comparatively more flexible material than the housing. Therefore, the displacement on the nodes of bush-housing interface is assumed to be zero. $d\u0304%5$0 0 0%T The journal bearing operating under load is required to maintain an appropriate minimum fluid-film thickness to minimize the chances of metal to metal contact. The fluid-film thickness is influenced by thermoelastic deformation of bearing and thermal deformation of the journal as shown schematically in Fig. 1~c!. For a rigid journal bearing system the fluid film thickness expression is given as h\u03045 h\u030401D h\u0304 where h\u03040 is the fluid-film thickness when the journal center is in the static equilibrium position and is given as h\u03040512X\u0304J cos a2Z\u0304J sin a D h\u0304 is the perturbation on the fluid-film thickness due to dynamic conditions. For a flexible bearing experiencing thermal distortion, modified film thickness is defined as h\u03045 h\u030401D h\u03041 d\u0304 f b p 2 d\u0304 f b T 2 d\u0304J T where d\u0304 f b p , d\u0304 f b T are the dimensionless radial deformations in bush due to fluid-film pressure and rise in bush temperature respectively at fluid-film bush interface and d\u0304J T is the radial deformation of journal due to rise in journal temperature" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003622_cca.2010.5611228-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003622_cca.2010.5611228-Figure4-1.png", "caption": "Fig. 4. Transformation of the system input with C\u2217 =C(u\u2217) =Creal(u \u2217 plant)", "texts": [ " The input affine system reads: \u02d9\u0303x = \u23a1 \u23a2 \u23a3 x2 1 M (FCyl\u2212FR(x\u03032)+Mg) \u03baRT m\u03071(u)\u2212 p1x\u03032\u03baA1 L0 + x\u03031 \u2212 \u03baRT m\u03072(u)+ p2x\u03032\u03baA2 L0 +L\u2212 x\u03031 \u23a4 \u23a5 \u23a6 (23) with m\u03071 = a12u\u03c10 pv\u03a8 ( p1 pv ,b ) \u2212a23u\u03c10 p1\u03a8 ( p0 p1 ,b ) (24) m\u03072 = a14u\u03c10 pv\u03a8 ( p2 pv ,b ) \u2212a45u\u03c10 p2\u03a8 ( p0 p2 ,b ) (25) and a12 = a45 = { 1 for u > 0 0 for u \u2264 0 , (26) a23 = a14 = { 0 for u \u2265 0 \u22121 for u < 0 . (27) Here it is assumed that C(u) is linear in u, i. e. C(u) = Cmax \u22c5u. In reality this is, however, not the case. Therefore, a signal transformation is performed before applying the control signal to the plant. This transformation maps the linear input u to the real input uplant : uplant =C\u22121 real(u), (28) where Creal is the true conductance function which is assumed to be strictly increasing and Cumax =Cmax. A graphical representation of the transformation is shown in figure 4. Given a control input u, the plant input uplant is calculated according to Equation (28). Following the SDRE control design description in Section III, the SDC parametrization for the pneumatic actuator reads: \u02d9\u0303x(t) = A(x\u0303)x\u0303+b(x\u0303)u(t). (29) In general there is an infinte number of possible choices for A(x\u0303) and b(x\u0303). As the LQ-like controller is calculated at each simulation step, the SDC matrices can even vary over time to account for varying operation conditions. In the context of the present contribution, however, SDC parametrization is kept constant" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003301_cae.20418-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003301_cae.20418-Figure2-1.png", "caption": "Figure 2 Vector/phasor diagram of a three-phase induction motor.", "texts": [ " In these equivalences, iFe=0 and is \u2212 i\u2032r = ism Vs = (Rsis) + ( j ms 2 Xs is ) + ( j ms 2 Xsmism ) (1) 0 = ( R\u2032 ris s ) + ( j mr 2 X\u2032 r i\u2032r ) \u2212 ( j ms 2 Xsmism ) (2) Rs, Xs , ms, Xsm, Rr, Xr , mr, s, is0 = ism Parameters are usually known and Vs, ir, is, Qs, E \u2032 r = Es can be drawn in the vector/phasor diagrams as follow steps. The vector/phasor diagram in Figure 3, 20 V = 1 cm and 10 A = 1 cm. Step 1: is0 current is taken, \u2212j(ms/2)Xsm \u00d7 ism = E\u2032 r = 0B Step 2: R\u2032 r, s, X \u2032 r , mr are known from tan r= ((mr/2)X\u2032 r )/ (R\u2032 r/2) and \u2032 r is found and the direction of i\u2032r is drawn. Step 3: E\u2032 r = |\u2212j(ms/2)Xsmism| diametric circle is drawn and the A point is shown. The A point shows that where the i\u2032r line cuts the circle (Fig. 2). The reduced rotor current is found using i\u2032r phasor equation as in Equation (3). A0 = R\u2032 r s i\u2032r (3) The real rotor current ir is calculated as in Equation (4) ir = ms mr i\u2032r (4) Step 4: is = 0Q can obtain from is = i\u2032r + ism relation. Step 5: OK = j(ms/2)Xsmism is drawn by taken Es = E\u2032 r Step 6: OM = Vs can obtain from Rsis = KL and j(ms/2)Xs is = LM The software is developed using Delphi 7.0 visual package program and it works in a Windows environment [23]. It is prepared to help students to improve their knowledge about the induction motors" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003887_19346182.2012.663534-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003887_19346182.2012.663534-Figure8-1.png", "caption": "Figure 8. Lines and marks on a ball can be used to determine the location of the spin axis. The black dot marks the position where the axis passes through the surface of the ball.", "texts": [ " The spin axis remains fixed while the ball travels toward the net, so the time for one revolution can be measured in terms of the number of video frames required for a particular mark or pattern on the ball to re-appear in the same orientation. When filming at 600 fps, the pattern of marks re-appeared after about 8\u201310 frames or about 8/600 to 10/600 seconds, giving a ball spin of 60 to 75 rev s21 or 3600 rpm to 4500 rpm. It is much more difficult to locate the spin axis, unless the spin axis happens to be exactly vertical or exactly horizontal. If the spin axis is vertical, then all marks on the ball rotate in a horizontal direction, and a horizontal line around the equator remains horizontal, as indicated in Figure 8(a). If the spin axis is horizontal and points to the camera, then all marks on the ball rotate in a circular path around the middle of the ball. Otherwise, the marks and lines rotate in a manner that can be difficult to interpret. For example, Figures 8(b) and (c) show two positions of the equator, one half revolution apart, when the axis is vertical and the equator line is inclined at an angle to the axis. It might appear that the ball is rotating with topspin, given the rotation of the equator line during half a revolution, but if the axis is vertical then there is no topspin at all. In that case, marks on the ball rotate purely in the horizontal direction and then disappear around the back of the ball. Figure 8(d) shows a situation where the axis is perpendicular to the equator line but the top end of D ow nl oa de d by [ U ni ve rs ity o f D el aw ar e] a t 1 7: 40 2 7 Ju ly 2 01 3 the axis is tilted to the left and is also tilted out of the page. In that case, the equator line appears stationary on video film, giving the false impression that there is no topspin. All marks on the ball rotate in circular paths around the axis, so the axis can be identified by the motion of those marks. If the axis is horizontal and pointing in the same direction as the motion of the ball, then the spin is classified as gyrospin" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000776_j.ijpvp.2005.01.008-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000776_j.ijpvp.2005.01.008-Figure1-1.png", "caption": "Fig. 1. Ring-stiffened cone-cylinder intersection under internal pressure.", "texts": [ " A more recent study by Zhao and Teng [7] examined the buckling strength of imperfect cone\u2013cylinder intersection under internal pressure and established a design proposal in the Eurocode format for practical use. For a review of most of the existing work on this topic, readers may refer to [8] or [9]. Although local thickening alone may be used, it is often desirable and convenient to provide an annular plate ring at the cone-to-cylinder joint to supplement local thickening or as an alternative strengthening measure [10], leading to a ring-stiffened cone\u2013cylinder intersection (Fig. 1). A common form of the ring stiffener is an annular plate. Only limited theoretical work has been carried out specifically on such ring-stiffened cone\u2013cylinder intersections under internal pressure. The plastic collapse strength was studied [10\u201312] and a simple method for predicting the plastic collapse strength has been developed [11,12]. More recently, the elastic buckling strength was studied by Teng and Ma [13]. Two types of buckling mode were identified: a shell buckling mode and a ring buckling mode", " This paper presents the first experimental study into the behaviour of imperfect ring-stiffened cone\u2013cylinder intersections under internal pressure. The paper first presents the results obtained from a careful model experiment, including geometric imperfections, failure behaviour, and the determination of buckling mode and load based on displacement measurements. In addition, results from nonlinear finite element analyses are also presented and compared with the experimental results. The model intersection had the following nominal dimensions (Fig. 1): middle surface radius of the cylinder RZ500 mm, thickness of the cylinder and the cone tcylinderZtconeZ1 mm, apex half angle of the cone aZ408, ring width BZ20 mm and ring thickness TZ1 mm. The length of the cylinder was 200 mm to ensure that the boundary effects at the far end of the cylinder would not affect the behaviour near the cone-to-cylinder joint. The model intersection was closed off by a thick bottom plate at a short distance from the lower end of the cylinder. The model intersection was fabricated using the method of sheet rolling followed by seam welding" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002027_978-1-4020-4535-6_27-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002027_978-1-4020-4535-6_27-Figure4-1.png", "caption": "Figure 4. Flux distribution for no-load condition. (a) Stator teeth b) Rotor pole.", "texts": [ " Lamination core for the magnetic circuit has the property to carry flux mainly in tangential and axial directions. However, due to the interlamination airgap radial flux is reduced considerably. To take into account this effect in the model, additional radial airgaps are introduced in the 3D-FE model. In this manner, flux is forced to flow in the ordinary directions given by the iron permeability and lamination. No-load operation For no-load operation, the only excitation present on the machine is provided for the magnets. Flux density distribution for this operation is depicted in Fig. 4. As expected, there is magnetic activity mostly over the magnet area. Due to the no stator current, armature reaction does not apply flux over the iron section of the airgap. As a result, flux density is negligible in this area. Total flux crossing airgap correspond that impressed by the PMs. On-load operation Under vector control strategy, stator current can be positioned in any location over the airgap respect to the PM flux. According to the required demagnetization effect, current angle (\u03b3 in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003898_17504971311312627-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003898_17504971311312627-Figure6-1.png", "caption": "Figure 6. Two examples of robots used in the Virtual Laboratory for Robotics at . . .University. . .", "texts": [ " Although diagrams showing the time histories of relevant variables represent an efficient way to monitor the system behavior, an animation is still necessary to complete the impression. The detailed image must therefore move according to the mathematical model, replacing the real robot. Let us impose some requirements on the animation: . the robot image must be detailed; . different viewpoints (camera positions) are necessary (including a camera mounted on the robot gripper); . different lighting is desired; and . the visualization must include the robot\u2019s environment. Figure 6 shows the visualized external view of two of the five robots used in the virtual laboratory at ETF. Figure 7 shows the gripper; image (a) is taken from a virtual camera mounted on the robot gripper and image (b) is taken from a virtual camera moving with the gripper but at a constant distance from it. Let us now introduce some additional requirements \u2013 some advanced options taking advantage of the differences between the virtual robot and the real device: . The virtual robot can leave a trace behind it when it moves. The trace may be a line showing the trajectory of the tip point (Figure 6(a)). Alternatively, the trace may represent the whole gripper or the entire robot arm (a kind of \u201cghost effect\u201d). The trace may be permanent in order to show the entire motion, or temporary, being deleted with some delay. . An important feature of the virtual robot is that its graphical representation includes all the inner mechanisms, such as the transmission systems (gears in a gearbox, belts, rack and pinion, spindles, etc.), the motor elements (stator with magnets, rotor), the encoder inner structure, etc. All these inner mechanisms move in complete accordance with the kinematic and dynamic model. The user can \u201copen\u201d the robot arm by removing the covers to expose its internal structures, enabling him to examine the behavior and the role of each of the inner elements. Figure 8 shows the drive and transmission mechanisms used to move the elbow of the jointed robot of Figure 6(a) (note that all these elements are inside the upperarm segment). The figure shows the full drive: the motor, the 3-stage gearbox, a pair of bevel gears to change direction, and a belt to transmit the motion to a distance. On the left side the motor is closed while on the right side the motor has been opened to reveal its inner structure. Figure 9 shows the drive used to move the shoulder joint. It is open, and one can see the magnets on the stator, the rotor armature, and the encoder. Distance learning in engineering 81 D ow nl oa de d by U ni ve rs ity o f N ew ca st le A t 0 8: 03 2 3 M ar ch 2 01 7 (P T ) This advanced visualization combined with the advanced user interface illustrates many of the advantages of the virtual laboratory" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003667_tmag.2010.2044875-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003667_tmag.2010.2044875-Figure7-1.png", "caption": "Fig. 7. Streamline computation and seeding strategy exemplified on a claw pole alternator. (a) Geometry and cutting plane of the claw pole alternator. (b) Lateral view of the claw pole alternator. (c) Front view of the claw pole alternator.", "texts": [ " 6(b) shows a load-dependent current distribution within the red PM of the FE model draws in Fig. 6(a). The streamline distribution given in Fig. 6(c) reflects the vectorial solution adequately. Since some current hotspots are located near the PM surface and edge, the field line endpoints of the corresponding seeding points are opened due to the fact that the integration process crosses the boundary. Even if clawpole alternators are currently manufactured in mass production, the simulation of their electromagnetic characteristics is very time consuming. Fig. 7(a) shows a pole pitch model of a clawpole alternator with about 200 000 elements, which has been analyzed with the FE package iMOOSE [6]. The cutting surface is placed under the claws; the solution within the excitation coils has not been represented to focus on the magnetic active part. The results of the proposed method are given in Fig. 7(b) from lateral view and in Fig. 7(c) from axial view. The stream lines envelope all three stator phases, penetrating the rotor claws at different locations. This allows a quick recognition of the general field distribution, e.g., an interpretable result for checking the FE model and the applied FE settings. Methods for the visualization of magnetic flux lines can provide a visual impression of the vectorial field direction, and therefore, support users to get a quick and qualitative recognition of the flow pattern. In this paper, an algorithm, which detects closed flux lines by extending the integration process by a monitoring routine, is presented" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.41-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.41-1.png", "caption": "Fig. 14.41a,b Principle of operation of a concrete pump with a valve system in the form of the conveying pipe\u2019s swing segment connecting conveying pipe with cylinders during the pumping phase (C-valves \u2013 elephant-type system). (a) valve system in pump; (b) valve in the form of a swing segment", "texts": [ " In the technical literature one can find information about concrete mix pumped to an elevation of 530 m and over a distance of 4000 m (the construction of the Schaeftlarn tunnel). Such pumping distances are achieved when specially designed pumpable concrete mixes are pumped. The design features of these mixes include: the consistency, the cement content, the additives content, and the shape and grading of the aggregates. Another crucial factor is the ambient temperature. A high ambient temperature accelerates concrete mix setting and limits the pumping distance. Concrete pumps (Fig. 14.41) are usually manufactured as self-propelled machines on chassis or trailers to be towed by a tractor. Concrete pumps mounted on self-propelled chassis are usually powered by the vehicle\u2019s engines while those mounted on trailers are driven by a separate engine. Stationary pumps are fitted with skids or driving axles, which are dismantled on the construction site. They are driven by diesel engines or electric motors. The choice of a concrete pump (self-propelled, mounted on a trailer or stationary) is determined by economic factors and the character of the construction project" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000724_physreve.69.011705-Figure14-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000724_physreve.69.011705-Figure14-1.png", "caption": "FIG. 14. Nucleation of a loop of edge dislocations from a screw dislocation, through a helical instability ~see description in the text!.", "texts": [ " in which the separation of two parallel surfaces ~plane and sphere or two cylinders! is slowly decreased. The resulting oscillatory plastic flow can be observed at macroscopic separations and is a consequence of the layered structure. References @9,10# discussed the mechanism responsible for the observed dynamics. Among several models, only the helical instability of screw dislocations was retained. Our observations strongly reinforce this model. First, let us recall the proposed mechanism, which is described in detail in Refs. @13,14#. Figure 14 depicts the nucleation of a loop from a screw dislocation under a helical instability. Figure 14~a! represents a straight screw dislocation which joins the boundaries of an unstrained sample. Under compression, the screw dislocation is unstable and adopts a helical shape @Fig. 14~b!#. The origin of this instability comes from the fact that one layer has disappeared inside the cylinder formed by the helix, which therefore decreases the stress. The helical dislocation has acquired a mixed character ~screw and edge! and can transform into a straight screw dislocation of the same Burgers vector plus a loop edge dislocation with a kink @Fig. 14~c!#. The loop then grows outward to relax the remaining stress. The main characteristic of this model is that the edge loop should appear in the middle of the sample, around a screw dislocation, and should leave it unchanged, ready to nucleate another loop. Our observations are in complete agreement with these points and strengthen the validity of the helical instability model. Our results raise the problem of the microscopic origin of the pinning and how the dislocations intersect. We will now discuss this point further, which was already approached in Refs" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.69-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.69-1.png", "caption": "Fig. 14.69 Tube\u2013coupler scaffold joint", "texts": [ ", during working shift Permanent scaffolds used for prolonged periods without dismantling 7 Material from which scaffold load-bearing Wooden scaffolds elements are made Aluminum scaffolds Steel scaffolds 8 Technical\u2013organizational Scaffolds in typical version and formal-legal aspects Individually designed scaffolds fold according to a blueprint ascertain the positions of all the elements which determine the dimensions of the structural grid and the verticality of the uprights. The basic components of the tube\u2013coupler scaffold are shown in Fig. 14.68. In tube\u2013coupler scaffolds, such elements as standards, transons, and bracings are joined eccentrically by right-angle or swivel couplers, as illustrated in Fig. 14.69. A characteristic feature of scaffolds made of prefabricated elements (system scaffolds) is that their dimensions (or some of their dimensions) are determined by the dimensions of their components. All frame and modular scaffold systems belong to this class. A general view of system scaffold constructions is shown in Fig. 14.70. Modular scaffolds and frame scaffolds are shown on the left and right, respectively. In the frame scaffold, the vertical structure is made up of prefabricated flat frames" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001622_b978-044451958-0.50002-1-Figure1.5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001622_b978-044451958-0.50002-1-Figure1.5-1.png", "caption": "Figure 1.5 A galvanic electrochemical cell. (for colour version: see colour section at the end of the book).", "texts": [ "2) Electrolytic processes are used by the pharmaceutical industry (electrochemical synthesis) and heavily by the metal-refining industry where they represent the only currently utilized means of production of aluminum from alumina ore. These industrial electrolytic processes, by some estimates, use a significant fraction of the electricity produced worldwide today (5). Cu/H //Cu /Pt+ +2 Cu Cu H SO Pt4 2/ , , /2+ + \u2212 20 1. Fundamentals The schematic depiction of Volta\u2019s battery (Section 1.1), which is composed of stacked copper and zinc disks separated by paper soaked in an acidic solution (Figure 1.5), could be represented as (1.6.3) This is an example of a primary (non-rechargeable) cell. Other types of important galvanic cells include secondary (rechargeable) cells and fuel cells. It should be noted that when a secondary galvanic cell is being recharged, it becomes an electrolytic cell, because a potential is being applied in order to reverse the direction of the spontaneous electrochemical reaction. While it may suffice to describe the electrodes of a galvanic cell as simply an anode or cathode, electrodes in electrolytic cells are called on to perform more specialized roles" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000975_6.2004-5308-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000975_6.2004-5308-Figure2-1.png", "caption": "Figure 2. A N-body Closed-Hub-And-Spoke Tethered Formation.", "texts": [ " The two types of formations under consideration are hereby named, for ease of discussion, hub-andspoke configuration and closed-hub-and-spoke configuration. The hub-and-spoke configuration is a system consisting of a central (or parent) body of mass mP in the hub and N tethers of length li connecting N bodies of mass mi to the hub (Fig. 1). The closed-hub-and-spoke configuration is the same system but to which are added N external tethers connecting the N peripheral bodies in pairs in a ring-type setting (Fig. 2 of 23 American Institute of Aeronautics and Astronautics 2) , i.e m1 is joined to m2, mN is joined to m1, etc. The tethers are considered straight but elastic and are subject to structural damping, and their mass is assumed to be negligible. The dimensions of the end-bodies are much smaller than the length of the tethers so that the former can be approximated as point masses. The free dynamics of both types of tethered satellite formations is studied assuming the center of mass of the system moves in a Keplerian orbit at LEO" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003951_j.apenergy.2012.03.026-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003951_j.apenergy.2012.03.026-Figure3-1.png", "caption": "Fig. 3. Experiment framework of adding flow channel in the anode.", "texts": [ " coli culture solution, and by quantitative method fetch required liquid volume and put into flask with culture solution, and leave at least 2/3 space for bacteria stirring space, finally put into 37 C incubator and conduct cultivation with 200 rpm for 18 h. In order to study the effect of biometric mixer and biometric flow channel, two experiments were set up, as described as follows: (1) Biometric flow channel experiment. Based on the common continuous MFC, the biometric flow channel was added into the anode chamber, and the schematic diagram was shown in Fig. 3. When the nutrient and bacteria flow through the MFC, the existing block could change the movement of fluid and cause a disturbance. Then the disturbance could improve the mixing of nutrient and bacteria. (2) Biometric mixer experiment. As shown in Fig. 4, the biometric mixer was added ahead of the MFC. Before the nutrient and bacteria affluxes to the MFC, they would mix in the biometric mixer first. It can make the mixing more uniformly. The related experiment operating parameters were as following: In the anode, the E", " The temperature is 27 C, and the density and viscosity coefficient is 997 kg m 3 and 8.91 10 4 kg/ms separately. The physical property, including density, molecular weight and viscosity coefficient, is all the same without changing. The simulation was conducted with the analysis settings below: (1) Flow channel simulation Total grid amount is N = 7.3 106, Solver is set as AMG (Algebraic Multi-Grid), and the convergent value as 10 4. The fluid flows into the MFC through the upper inlet, with the Reynolds number as Re = 3.73, and flow out from the lower part (as shown in Fig. 3). (2) Flow channel combining biometric mixer simulation The total grid number is N = 4.7 105, Solver is set as AMG, and the convergent value as 10 4. As shown in Fig. 4, the fluid flows into MFC through the upper three inlets with the Reynolds number ratio Rer = 1, and out from the lower outlet. In the fluid mixing simulation, water \u2018\u2018A\u2019\u2019 is supposed to flow through the major inlet in the middle, while water \u2018\u2018B\u2019\u2019 is supposed to flow through the other two inlets. In addition, in the experiment, gray scale technique was employed to quantize the fluid mixing effect through the color changes caused by different fluid density" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003423_1.4002259-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003423_1.4002259-Figure3-1.png", "caption": "Fig. 3 General material stress-strain curves of tube and tubesheet", "texts": [ " Analytical solutions of the residual contact pressure and/or the required expansion pressure for the hydraulically expanded tube-to-tubesheet joint will be described in Sec. 3. 3 Analytical Solutions 3.1 Material Model. Corresponding to the process of tube-totubesheet expansion, see Fig. 2, the elastic-plastic deformation will take place in the tube and tubesheet. For describing the behaviors of tube and tubesheet materials in the process, a general material model proposed by present author 12 is adopted. The material stress-strain curves of tube and tubesheet is shown in Fig. 3 and expressed by Eqs. 2 and 4 for the loading stage, and Eqs. 3 and 5 for the unloading stage, respectively. For tube material in the loading stage: i = Et i i st A1 + A2 i B1 i st 2 For tube material in the unloading stage, i = Et i 3 For tubesheet material in the loading stage, i = Es i i ss A3 + A4 i B2 i ss 4 For tubesheet material in the unloading stage, i = Es i 5 Figure 1 shows the geometrical character of the simplified tubeto-tubesheet joint model. The material model described by the above equations is adopted in the analytical solution derivation", " In equilibrium equation, d r dr + r \u2212 r = 0 11 n geometrical equation, r = du dr 12 = u r 13 n Hencky deformation theory constitutive equations, r = 3 i 2 i r \u2212 m 14 = 3 i 2 i \u2212 m 15 z = 3 i 2 i z \u2212 m 16 m = 1 3 r + + z 17 3.2.3 Basic Derived Equations. Substituting Eqs. 9 and 17 nto Eq. 16 gives z = 1 2 + r 18 ubstituting Eqs. 17 and 18 into Eqs. 14 and 15 respecively, gives = 3 i 4 i \u2212 r 19 r = 3 i 4 i r \u2212 20 ubstituting Eq. 18 into Eq. 6 gives i = 3 2 \u2212 r 21 3.3 Tube in Loading Stage 3.3.1 Tube Elastic Deformation Phase. The loading stress nalysis is based on the Cartesian coordinate system O , shown n Fig. 3. The radii of elastic-plastic zones in the tube wall are hown in Fig. 4. Before the tube contact with the tubesheet, the nternal pressure is pi, and the external pressure is zero. The inner and outer radii and the elastic-plastic interface radius f tube are shown in Fig. 4. Tube stress and deformation are given y Lam\u00e9 equations when tube deformed elastically. ournal of Pressure Vessel Technology om: http://pressurevesseltech.asmedigitalcollection.asme.org/ on 01/27/20 r = r2 \u2212 ro 2 ri 2 ro 2 \u2212 ri 2 r2 pi = r2 + ro 2 ri 2 ro 2 \u2212 ri 2 r2 pi u = k1 1 r pi 22 k1 = 3ro 2ri 2 2Et ro 2 \u2212 ri 2 23 i = 3ri 2ro 2 ro 2 \u2212 ri 2 r2 pi 24 3", "org/ on 01/27/20 pi = ss Ro 2 \u2212 Rc 2 3Ro 2 + 2 3 A3 ln Rc Ri + 2 3 A1 ln ro ri \u2212 1 2B2 2 3 B2+1 A4 ro co + 3 ssRc 2 2EsRi B2 Rc \u22122B2 \u2212 Ri \u22122B2 \u2212 1 2B1 2 3 B1+1 A2 ro co + 3 ssRc 2 2EsRi B1 ro \u22122B1 \u2212 ri \u22122B1 58 When the outer side of tubesheet starts to yield, replacing Rc with Ro in Eq. 58 , the collapse pressure psyo is determined by the following equation: psyo = 2 3 A3 ln Ro Ri + 2 3 A1 ln ro ri \u2212 1 2B2 2 3 B2+1 A4 ro co + 3 ssRo 2 2EsRi B2 Ro \u22122B2 \u2212 Ri \u22122B2 \u2212 1 2B1 2 3 B1+1 A2 ro co + 3 ssRo 2 2EsRi B1 ro \u22122B1 \u2212 ri \u22122B1 59 3.5 Unloading Stage. Unloading stress analysis is based on the Cartesian coordinate system B , shown in Fig. 3. The radii of elastic-plastic zones in the tube and tubesheet wall are shown in Fig. 5. The procedure for unloading stress analysis is analogous to that for elastic loading stress analysis. As the expansion pressure is totally removed, the pressure at the inner surface of the tube is zero and the pressure at the contact surface is pc . The change in internal pressure of the tube is equal to \u2212pi, the pressure at the contact surface is from pc to pc . The unloading is assumed perfectly elastic. In this case, the unloading stresses in the whole tube wall ri r ro and tubesheet wall Ri r Ro follow the Lame\u0307 equations and the changes of radial displacement at the contact surface of the tube and the tubesheet are the same for tube and tubesheet" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003086_tsmca.2010.2076405-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003086_tsmca.2010.2076405-Figure2-1.png", "caption": "Fig. 2. Transformation of coordinate systems.", "texts": [ " Section III shows how to compute the independent components of the link lengths and analyze the relationships between the independent components and the positional variables. The numerical algorithm for obtaining the value of the positional variables is presented in Section IV. In Section V, simulation experiments are carried out, and the experimental results are compared with those provided by other published algorithms. The conclusions are drawn in Section VI. For the purpose of analysis, the transformation of a Cartesian coordinate system as shown in Fig. 2 is used. Given an original coordinate system Oxyz and a new coordinate system Ox\u2032y\u2032z\u2032, they share the same coordinate origin O. For any point M , its coordinates are (x, y, z) in the coordinate system Oxyz and (x\u2032, y\u2032, z\u2032) in the coordinate system Ox\u2032y\u2032z\u2032. Thus, they satisfy the following: (x, y, z)T = R(x\u2032, y\u2032, z\u2032)T (1) where R is the transformation matrix from the coordinate system Oxyz to Ox\u2032y\u2032z\u2032. The matrix R may be represented uniquely by three angles (\u03b8, \u03c8, \u03d5) = (\u03b8, \u03b1+ 3\u03c0/2, \u03b1+ \u03b2), where (\u03b1, \u03b8, \u03b2) are the Euler\u2019s angles by defining the relative rotations z\u2212x\u2212z as shown in Fig. 2, as follows: 1) Rotate the system Oxyz about the z-axis by \u03b1, and the x-axis now lies on the line ON ; 2) rotate the system Oxyz again about the new rotated x-axis by \u03b8, and the z-axis is now in its final orientation; and 3) rotate the system Oxyz a third time about the new z-axis by \u03b2. R can be easily derived, and it is given by (2), as shown at the bottom of the page. For the convenience of expression, \u03b8, \u03c8, and \u03d5 are called transformation angles from the coordinate system Oxyz to Ox\u2032y\u2032z\u2032. For the convenience of forward kinematics analysis, three Cartesian coordinate systems have been set up for the Stewart platform" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.97-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.97-1.png", "caption": "Fig. 14.97 Portable crane mounted on building\u2019s roof", "texts": [ " The numbers above the curves specify the allowable hoisting capacity for a given jib and hoisting height A range of portable machines, based on winches and other accessories, for handling materials and transferring light equipment on construction sites has been developed. This machinery includes: \u2022 Scaffold cranes mounted on scaffolds (Fig. 14.92 items 5 and 6, and Fig. 14.95)\u2022 Portable cranes (Fig. 14.92 item 4, and Figs. 14.96, 14.98) fixed to steel supports installed between floors, in window openings or on the roof (Fig. 14.97, basic parameters are shown in Table 14.17)\u2022 Gantries mounted on the roof (Fig. 14.92 item 3), in an opening in the building\u2019s elevation (Fig. 14.92 item 2) or on a scaffold (Fig. 14.89 item 1) Part B 1 4 .6 The main component of the above machines is a universal winch that can work in tandem with various accessories. Figure 14.92 shows the use of scaffold cranes, portable cranes, and small-capacity gantries during building erection. Machinery of this type is intended for lifting and transferring loads of up to 200 kg to a height of 80 m" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002090_9780470264003-Figure12.6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002090_9780470264003-Figure12.6-1.png", "caption": "Figure 12.6 Datum design exploded view.", "texts": [], "surrounding_texts": [ "design and decide on the theoretical possible minimum number of parts. We need to simplify by achieving the minimum number of parts as follows: \u2022 The motor and the sensor comprise standard purchased subassembly. Thus, no further analysis is required. \u2022 The base is assembled into a fixture, and since there is no other part to assemble to, it is a \u201ccritical\u201d part. \u2022 The two bushings don\u2019t satisfy the criteria in Section 12.4.1. Theoretically, they can be assembled to the base or can be manufactured of the same material as the end plate and combined with it. \u2022 The setscrew, the four cover screws, and the end-plate screws are theoretically unnecessary. An integral fastening arrangement is possible most of the time. \u2022 The two standoffs can be assembled to the base (do not meet the criteria). \u2022 The end plate is a critical part for accessibility. 278 MEDICAL DEVICE DESIGN FOR X" ] }, { "image_filename": "designv11_20_0003703_cctc.201000130-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003703_cctc.201000130-Figure4-1.png", "caption": "Figure 4. Immobilization of CalB on Accurel MP1001 at different enzyme loadings. Immobilized CalB on oxygen plasma-modified support (dark gray), immobilized CalB on untreated support (light gray).", "texts": [ " Treatment of the support with ammonia plasma showed no positive effect on enzyme immobilization (data not shown). During these experiments, we learned that Accurel EP100 and MP1000 are no longer commercially available, thus MP1001 was used in further studies. To determine the loading capacity of MP1001, saturation experiments were performed using different amounts of enzymes. Thus, for PestE, an up to 14-fold increase in immobilization yield compared to untreated support was observed (Figure 3). For CalB, a similar effect was found, resulting in a fivefold increase of immobilized activity (Figure 4). The molecular reasons for the increase in enzyme immobilization is currently under investigation. Transesterification activity of the immobilized enzymes In order to prove that PestE immobilized on the plasma-modified supports Accurel EP100, MP1000, and MP1001 is active in transesterifications, the kinetic resolution of rac-a-phenylethanol with vinyl acetate was performed (Table 3). From these experiments, the immobilized PestE gave faster conversion than the free enzyme and there was also a slight increase in enantioselectivity" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003908_978-94-007-4201-7_1-Figure1.8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003908_978-94-007-4201-7_1-Figure1.8-1.png", "caption": "Fig. 1.8 Twist motion of a body", "texts": [ " For this reason,S is called a free vector and the motion can be quantified by taking a scalar multiple of the free vector v\u00f00; S\u00de or \u00f00; n\u00de. Some readers may prefer to consider instantaneous translation to be an instantaneous rotation about an axis that is orthogonal to S and that lies in the plane at infinity. The Pl\u20acucker coordinates of this axis are \u00f00; S\u00de, and the instantaneous rotation about this axis can be expressed as the scalar multiple v\u00f00; S\u00de. When the motion of one body 2 relative to another includes rotation about the axis S1 and translation in the direction S1, the situation is more complex, as shown in Fig. 1.8. The body rotates about the axis S1, with the instantaneous wrench o1\u00f0S1 ; S01\u00de, where \u00f0S1; S01\u00de is unit screw. The body also translates with screw v2\u00f00;S1\u00de along the axisS1 at the same time. The absolute motion of the body is the sum of the two parts. That is, oi $i \u00bc \u00f0o1S1;o1S01\u00de \u00fe \u00f00; v1S1\u00de \u00bc \u00f0o1S1;o1S01 \u00fe ho1S1\u00de \u00bc o1 S1; S 0 1 ; (1.52) or oi $i \u00bc o1 S1; S 0 1 \u00bc \u00f0v1; v 0\u00de; (1.53) where v1 is the angular velocity of the body and v0 is the velocity of a point in the body coincident with the origin" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002086_j.mechmachtheory.2007.04.006-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002086_j.mechmachtheory.2007.04.006-Figure4-1.png", "caption": "Fig. 4. Small change of position of P1 due to a small twist of body A.", "texts": [ " (6), oS oa and oS ob can be explicitly written by oS oa \u00bc sin b sin a sin b cos a 0 2 64 3 75; \u00f010\u00de oS ob \u00bc cos b cos a cos b sin a sin b 2 64 3 75: \u00f011\u00de Since oS oa is not a unit vector, a unit vector oS oa 0 is introduced as oS oa 0 \u00bc sin a cos a 0 2 64 3 75; \u00f012\u00de oS oa \u00bc sin b oS0 oa : \u00f013\u00de Hence Eq. (7) can be rewritten as dw \u00bc k dl$\u00fe k 1 l0 l o$0 oa l sin bda\u00fe o$ ob ldb ; \u00f014\u00de where o$0 oa \u00bc oS0 oa ErE P0 oS0 oa \" # : \u00f015\u00de It is important to note that o$0 oa and o$ ob are the unitized Plu\u0308cker coordinates of the lines perpendicular to S and go through the pivot point P0. In Eq. (14) dl, l sinbda, and ldb can be considered as the change of the spring length and the changes of the direction of the spring (see Fig. 4). These values correspond to the projections of the variation of position P1, EdrA P1, into S; oS0 oa , and oS ob, respectively. Thus EdrA P1 \u00bc \u00f0EdrA P1 S\u00deS\u00fe EdrA P1 oS0 oa oS0 oa \u00fe EdrA P1 oS ob oS ob \u00bc dlS\u00fe l sin bda oS0 oa \u00fe ldb oS ob : \u00f016\u00de From the twist equation, the variation of position P1 can be written as EdrA P1 \u00bc EdrA 0 \u00fe EduA ErA P1; \u00f017\u00de where EdrA 0 is the differential of the position of point O in body A which is coincident with the origin of the inertial frame E measured with respect to the inertial frame", " From the twist equation, the variation of position of point P2 in body B with respect to body A can be expressed as AdrP2 B \u00bc AdrB 0 \u00fe AduB ArB P2; \u00f033\u00de where ArB P2 is the position of P2, which is embedded in body B, measured with respect to a coordinate system embedded in body A which at this instant is coincident and aligned with the reference system attached to ground. It can also be decomposed into three perpendicular vectors along S; AoS0 oa , and AoS ob which are defined in a similar way as Eqs. (6), (12), and (11). These three vectors correspond to the change of the spring length dl and the directional changes of the spring such as l sinbda and ldb in terms of body A in a way that is analogous to that shown in Fig. 4. Thus the variation of position of point P2 in body B in terms of body A can be written as AdrB P2 \u00bc \u00f0AdrB P2 S\u00deS\u00fe AdrB P2 AoS0 oa AoS0 oa \u00fe AdrB P2 AoS ob AoS ob \u00bc dl S\u00fe l sin bda AoS0 oa \u00fe ldb AoS ob : \u00f034\u00de From Eqs. (33) and (34), dl in Eq. (32) can be obtained as dl \u00bc AdrB P2 S \u00bc AdrB 0 S\u00fe AduB ArB P2 S \u00bc AdrB 0 S\u00fe AduB ArB P2 S \u00bc $T AdDB: \u00f035\u00de In the same way, l sinbda and ldb can be expressed as l sin bda \u00bc AdrB P2 AoS0 oa \u00bc Ao$00T oa AdDB; \u00f036\u00de ldb \u00bc AdrB P2 AoS ob \u00bc Ao$00T ob AdDB; \u00f037\u00de where Ao$ oa 00 \u00bc AoS oa 0 ArB P2 AoS oa 0 2 4 3 5; \u00f038\u00de Ao$00 ob \u00bc AoS ob ArB P2 AoS ob 2 4 3 5: \u00f039\u00de Now in Eq", " (42) can be written as AdS ErA P1 AdS \" # \u00bc AoS oa da\u00fe AoS ob db ErA P1 AoS oa da\u00fe AoS ob db 2 64 3 75 \u00bc AoS oa da ErA P1 AoS oa da 2 4 3 5\u00fe AoS ob db ErA P1 AoS ob db 2 4 3 5 \u00bc Ao$0 oa 1 l l sin bda\u00fe Ao$ ob 1 l ldb \u00bc 1 l Ao$0 oa Ao$00 oa T \u00fe Ao$ ob Ao$00 ob T ! AdDB: \u00f043\u00de As to the second screw in Eq. (42), EduA \u00b7 S can be decomposed into three perpendicular vectors along S, AoS0 oa , and AoS ob , respectively, as EduA S \u00bc \u00f0EduA S\u00de S n o S\u00fe \u00f0EduA S\u00de AoS0 oa AoS0 oa \u00fe \u00f0EduA S\u00de AoS ob AoS ob : \u00f044\u00de From the fact that S, AoS0 oa , and AoS ob are unit vectors and perpendicular to each other (see Fig. 4), each dot product of Eq. (44) can be expressed as \u00f0EduA S\u00de S \u00bc 0; \u00f045\u00de \u00f0EduA S\u00de AoS0 oa \u00bc EduA S AoS0 oa \u00bc EduA AoS ob \u00bc 0 AoS ob 2 4 3 5 T EdrA 0 EduA \" # \u00bc 0 AoS ob 2 4 3 5 T EdDA; \u00f046\u00de \u00f0EduA S\u00de AoS ob \u00bc EduA S AoS ob \u00bc EduA AoS0 oa \u00bc 0 AoS0 oa 2 4 3 5 T EdrA 0 EduA \" # \u00bc 0 AoS0 oa 2 4 3 5 T EdDA; \u00f047\u00de where 0 \u00bc 0 0 0\u00bd T . Hence, EduA \u00b7 S can be rewritten as EduA S \u00bc AoS0 oa 0 AoS ob \" #T EdDA \u00fe AoS ob 0 AoS0 oa \" #T EdDA \u00f048\u00de and the second screw in Eq. (42) can be expressed as EduA S ErA P1 \u00f0EduA S\u00de \" # \u00bc AoS0 oa 0 AoS ob \" #T EdDA \u00fe AoS ob 0 AoS0 oa \" #T EdDA ErA P1 AoS0 oa 0 AoS ob \" #T EdDA \u00fe AoS ob 0 AoS0 oa \" #T EdDA 8< : 9= ; 2 66666664 3 77777775 \u00bc AoS0 oa ErA P1 AoS0 oa 2 4 3 5 0 AoS ob \" #T EdDA \u00fe AoS ob ErA P1 AoS ob 2 4 3 5 0 AoS oa 0 \" #T EdDA \u00bc Ao$ oa 00 0 AoS ob \" #T \u00fe Ao$ ob 00 0 AoS oa 0 \" #T 0 @ 1 AEdDA: \u00f049\u00de As to the third screw in Eq", " (50) can be written as EdrA P1 AoS oa 0 \u00bc EdrA 0 AoS oa 0 \u00fe EduA rA P1 AoS oa 0 \u00bc EdrA 0 AoS oa 0 \u00fe EduA rA P1 AoS oa 0 \u00bc Ao$ oa 00T EdDA; \u00f052\u00de EdrA P1 AoS ob \u00bc EdrA 0 AoS ob \u00fe EduA rA P1 AoS ob \u00bc EdrA 0 AoS ob \u00fe EduA rA P1 AoS ob \u00bc Ao$ ob 00T EdDA: \u00f053\u00de Finally, EdrA P1 S of the third screw in Eq. (42) can be expressed as EdrA P1 S \u00bc \u00f0$T EdDA\u00deS\u00fe Ao$ oa 00T EdDA ! AoS oa 0 \u00fe Ao$ ob 00T EdDA ! AoS ob ( ) S \u00bc Ao$ oa 00T EdDA ! AoS oa 0 S\u00fe Ao$ ob 00T EdDA ! AoS ob S \u00bc Ao$ oa 00T EdDA ! AoS ob Ao$ ob 00T EdDA ! AoS oa 0 ; \u00f054\u00de since S, AoS oa 0 , and AoS ob are unit vectors and perpendicular to each other (see Fig. 4). Substituting Eq. (54) for EdrA P1 S of the third screw in Eq. (42) yields 0 EdrA P1 S \" # \u00bc 0 Ao$ oa 00T EdDA AoS ob Ao$ ob 00T EdDA AoS oa 0 2 64 3 75 \u00bc 0 AoS ob 2 4 3 5 Ao$ oa 00T EdDA ! 0 AoS oa 0 2 4 3 5 Ao$ ob 00T EdDA ! \u00bc 0 AoS ob 2 4 3 5Ao$ oa 00T 0 AoS oa 0 2 4 3 5Ao$ ob 00T 0 @ 1 AEdDA: \u00f055\u00de By replacing dl and Ed$ in Eq. (32) with Eqs. (35), (43), (49), and (55) and sorting it into the twists, the derivative of the spring wrench can be rewritten as Edw \u00bc k dl$\u00fe k\u00f0l l0\u00deEd$ \u00bc \u00bdKF AdDB \u00fe \u00bdKM E dDA; \u00f056\u00de where \u00bdKF \u00bc k$$T \u00fe k 1 l0 l Ao$ oa 0Ao$ oa 00T \u00fe Ao$ ob Ao$ ob 00T " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000610_00423110600870055-Figure13-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000610_00423110600870055-Figure13-1.png", "caption": "Figure 13. Free differential (a); locking action of the differential (b); and the three-wheel reference model (c) in a \u03bc-split condition.", "texts": [ " Power-off starting and ending conditions are detected by observing the driver\u2019s input, with particular attention to the torque demand. The system implemented foresees some subsystems dedicated to particular driving conditions such as start-on \u03bc-split. In this condition, one driven wheel is on a low-adherence (low \u03bc) surface, such as ice or mud, and it is not able to transmit the driving torque to the ground. A free differential would allow the wheel on low \u03bc to spin, dissipating the whole driving power although the car does not move (figure 13(a)). A locking action of the differential transfers torque to the wheel on high \u03bc, improving traction, but, of course, it generates an oversteering moment on the vehicle, because of the non-uniform distribution of longitudinal contact forces (figure 13(b)) which the driver has to contrast acting on the steering wheel. A good control system has to transfer to the high-\u03bc wheel the right amount of torque in order to reach a good compromise between longitudinal performance and driving feeling. The algorithm developed to evaluate the reference locking torque Tf,ref , in the vehicle start condition, adopts a feed-back component based on a PI controller. This regulator gets as input the error between the actual difference of wheel speeds VX,ist and the threshold one VX,\u03bc-split that is calculated in feed-forward using a kinematical three-wheel vehicle model (figure 13(c)). As this driving condition is characterized by very low side-slip angles, low speed, and low lateral acceleration, a kinematical approach is accurate enough, and it can be assumed as the reference. The model is used to calculate the maximum wheels speed difference VX,\u03bc-split that the system can tolerate before actuating some locking torque. VX,\u03bc-split is mainly a function of the steer angle and the vehicle longitudinal speed according to equation (9), where VX is the vehicle longitudinal speed coming from a CAN bus, \u03b4 the steer angle at the wheel, b the car base, and t the car rear track" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002261_j.mechmachtheory.2009.09.006-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002261_j.mechmachtheory.2009.09.006-Figure8-1.png", "caption": "Fig. 8. 3D-FEM mesh dividing pattern.", "texts": [ " Seven points along the tooth width and five points along the tooth height are sufficient to obtain correct results. Finally, a total of thirty five points is necessarily to determine the bending definition matrix. Table 4 Maximal and mean errors for the calculation of displacements for the 2nd comparison. Maximal error 7.3% Mean error 1.3% Fig. 12. Load Sharing. All these points have been imposed as nodes in the meshing. The size of the mesh is small on the teeth having points, in order to have correct results in term of displacement computations. The global mesh can be coarser, as shown in Fig. 8 Fig. 9 presents the different points taken into account in the comparison. The figure corresponds to the points on the spiral bevel gear. All the different points have the same parameter L20. Only parameter L10 differs from one point to another. The principle is the same for studying the pinion. Applying a unit load successively on different points, the displacements are first calculated by FEM and then calculated with the functions. The different results are presented in Fig. 10. This figure shows the displacements of the different points for several load cases" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000495_2006-01-0426-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000495_2006-01-0426-Figure1-1.png", "caption": "Figure 1 Grasshopper linkage and connecting rod", "texts": [], "surrounding_texts": [ "Piston assembly friction was measured initially at 16 different load/speed conditions with oil temperature controlled to 90\u00b0C and using 5W30 grade oil. \u2022 Engine speed - 1000, 1500, 2000 and 2500 rpm \u2022 Engine load \u2013 motored, quarter, half and full The oil was then changed to 15W40 and 8 further tests were performed. \u2022 1000 rpm \u2013 quarter, half and full load \u2022 2000 rpm \u2013 quarter and full load \u2022 2500 rpm \u2013 quarter, half and full load" ] }, { "image_filename": "designv11_20_0001920_1.2988480-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001920_1.2988480-Figure1-1.png", "caption": "Fig. 1 Coordinate systems O-xyz, O-xgygzg, and O-XYZ", "texts": [ " Therefore, the ratio of roll i between the quasiomplementary crown gear and the workpiece is cos f /sin 0. his quasi-complementary crown gear generates the tooth surface f the straight bevel gear with a modified profile and a depthwise ooth taper taking the theoretical generation method and the posiion of tooth bearing into consideration, respectively. Therefore, a ood tooth bearing is expected to get at an indicated position of he tooth surface without trial and error. 2.2 Mathematical Expression of Tooth Surface of Straight evel Gear. Figure 1 shows three coordinate systems: O-xyz, -xgygzg, and O-XYZ. O-xyz and O-xgygzg are attached to the uasi-complementary crown gear and the generated gear, respecively. O-XYZ is fixed in space. Origin O is the machine center. Y plane is the cradle plane. z is the cradle axis, which is equivaent to the axis of the quasi-complementary crown gear. zg is the xis of the generated gear and is inclined at an angle r to the Y xis. yg axis is omitted. O-xyz coincides with O-XYZ when the otation angle of the quasi-complementary crown gear is zero", " In addition, a complicated grid model is not needed beause the real values related to gear accuracy can be detected sing the coordinates of arbitrary points. 3.1 Outline of Measurement Method. The aim of the meaurement of the straight bevel gear accuracy is fundamentally to ompare the measured real tooth surface with the theoretical one. he coordinate measurement of the real gear tooth surface proides the information about some factors related to the gear accuacy mentioned in the previous section. We consider the pressure ngle , tooth angle error as shown in Fig. 2, workpiece setting ngle r as shown in Fig. 1, and apex to back l as shown in Fig. 3 s the factors related to the gear accuracy in this paper. These actors are invariable because these factors do not change during he end milling process. For measurement, the straight bevel gear is set arbitrary on a MM. The positions of the gear axis and the datum plane must be etermined by measurement independent of the tooth surface easurement because the gear is set arbitrarily. We can make the rigin O and gear axis zg in the coordinate system O-xgygzg atached to the gear coincide with the origin Om and axis zm in the oordinate system Om-xmymzm of the CMM, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003074_j.engfailanal.2010.03.009-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003074_j.engfailanal.2010.03.009-Figure10-1.png", "caption": "Fig. 10. (a) A shaft coupling in place and a cut-away view showing bushings and (b) worn bushings and wear debris.", "texts": [ " The impeller and motor drive shaft are connected via a pin and coupling with elastomeric bushings, used to minimize effects of mechanical misalignment and to dampen torsional oscillations. The coupling has a stiffness that is variable depending on transmitted torque. Examination of other components revealed damage to the shaft and coupling bushings. Damaged elastomeric elements are signs of high dynamic torque in the coupling. The shaft exhibited a crack emanating from a keyway oriented at approximately 45 from the shaft surface (Fig. 9), which is typical of failures due to high torsional vibration. Images of a coupling in place and a cut away to reveal bushings are shown in Fig. 10a. Coupling bushings were heavily worn after service (see Fig. 10b) and were directly observed to be losing material as fine shavings during operation. These two pieces of evidence indicated that there was a significant source of torsional stress that had been unanticipated during design and was previously undetected during operation. Vibration monitoring is common on rotating machinery to detect mechanical imbalance and bearing condition. In the present case, the transverse and axial components of vibration were monitored continuously and were within the \u2018normal\u2019 range at the time of the failure" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000319_095440605x8478-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000319_095440605x8478-Figure3-1.png", "caption": "Fig. 3 Two rigid bodies connected by a flexure hinge", "texts": [ " 2, the force vector consisting of force and moment components applied to the point oi by the small displacements Dri and Dui is Qi \u00bc \u00bdFij j Tij \u00fe dij Fij T (20) Substituting equation (19) into equation (20) results in Qi \u00bc ~DijK ji ~D T ij Dqi (21) In addition, the force vector applied to the point oji on the fixed frame is Q ji \u00bc \u00bdFij j Tij \u00fe h ji Fij T (22) wherehji is the distance vector from the point oji to the point oij and is given by h ji \u00bc \u00bdhxji hy ji hz ji T (23) Substituting equation (19) into equation (22) results in Q ji \u00bc ~H jiKij ~D T ij Dqi (24) where ~H ji \u00bc I 0 H ji I (25) and H ji \u00bc 0 hzji hy ji hz ji 0 hxji hyji hx ji 0 2 4 3 5 (26) Whena rigid body i is connected to a rigid body jwith a flexure hinge, as shown in Fig. 3, the force vector (consisting of force and moment components) applied to the point oi by the small displacements Dqi and Dqj is Qi \u00bc ~Dij K ji ~D T ij Dqi ~HijKij ~D T ji Dqj (27) Consider a rigid body i connected to multiple rigid bodies with multiple flexure hinges, as shown in Fig. 4. The rigid body i is connected to nb rigid bodies with flexure hinges. In addition, the rigid body i and Proc. IMechE Vol. 219 Part C: J. Mechanical Engineering Science C19203 # IMechE 2005 at NATIONAL UNIV SINGAPORE on June 27, 2015pic" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001233_te.2005.858389-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001233_te.2005.858389-Figure3-1.png", "caption": "Fig. 3. UVM CricketSat wireless temperature sensor (Revision F, 2005).", "texts": [ ") but not beyond, because of loss of communication, and later because of radiation heating of the black epoxy structure in the upper atmosphere. Two main objectives were considered in adapting the CricketSat design for educational programs at UVM. First, the circuit needed to be easy to fabricate and test, robust to common fabrication errors, and low cost 10 . Since those assembling the printed circuit boards (PCBs) could be novices, the board also needed to withstand less than refined soldering techniques. As such, the PCB (Fig. 3) was designed with detailed labeling, strain relief for battery and antenna leads, thermal relief, and a solder mask. In addition a fully illustrated assembly procedure was developed to ensure ease of fabrication and explanation of operation. Second, the circuit needed to be adaptable to a variety of applications beyond temperature sensing. For example, humidity, pressure, and solar radiation are all parameters of interest in the ballooning programs. Thus, an enlarged prototype development area was included on PCB for sensors and/or other components" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.55-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.55-1.png", "caption": "Fig. 14.55 Two-disk floating machine equipped with blades", "texts": [ " Modern single-disk floating machines have the following features: \u2022 A long handle, enabling access to floated surfaces with no need to walk on them (during transport of the floating machine the shaft is folded).\u2022 The angle of inclination of the blades can be adjusted from the handle.\u2022 The machine is equipped with a safety cutout switch (the so-called dead-man\u2019s grip), which automatically stops it once the operator\u2019s grip on the handle is released.\u2022 Electric protection against switching the opposite direction of rotation.\u2022 Road wheels for short-distance transport.\u2022 The blades and the floating disks are made of highquality materials to ensure long lifetime. Two-disk floating machines (Fig. 14.55) are used for floating large surfaces. Part B 1 4 .3 For these machines the width of the floated strip in one pass ranges from 1700 to 2400 mm. The floating machine\u2019s control system consists of push-buttons, two joysticks for controlling the running direction, and two knobs for setting the angles of inclination of the blades. The operator controls the floating machine while sitting in a centrally situated seat whose symmetry axis coincides with the machine\u2019s lateral axis. There are also tandem systems in which the symmetry axis of the operator\u2019s seat coincides with the machine\u2019s longitudinal axis" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003887_19346182.2012.663534-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003887_19346182.2012.663534-Figure1-1.png", "caption": "Figure 1. Direction of racquet head and spin axis in a kick serve, as viewed by (a) a right hander and (b) a left hander. The ball is traveling into the page toward the net. The spin is primarily sidespin here with a small topspin component. The aerodynamic Magnus force, F, acts at right angles to the spin axis, pushing the ball down onto the court and causing it to curve to the left in (a) or to the right in (b). F is in the opposite direction to the friction force on the ball generated by string motion across the back of the ball.", "texts": [ " The amount of topspin has not previously been measured, although it is known than spin rates of around 4000 rpm are generated by elite tennis players when serving a ball (Choppin et al., 2007; Goodwill et al., 2007). When serving a kick serve, right-handed players toss the ball over their left shoulder, arch their back, bend at the knees and then jump up off the court. The end result is that the racquet head strikes the ball in a direction that is partly sideways across the back of the ball and partly vertical up the back of the ball, as shown in Figure 1. The sideways component generates sidespin and the vertical component generates topspin. In Figure 1, the ball will have more sidespin than topspin since the horizontal speed of the racquet head is greater than the vertical speed. D ow nl oa de d by [ U ni ve rs ity o f D el aw ar e] a t 1 7: 40 2 7 Ju ly 2 01 3 Sidespin causes the ball to curve from right to left through the air, as viewed by a right\u2013handed server, or from left to right for a left\u2013hander. Topspin causes the ball to curve down onto the court at a rate that is faster than the effect of gravity alone. The situation shown in Figure 1 is the one normally used in coaching manuals to describe how players need to strike the ball in a kick serve (see, for example, Braden, & Bruns, 1998), and it shows how sidespin is generated as well as topspin. That does not mean that the ball spins about two separate axes. There is only one spin axis and it is tilted away from the vertical. If the axis in Figure 1 is vertical then there is no topspin, just sidespin. If the axis is horizontal then there is no sidespin, just topspin. If the axis is tilted then any point on the ball will rotate in a circle around the axis, and it rotates simultaneously in vertical and horizontal directions. The amount of topspin generated in a serve due to vertical motion of the racquet head can be estimated by considering the situation shown in Figure 2. The racquet head is vertical and is approaching the ball rapidly at speed V and angle A", " The latter result follows from the fact that the rebound angle B in Figure 3 increases when the angle of incidence increases. The result, when transformed back to the reference frame in Figure 2, is an increase in the launch angle as the approach angle, A, increases. Figure 2 also describes the result when the racquet is moving sideways across the back of the ball, and is approaching the ball at a sideways angle A. In that case, the ball acquires sidespin, and the amount of sidespin is given by the same expression. In practice, the racquet head usually approaches the ball as shown in Figure 1, with a large sideways approach angle and a relatively small vertical approach angle. D ow nl oa de d by [ U ni ve rs ity o f D el aw ar e] a t 1 7: 40 2 7 Ju ly 2 01 3 As a result, the ball is usually served with about 4000 rpm of spin in a kick serve, but the spin is mostly sidespin and the amount of topspin is relatively small. That is, the spin axis is almost vertical, as indicated in Figure 1. Suppose a racquet approaches a ball in a horizontal direction at speed V and the ball is falling vertically at speed v just prior to impact, as shown in Figure 4(a). In a reference frame where the ball is at rest, the racquet is rising vertically at speed v while simultaneously moving horizontally at speed V, as shown in Figure 4(b). The situation is then the same as that shown in Figure 2 and the spin is given by Equation (1). If the ball falls say one metre before it is struck then it will be falling at 4", " All marks on the ball rotate in circular paths around the axis, so the axis can be identified by the motion of those marks. If the axis is horizontal and pointing in the same direction as the motion of the ball, then the spin is classified as gyrospin. That type of spin is used when throwing an oval shape football. If the spin axis is essentially vertical but tilted in a direction toward the net in a serve, then the ball will have a small gyrospin component. If the spin axis is tilted sideways, in a direction parallel to the net or the baseline (as in Figure 1) then the ball will have a small topspin component. The approach used by the author to determine the spin axis was to mount a ball in such a way that its axis could be fixed in any given position, and then to rotate the ball about that axis in order to compare the result with the video film. A certain amount of trial and error was needed to identify the spin axis, but it was usually close to the orientation shown in Figure 1. The inclination of the spin axis could easily be determined to within five degrees by this method. For some players, the axis was tilted away from the vertical by about 10 or 208, while others served the ball with an axis tilt of about 30 to 408. In some cases, the axis was also tilted slightly toward the net, meaning that the ball was struck slightly toward the front of the ball rather than exactly at the rear of the ball. For most of the serves, the racquet head was rising at about 58 just before impact and either continued to rise at a smaller angle at the end of the impact or travelled in a horizontal or downward direction at the end of the impact" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000923_tasc.2005.849119-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000923_tasc.2005.849119-Figure4-1.png", "caption": "Fig. 4. Fragment of a tilted Bitter coil with elliptical disks. Holes for tie rods are shown (neglected in the analysis).", "texts": [ " Furthermore, theoretically, even a wire of variable cross-section may be used. In both cases the field uniformity is not affected, as the current is still constant along the wire, and the above proof holds. These options may have some advantages. For example, the gaps (see (5)) can be eliminated using a wire of variable section. As a result, the total resistance of the winding can be decreased in resistive coils, and the mechanical strength can be increased. As opposed to wire-wound tilted coils, Bitter tilted coils (Fig. 4) are much more complicated for analysis. The magnetic field from a Bitter tilted coil can be adequately approximated by the field of a number of identical elliptical disks parallel to each other. In our analysis, we neglect the presence of numerous holes and slits in the disks. While for a conventional circular Bitter disk the current density distribution obeys a very simple analytical formula [7], this is not the case for the current density in an elliptical disk (Fig. 5). However, a more complex exact solution was found for the latter for an arbitrary tilt angle" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000930_robot.1985.1087262-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000930_robot.1985.1087262-Figure7-1.png", "caption": "Figure 7. Gear Eccentricity", "texts": [ " The PUMA base joint is not preloaded by gravity, and C* has to be constantly updated if forward motion is executed without any reversal of the motor shaft. Then ,the gear's are in contact, i.e. C+ = 0 . E 1 Backlash Compensation The backlash may now be compensated on-line by altering the demand signal to the motor as: Om = ngnom Iearm + Cn+ Kg, deb] (14) 3. Errors in Joint Angle Due to Variation In Reduction Ratio One other problem which effects the positioning accuracy of the joint is the gear eccentricity. Machined gears tend to be elliptical, and a very small variation in the dimensions between the axes exist, see Fig. 7. The input gear will usually have a much smaller variation in dimensions than the output gear, if a reduction ratio exists. As a result the smaller drive gear can be assumed to be near perfectly round, of radius ro (see Fig. 7). The larger gear has a radius which varies from rgnom to r max. As the gear turns through an angle Og, the radius of the contact point is given as : r,(O) = rgnom (1 + cg Sin 0,) (15) where t, = [ rgmax ] - rgnom rgnom Gear Eccentricity usually E < Le. a small eccentricity. The reduction ratio is tken given as: and ngnom is the ideal gear ratio. As ngnom > 1, for any joint, the final gear pair will have the most significant effect in terms of the variation of the reduction ratio with respect gear angle Bg : Variation of reduction ratio with respect to two gears in the drive train is given in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000398_1.1814651-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000398_1.1814651-Figure1-1.png", "caption": "Fig. 1 A symmetrical Stewart parallel manipulator", "texts": [ " Through investigating the concrete moving behaviors of parallel manipulators at singular points and by adding disturbing control function, Wang et al. @18# presented a method for a 3-dof planar manipulator to pass through the singular positions with the persistent configuration. In this paper, we will investigate the detailed bifurcation characteristics of a symmetrical Stewart parallel manipulator in the vicinity of the singular points, so as to find the way to control the moving platform passing through the singular points with the persistent configuration in the further study. Figure 1 shows a symmetrical Stewart platform manipulator with the fixed base frame and movable platform ~end-effector! connected by six extensible legs with spherical joints at both ends. Each leg is a system of two bodies connected by a prismatic actuation giving the variable length. The fixed coordinate system Oxyz is established on the fixed base frame. The centers of six spherical joints Bi(xi0 ,yi0 ,zi0) connected on the fixed base frame are distributed on a circle with radius R2 , and joints B1 , B3 , B5 and joints B2 , B4 , B6 form two equilateral triangles, respectively, and the relative angle between them is a2 " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002332_j.mechmachtheory.2009.05.007-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002332_j.mechmachtheory.2009.05.007-Figure6-1.png", "caption": "Fig. 6. Kinematic mechanism design from \u2018\u2018Results\u201d.", "texts": [ " As mentioned this book was very popular: One edition edited by his successor, Grashof, in 1875 was in the library of the well-known American engineer-professor Robert Thurston of Stevens Institute and later in Cornell University. There are mathematical curves describing the cycloid curves used in gear design (Fig. 5). In another section there are designs for balancer beams for steam engines that use a variable cross-section to obtain a section with constant maximum bending stress. Redtenbacher also designed a number of kinematic mechanisms as shown in Fig. 6. This combined with his published books in water wheels, turbines and railroad vehicles demonstrates his wide range of expertise in mechanical engineering and his expectations for students in the Karlsruhe Polytechnic School. A comparable engineer-scientist of his era would have been Rankine who also published in a wide range of mechanical sciences. Mechanics was omnipresent in Redtenbacher\u2019s work: \u2018\u2018All-around where something moves, mechanics has a finger in the pie but the spirits do not stir by mechanics\u201d, could be read 1856 in the caption of a Redtenbacher portrait" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000286_00021369.1984.10866431-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000286_00021369.1984.10866431-Figure2-1.png", "caption": "FIG. 2. Polarograms of (A), 0.5 mM p-Benzoquinone in pH 5.0 Acetate Buffer and (B), (A) Plus 20mM D-Glucose at a GOD-Immobilized Graphite Disk Electrode Rotated at 600 rpm.", "texts": [ " The magnitude of the limiting current was the same as that obtained with a rotating-disk bare-graphite electrode, though the half-wave potential was shifted by about -60mV, indicating that the presence of the immobilized enzyme layer does not markedly affect the electron exchange of BQ with the electrode. In the presence of 0.5 mM BQ and 20 mM D-glucose in basal buf fer solution, the anodic current was observed at potentials more positive than 0.2 V, indicat ing that BQ can work as an electron transfer mediator at a GOD-immobilized electrode. Figure 2 shows polarograms of (A) 0.5 mM BQ in, a deaerated basal buffer solution and (B) (A) plus 20 mM D-glucose at a GOD immobilized graphite electrode rotating at 600 rpm. In recording the polarograms in Fig. 2, we used a disk graphite electrode, on which a larger amount of GOD (some five times larger than the amounts given above) was \"immobilized\" by covering the enzyme-loaded electrode s~rface with a thin collodion mem brane before washing. In the presence of D glucose (curve B), the anodic current started to appear at 0.2 V, and increased with increasing D ow nl oa de d by [ 19 0. 10 6. 61 .9 7] a t 0 6: 50 1 7 Ja nu ar y 20 16 1972 T. IKEDA et al. I III (Ered)im, according to reaction schemes I and II, respectively, and that hydroquinone (Mred) produced in scheme II is oxidized at the electrode surface to regenerate BQ (Mox) by scheme III, giving the anodic current" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001984_robot.2008.4543370-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001984_robot.2008.4543370-Figure3-1.png", "caption": "Fig. 3. (a)-(b) Center-of-mass stable equilibrium regions for two stances with different radii of gyration, (c) Illustration of a stable equilibrium stance.", "texts": [ " The shaded regions in both figures are center-of-mass regions satisfying the kinematic-strong equilibrium condition. The dashed regions are center-of-mass regions satisfying the clattering convergence condition (10). The intersection of these two regions gives center-of-mass locations achieving equilibrium and frictional stability. Note that there is a tradeoff regarding the radius of gyration of B, as follows. Increasing \u03c1 results in larger region of kinematicstrong equilibrium, but smaller region of clattering convergence, as in the example of Fig. 3a, where the intersection region lies entirely under the terrain. In the example of Fig. 3b, with \u03c1 twice smaller, the kinematic-strong equilibrium region is smaller, but the clattering convergence region is enlarged, and the intersection region has a portion above the terrain, making stable equilibrium postures practically achievable. Finally, Fig. 3c shows an illustration of a twolegged mechanism (treated as a single rigid body) positioned in a stable equilibrium posture by keeping the center-ofmass location sufficiently low. The small radius of gyration is achieved by a massive central body and relatively thin limbs. This paper analyzed the impact-induced hybrid dynamics of a planar rigid body with two contacts, and derived a condition guaranteeing that its solution converges in finite time to a Zeno point of either one- or two-contact re-establishment" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003032_j.wear.2009.01.047-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003032_j.wear.2009.01.047-Figure3-1.png", "caption": "Fig. 3. Different oil groove shapes: (a) no groove, (b) two groove, (c) three grooves, (d) Z shape, and (e) geometry of oil groove (all dimensions in mm).", "texts": [ " Grooves of three configurations as shown below were milled on the surface of the blocks. The grooves were provided at right angles to the direction of motion and connected together internally for the distribution of the lubricant over the entire surface. The grooves are provided on only one surface of the blocks. ith li T m o u C K D F P he locating dimensions and pitch [7] are as shown in Fig. 5. Comercially available lubricant is introduced through the block and n to the slide by gravity flow. The specifications of the lubricant sed are (Fig. 3): olor Golden yellow inematic viscosity@ 40 \u25e6C (cSt) 68 ensity@ 15 \u25e6C (kg/l) 0.882 lash point (\u25e6C) 219 our point (\u25e6C) \u221227 near velocity for different groove shapes. Experiments were carried out at room temperature and under lubricated conditions. The slide way and the block are cleaned with acetone. The block is fixed to the slide way and connected to the force transducer. The load hanger attached to the lever arm is loaded with suitable weights such that the desired pressure is applied on the block" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001108_1.2401215-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001108_1.2401215-Figure3-1.png", "caption": "Fig. 3 Experimental apparatus", "texts": [ " The ball locations in the two upper rows i=1,2 are shifted half the distance s between the adjacent balls to those in the two lower rows i=3,4 ; 2. Ball grouping II. The ball locations in the two left rows i =2,3 are shifted s /2 to those in the two right rows i =1,4 ; and 3. Ball grouping III. The ball locations in the second row i =2 , in the third row i=3 , and in the fourth row i=4 are shifted s /2, s /3, and s /5, to those in the first row i=1 , respectively. 2.2 Experimental Method. The experimental apparatus is shown in Fig. 3. The rail of the test bearing is fixed on the concrete bed by bolts. The carriage of the test bearing is driven at a 2007 by ASME Transactions of the ASME of Use: http://www.asme.org/about-asme/terms-of-use c b L T m s t G w r b t w p v E a t 3 b m b v o n g o s g s A C C C R R R B D N C O C C J Downloaded Fr ertain linear velocity. The pitching and yawing motion of the carriage was detected y using a laser autocollimator Chuo Precision Industrial Co., td: LAC-S and a mirror, and was stored in a personal computer. he positive and negative directions of the pitching and yawing otion of the carriage were defined as shown in Fig. 3. Figure 4 hows the mounting position of the mirror. Because the gravitaional center Gm of the mirror was right above that of the carriage c, the gravitational center G of the mirror attached the mirror as also right above Gc. In the measurement using the laser autocollimator and the miror, the motion caused by the deformation of the rail as well as the all passage vibrations were detected 11 . To avoid the effect of he motion caused by the deformation of the rail, a high pass filter as used. The cutoff frequency was 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001980_sisy.2008.4664900-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001980_sisy.2008.4664900-Figure6-1.png", "caption": "Fig. 6. Illustration of the notion of dynamic balance", "texts": [ " Namely, if we assumed that the foot is immobile with respect to the ground and the calculation gives a ZMP position outside support area, this means that the assumption about foot immobility was not correct, and the mechanism will rotate about the edge of the support area. In such a situation, none of the points inside support area can be declared ZMP or, in other words, the mechanism is not in the state of dynamic balance. Let us try to demonstrate the essence of dynamic balance on one more example. We consider the four-link physical inverted pendulum supported on the ground by its bottom link (Fig. 6). All the joints of the pendulum (J1, J2 and J3) are powered. If the links L2, L3, and L4 move so that the link L1 remains immobile with respect to the ground (Fig. 6a), then we say that the system as a whole is dynamically balanced. However, if the motion of the links L2, L3 and L4 causes the contact of the L1 and the ground becomes as shown in Fig. 6b, the system will not be dynamically balanced. Further, let us assume the link L1 consists of two links, link L1A and link L1B, connected via the joint J0 (Fig. 6c). In the examples shown in Figs. 6a and 6b, the joint J0 is locked, so that the links L1A and L1B behave as a single sol- id body. If we assume that the joint J0 is also active, the system will be dynamically balanced only if the link L1A is immobile with respect to the ground, as presented in Fig. 6c. In this case, the link L1B may move like any other link (L2, L3 or L4), and need not be in contact with the ground. Let us show once more, this time using the examples from Fig. 6, where ZMP and FZMP can be located. In the examples where the system is dynamically balanced (Figs. 6a and 6c) the ZMP is inside the contact area of the ground and the link L1 (Fig. 6a), and L1A (Fig. 6c). In the example shown in Fig. 6b the ZMP does not exist, but in that case it would be possible to determine the position of the FZMP on the ground surface that would be outside the area which in Fig. 6a belonged to the ZMP. If, however, the supporting link were larger than the link L1 shown in Fig. 6a (i.e. large enough to cover the FZMP), the system would remain dynamically balanced and the situation shown in Fig. 6b would not arise. (Not to blurr the essence of dynamic balance, neither ZMP nor FZMP were shown in Fig. 6). In summary, the ZMP can be only inside the support area (as long as the foot remains immobile on the ground), and by no means at its edge when the foot loses immobility. When the foot rotates, the ZMP does not exist. But, the CoP (Center of Pressure) will still exist. This is the difference between the ZMP and CoP. Hence, the ZMP should be used as the indicator of dynamic balance, but not the CoP. And, let us now conclude. We can speak of ZMP only while the humanoid robot is dynamically balanced" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002239_j.euromechsol.2008.06.008-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002239_j.euromechsol.2008.06.008-Figure1-1.png", "caption": "Fig. 1. Stable region and ZMP. The stable region is approximated by the largest ellipse inscribed in the WMM sustentation polygon. (a) Hexagonally-shaped sustentation polygon; (b) Diamond-shaped sustentation polygon.", "texts": [ " In their model, inertia effects of the rigid bodies are taken into account, which is very important to the system dynamics. The ZMP coordinates can be computed by Newton\u2013Euler formulation as follows (Kim et al., 2002): x0 = \u2212 N y fz and y0 = Nx fz . (4) N and f are, respectively, the resultant moment and force applied to the platform. The above coordinates are expressed in the moving frame MP R \u2032 = (M S P ,MP x,MP y,MP z), where MSP is the Most Stable Point defined as the center of the contact polygon between the platform and the ground. The area enclosed in this polygon is the stable region (Fig. 1). The system is stable as long as the ZMP belongs to this region (Huang et al., 2000). In most cases, WMMs have a symmetric platform (Kim et al., 2002) so that the stable region of these systems is usually approximated by the largest circle/ellipse inscribed in the sustentation polygon (Fig. 1). With this simplification, the stability index \u03a6 is defined as follows: \u03a6(q, q\u0307, q\u0308) = 1 \u2212 [( x0 a )2 + ( y0 b )2] (5) where a and b are the half-diameters of the inscribed ellipse. We note that \u03a6 is a dimensionless number that takes positive (negative) values if the ZMP is inside (outside) the stable region. In particular, \u03a6 = 0 if the ZMP is at the boundary of the stable region and \u03a6 = 1 if the ZMP coincides with the MSP. The WMM is required to move freely from a given initial configuration \u03a9START = [(XS p)T, (qS a) T]T to a given final configuration \u03a9GOAL = [(XG p)T, (qG a )T]T (in the case of a GPP task, Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001460_1.2713788-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001460_1.2713788-Figure1-1.png", "caption": "Fig. 1. Arrangement used for first experiment.", "texts": [ " Penner1 provided a theoretical description of the gear effect to determine the optimum shape of the head of a golf club, assuming that the ball exits the club in a rolling mode. That is, the tangential speed of the spinning ball is equal to the tangential speed of the club face. The objective of this paper is to provide experimental data on the gear effect to elucidate the physics of this phenomenon. For that purpose, two experiments were conducted. The first was a simple experiment in which a tennis ball was placed at rest on a horizontal surface, as shown in Fig. 1. The surface was then accelerated in a horizontal direction to measure the resulting speed and spin of the ball. In the second experiment, a golf ball was swung as a pendulum bob to impact a rectangular block of wood, as shown in Fig. 2. The spin acquired by the ball was measured for two values of the impact parameter and at various angles of incidence. We found that the spin imparted to a ball by the gear effect arises from the static friction force that is generated when there is no slip between the contacting surfaces and when both surfaces are accelerating in a direction tangential to the two surfaces", " Second, the normal reaction force, N, continues to act during the grip phase and generates a tangential acceleration of the surface if N acts along a line that does not pass through the center of mass of the colliding object and the center of mass does not lie within the plane of the surface.3 Depending on the direction of the tangential motion of the surface, the gear effect during the grip phase can either increase or decrease the spin of the ball acquired during the sliding phase. In the second experiment, the gear effect was especially evident by the strong asymmetry in the outgoing ball spin with respect to the sign of the angle of incidence. Suppose that a ball of mass m and radius r is at rest on a horizontal surface, as shown in Fig. 1. If the surface is initially at rest and then moves horizontally to the right with acceleration as, the ball will be subject to a horizontal friction force F acting to the right. If a is the acceleration of the ball center of mass and the ball rotates with angular acceleration , then F=ma and Fr= I , where I= mr2 is the moment of inertia of the ball about an axis through its center of mass. For a solid sphere, =0.4. For a hollow tennis ball with a 6 mm thick wall, =0.55. If the ball center of mass has velocity v and the ball has angular velocity , then a point P on the ball in contact with the surface will have velocity vP=v+r and an acceleration aP=a+r , assuming that the ball rotates in a counter-clockwise direction, as indicated in Fig. 1. An interesting physics question is whether P has the same acceleration as the surface or whether there is some slip. Also of interest is the relation between F and the normal reaction force N=mg. If we define the coefficient of friction between the ball and the surface as =F /N, then the ques- 658 658Am. J. Phys. 75 7 , July 2007 http://aapt.org/ajp \u00a9 2007 American Association of Physics Teachers This article is copyrighted as indicated in the article. Reuse of AAPT content is subject to the terms at: http://scitation", "04 s by filming the block and ball motion with a digital video camera at 25 frames/s. Quadratic fits to the displacement data were used to calculate the relevant velocities and accelerations to within 2%. The accelerations a, as, and remained constant to within 2% while the ball travelled the 400 mm distance from one end of the surface to the other. The results are shown in Table I, including the calculated distance D between the line of action of the normal reaction force N and the ball center of mass. If N acts at a point to the left of the ball center of mass in Fig. 1, then the torque on the ball is given by Fr\u2212ND= I . In all cases, it was found that D was 0.9 mm. For low values of as, the ball accelerated with a=as and without rotation. Such a result can be explained if ND is equal and opposite to Fr. A positive value of D can be explained in terms of ball rotation. The leading edge of the ball rotates into the surface, thereby increasing the normal reaction force at the front edge. In an analogous fashion, the front end of a vehicle dips down when the brakes are applied" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001536_acs.1009-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001536_acs.1009-Figure4-1.png", "caption": "Figure 4. k\u2212 f and (k) for the dead zone and the second-order dead zone.", "texts": [ " Lemma 1 The functions d F and sd F fulfill the property P(BF ). Copyright q 2007 John Wiley & Sons, Ltd. Int. J. Adapt. Control Signal Process. 2008; 22:590\u2013608 DOI: 10.1002/acs Proof First, note that (K )=0 for K inside the set BF is trivial. Second, recall that the scalar product is such that Tr( (K )(K \u2212F)\u2217)= m\u2211 i=1 l\u2211 j=1 F (i, j) (K(i, j))(K(i, j)\u2212F(i, j)) \u2217 which is the sum of all terms of elements such as f (k)(k\u2212 f )\u2217. These terms are scalar products of two vectors in the complex plane as illustrated in Figure 4. They are necessarily non-negative. Moreover, the sum is strictly positive as soon as one element of K is outside the set BF . A first series of simulations is made taking initial conditions x(0)=(10 0 0 0)T, K (0)=[0 0 0] Copyright q 2007 John Wiley & Sons, Ltd. Int. J. Adapt. Control Signal Process. 2008; 22:590\u2013608 DOI: 10.1002/acs where =1, F=[10 10 10] and = d F . Robustness with respect to is illustrated for =0, 5 and 9.655. Time histories of both the output y(t) and the control gain K (t) are plotted in Figures 5\u20137, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001838_0094-114x(75)90072-5-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001838_0094-114x(75)90072-5-Figure5-1.png", "caption": "Figure 5. Rotating elliptic disks connected by a crossed belt.", "texts": [ " when the elliptical base g is reduced to the focal segment EF. Without pretending that these heuristic considerations are a rigorous proof of Graves' theorem--an exact proof will be indicated in section 3--we can complete Fig. 4 by adding its mirror image with respect to the tangent t and thus obtain a mechanism equivalent to the pair of elliptic wheels pj, p2: it consists of two equal elliptic disks gl and g2, confocal with p, and pz, respectively, and connected by a closed crossed belt which transmits the rotation about E, into a rotation about F~ (Fig. 5). Of course the transmission ratio of the angular velocities is not constant. The reasoning was essentially based upon the fact that the rolling curves p, and p: are always symmetrically situated with respect to their common tangent t. It is easy to see that this symmetry occurs only for elliptic wheels. Nevertheless the transmission problem can be posed for other noncircular wheels too, e.g. for those which were investigated in [4, section 45]. The existence of closed transmission belts has been led back by Zenow to a difficult system of functional equations, but it seems that there do not exist solutions in the general case", " We propose to take for c, an involute of some ellipse g,, confocal with pl. Thus each normal t, of c~ is tangent to g~ and cuts p] in a first point P~. According to Reuleaux this normal t, has to be transferred to the corresponding point P2 of p=, forming there the same angle a with p2 as t, forms with p,. This position t2 is then a normal of the second tooth profile c2. Corresponding to Poncelet's theorem (section 2), t2 is a tangent of the ellipse g2 which is equal to g, and confocal with p=. Hence the profile c2 is an involute of g= (Fig. 5). Corresponding points C, on c~ and C= on c2 are determined by equal segments P,C~ of t~ and P=C2 of &, as they become the contact point of the tooth profiles when, after a certain rotation of the wheels, P, coincides with P., (on the line E,F=). As pointed out in [4, section 40] for circular wheels, but valid also for non-circular ones, the contact element of corresponding tooth profiles c, and c= varies in such a way that the normal of the element is rolling on the evolute gl of c~ and on the evolute g2 of c2 simultaneously; this fact is again a simple consequence of the three-pole theorem", " Thus the normal might be realized by a belt wound around g, and g2. Repeating this conclusion for opposite tooth profiles in the case of elliptic wheels, we have a rigorous proof for Graves' theorem without needing the usual elliptic integrals. Summarizing we may consider our corresponding tooth profiles c,, c2 for elliptic gears as the paths of a point C fixed on the straight part of an auxiliary transmission belt, wound around elliptic base curves g,, g2 which are confocal with the rolling pitch ellipses p,, p2 (Fig. 5 and 6)/: This is a natural generalization of the known interpretation of the usual involute profiles for tit seems attractive to use the degenerate ellipses g, = E~F, and g= = E2F2, as then the tooth profiles c, and c2 would become circular arcs, but it is obvious that teeth of.this simple shape would not work in critical phases of the motion. In fact, in those positions where E,, E2, F,,F. are on the same straight line, the commoh normal in the contact point of corresponding circular tooth profiles passes through the rotation centers E, and F2, and thus these profiles are not able to transmit rotation", " Corresponding developments could be performed for elliptic bevel gears by means of operations on the sphere analogous to those used here in the plane. Elliptic gear wheels are based upon the motion of an antiparallelogram E,F2E2F, and its pole curves which consist of two equal ellipses pl and p2 with foci E,, FI and E2, F2 respectively (Fig. 1). Any two smaller equal ellipses gl and g2, confocal with the rolling ellipses pl and p: and rigidly connected with them, have the property that a crossed transmission belt slung around them does not disturb the original motion (Fig. 5). This fact gaves the possibility for an apparently new gear system for elliptic wheels which represents a natural generalization of the well-known involute system for circular wheels: The tooth profiles are throughout involutes of the \"base\" ellipses g, and g~ (Fig. 6). Corresponding tooth profiles c,, c2 may be considered as the relative paths of a point C on the straight part of the mentioned transmission belt (Fig. 5). The problem of similar developments for other non-circular wheels remains open. [1] DINGELDEY F., Kegelschnitte und Kegelschnittsysteme. Enzykl. math, Wiss. III C1 (1903). [2] HINKLE R. T., Kinematics o/Machines. Prentice-Hall, Englewood Cliffs (1960). [3] KLEIN F. und BLASCHKE W., Vorlesungen iiber h6here Geometrie. Grundlehr. math. Wiss. 22, (1926). [4] WUNDERLICH W., Ebene Kinematik. Hochschultaschenbiicher, 433/433a. (1970). Beltrag zur Geometrle elliptischer Zahnriider Kurzfassung--Die zur Relativbewegung der beiden k~rzeren Seiten eines gelenkigen Antiparallelogramms gehSrigen Polkurven sind bekanntlich zwei kongruente Ellipsen, die ihre Brennpunkte in den Gelenken haben und deren Hauptachsenl~nge mit jener der I~ngeren Antiparallelogrammseiten fibereinstimmt (Bild 1, gleichliiufiges Zwillingskurbelgetriebe)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002575_tmech.2008.2000638-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002575_tmech.2008.2000638-Figure5-1.png", "caption": "Fig. 5 Dynamical model of cart and inverted pendulum system.", "texts": [ " By using the renewed virtual reference signal V \u21d0 V + \u03b2\u2206V (21) procedures of (18)\u2013(21) are iterated until ef is sufficiently small and the final state after the \u03bbth learning iteration becomes as follows: x\u0304(\u03bb+1) (N ) \u223c= (1 \u2212 \u03b2)\u03bbx\u0304(1) (N ) + \u03bb\u2211 \u03b4=1 (1 \u2212 \u03b2)(\u03b4\u22121)\u03b2 x\u0304o . (22) When the iteration times \u03bbth become sufficiently large, the final state x\u0304(\u03bb+1) (N ) converges to the desired state x\u0304o . After error learning, by adding V renewed into the augmented system (15), r(j) is obtained from the state variables of x\u0304. III. 2-DOF CONTROL SYSTEM DESIGN FOR A CART AND INVERTED PENDULUM SYSTEM The dynamical model of a cart and inverted pendulum system dealt with in this research as a controlled object is shown in Fig. 5. Because this plant has only a single input, the subscript i = 1 is omitted hereafter. The specifications of the experimental setup are shown in Table I. The control input us is a voltage signal input to an actuator to move the cart. KT is the thrust constant and the actuator thrust is KT us [N]. We assume that the input voltage signal us saturates at a magnitude of \u03b1 = 1 V. First, we show the GS feedback controller designed applying our proposed method [1]. The state equation and the output equation are formulated as d dt [ x us ] = [ A BCs 01\u00d74 As ] [ x us ] + [ 04\u00d71 Bs (p) ] u\u0307 (23) [ y ue ] = [ C 02\u00d71 01\u00d74 De (q) ] [ x us ] (24) where Bs (p) = p = sech2 (u/\u03b1) and De (q) = q = ue/us = u/us \u2212 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003740_cdc.2010.5717382-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003740_cdc.2010.5717382-Figure3-1.png", "caption": "Fig. 3: Perspective Projection Model", "texts": [ " A coordinates transformation yields the positions of feature points relative to a frame \u03a3i, denoted by pijk , as below. pijk = gijpjjk , where pijk and pjjk should be regarded, with a slight abuse of notation, as [pT ijk 1]T and [pT jjk 1]T via the wellknown homogeneous coordinate representation in robotics, respectively (see, e.g., [19]). Let us now consider visual measurements of each rigid body. We denote the k-th feature point onto the image plane as fijk := [fxijk fyijk ]T \u2208 R2, k \u2208 {1, \u00b7 \u00b7 \u00b7 ,m}. Then, by perspective projection (Fig. 3), this is given by fijk = \u03bb zijk [ xijk yijk ] , where \u03bb \u2208 R is a focal length and pijk = [xijk yijk zijk ]T [19]. Moreover, let fij be the stuck vector of m feature points on image plane coordinates, i.e., fij := [ fT ij1 \u00b7 \u00b7 \u00b7 fT ijm ]T \u2208 R2m. We assume each body can extract the feature points of visible bodies from image data, namely, the measured output of body i is fi := (fij)j\u2208Ni , i \u2208 V. (4) Hereafter, the aggregate system consisting of n rigid bodies with kinematic model (1), the visibility structure (3) satisfying Assumption 1 and measured output (4) is called visual robotic network \u03a3" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002691_cav.194-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002691_cav.194-Figure4-1.png", "caption": "Figure 4. Tetrahedral mass-spring system.", "texts": [ " Here, soft tissues are modeled with viscoelastic behavior by a volumetric tetrahedral mass-spring system with attention focused on small deformations restricted into a local area. Viscoelastic Behavior Human organs in the study are constructed by volumetric tetrahedral meshes. It is convenient to extract multiple iso-surfaces among the different tissues. Besides, the tetrahedral element can support modeling of 3D organs with arbitrary shape. The volumetric tetrahedral mass-spring system consists of mass points and connected springs along the edges (Figure 4). The Viogt rheological model in Figure 4 (left) is used to depict the time-dependent viscoelastic behavior of tissues. The linear springs obey the Hook\u2019s law, whereas the viscous dampers generate a resistance force proportional to the velocity. The dynamics of points are governed by the Newton\u2019s Second Law of motion. The nodal displacement of the ith point (ui2R3) due to an external force Fi is given by the following, mi\u20acui \u00fe di _ui \u00fe P j sij\u00f0 ~rijj j lij\u00de ~rijj j ~rij \u00bc Fi (10) where mi is the mass of the point i, di is the damping constant of the same point, r * ij is the vector distance between point i and point j, lij is the rest length, and sij is the stiffness of the spring connecting two mass points" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002140_1.5061038-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002140_1.5061038-Figure2-1.png", "caption": "Figure 2. The coaxial deposition nozzle used for model verification (vertical section).", "texts": [ " For example, by first inspection, increasing the mass flow of the deposited material, its specific heat capacity or its conductivity, or reducing total laser power or absorptivity are all predicted to give lower temperatures and hence a reduced overall melt pool size. The model was tested by applying it to deposition of gas-atomised Inconel 718 powder using a Laserline LDL160-1500 diode laser fitted with optical conveyance fibre and optics and a coaxial nozzle assembly made by DeBe Lasers Ltd (http://www.debe.co.uk/). The nozzle (Figure 2) is of conventional design, with a single powder outlet annulus and central gas stream to protect the focussing optics. Powder was conveyed from a disk type powder feeder using Argon gas. The nozzle was positioned vertically so that the powder stream converged to a single stream in the laser focal plane, 7.5 mm below the nozzle, which is where the substrate was placed during deposition. Page 728 Laser Materials Processing ConferenceICALEO\u00ae 2007 Congress Proceedings The powder distribution beneath the nozzle due to conveyance of the Inconel 718 powder was first tested using an optical light sheet method and the results compared with the model" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002891_1.4000647-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002891_1.4000647-Figure4-1.png", "caption": "Fig. 4 Scroll compression chambers", "texts": [ " ,z 5 The relationship between the normal components of radius- vector on two pitch lines of identical type can be expressed as Rn,l,1 = Rn,l,i + 2 c i \u2212 1 for l,1,s l,1,e 6 The pitch lines and scroll profiles can be expressed as Pl,i = Rn,l,i exp + 2 c i \u2212 1 + Rt,l,i exp j + 2 c i \u2212 1 Pl,i,m = Pl,i Ror 2 exp j + 2 c i \u2212 1 for l,1,s l,1,e 7 where the positive sign \u201c+\u201d is used for m=in and negative sign \u201c\u2212\u201d for m=out. Effective turns of every scroll pitch line is the same as nl,i = l,i,e \u2212 l,i,s = l,1,e \u2212 l,1,s = n 8 2 2 JANUARY 2010, Vol. 132 / 014501-110 by ASME Use: http://www.asme.org/about-asme/terms-of-use C f w t o b p d f s m t p 0 Downloaded From: 3.2 General Geometrical Model of Scroll Compression hamber. As shown in Fig. 4, every compression chamber is ormed by moving the outer or inner profile of the orbiting scroll rap in the direction of the profile normal at the meshing point up o the inner or outer profile of the fixed scroll wrap by a distance f Ror. The normal angle region of one scroll compression chamer is l,i , l,i+2 , where l,i is the normal angle at the meshing oint. The compression chamber formed by A-type pitch lines is efined as A-series scroll compression chamber, and the chamber ormed by B-type pitch lines is defined as B-series scroll compresion chamber" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002492_iemdc.2009.5075175-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002492_iemdc.2009.5075175-Figure9-1.png", "caption": "Fig. 9 Approximate finite-element solution: 2 blocks conducting", "texts": [ " One could assume that this is characteristic of \u201cresistance-limited\u201d eddy-currents, but it is more correct to describe it as the neglect of proximity effect, which in other situations is known to be dangerous. Nevertheless, engineers in all disciplines have long relied on simplified calculations combined with other means of verification than simply throwing time and money at the problem through very expensive calculating tools; and if this approach is still permitted, the method is surely worth trying. A \u201crefinement\u201d, if it can be so described, is to suppress the conductivity judiciously in pairs or patterns of magnets, and Fig. 9 shows an example where magnets A and C are treated as conductive while magnets B and D are not. The total loss calculated in this case is 1@02 W, again scaled up to include all magnets. This is reasonably close to the 0@97 W calculated using the single-block approximation, and the difference may give some idea of the accuracy of the method. The only way to be sure is to measure it, although some confidence might be obtained through more sophisticated calculation tools. An interesting observation in Figs" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003264_syscon.2011.5929071-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003264_syscon.2011.5929071-Figure1-1.png", "caption": "Figure 1. Body axis coordinate system [5].", "texts": [ " The motion of an aircraft is defined by translational motion and 978-1-4244-9493-4/11/$26.00 \u00a92011 IEEE rotational motion around a fixed set of pre-determined axes [4]. Translational motion is the one by which a vehicle travels from one point to another in space. For an orthodox aircraft there is only one direction in which translational motion occurs, that is the direction in which the aircraft is flying which is also the direction in which it is pointing. The rotational motion relates to the motion of the aircraft around three defined axes, pitch, roll and yaw [4]. Fig. 1 shows the direction of aircraft velocity vector in relation to the pitch, roll and yaw axes. For most of the flight an aircraft will be flying straight and level and velocity vector will be parallel with the surface of earth and proceeding upon a heading that the pilot has chosen. III. LONGITUDINAL DYNAMICS The longitudinal equations of motion were described by a set of linear differential equations. A very useful concept in analysis and design of control system is the transfer function. In the case of dynamics it specifies the relationship between the motion variables and the control input [1, 5]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002780_cae.20393-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002780_cae.20393-Figure5-1.png", "caption": "Figure 5 (a) Initial interface of the operation. (b) Graphs interface. [Color figure can be viewed in the online issue, which is available at www.interscience.wiley.com.]", "texts": [ " The main objective of this module is to contribute to the establishment of a relationship between the dynamic models of HUE operation with the VR environment, through visualization of the mechanical process of the turbine distributor. In this module, it is possible to analyze the dynamic behavior of the turbine-generator assemblage in a HUE, operating in both the permanent and transitory states, which allows the visualization of a HUE linked to a power system and an animated vision of the operation of the hydropower plant. By clicking on the \u2018\u2018Turn button,\u2019\u2019 as shown in Figure 5a, HUE starts operation, detaching the animation of the rotating parts for the turbine-generator assemblage in the virtual world. The animation of the jointed parts of HUE is governed by the electromechanical dynamics of the turbine-generator assemblage and, according to the event simulated, a certain contingency can dynamically alter the values of the rotor speed as the mechanical torque, the opening of the distributor, and the terminal voltage, among others. In the same sequence, by clicking on the \u2018\u2018Apply Contingency\u2019\u2019 button, the application of a short-circuit is accomplished in the terminal of the HUE substation", " By the application of this contingency, it is possible to verify the performance of the speed governor through the animation of the turbine-distributor system. For a Francis turbine, the control is exerted by an assemblage formed by the guide vane, servomotors, and regulation rings, collectively called the \u2018\u2018distributor.\u2019\u2019 Control is exercised by altering the position of the guide vane. With the purpose of procuring a graphic visualization of the main electromechanical magnitudes of HUE during the operation, the Graphs interface (see Fig. 5b) has been developed. This interface is enabled by clicking on the Oscillographies button as shown in Figure 5a. The visualization of the process of short-circuit in a hydraulic generator, obtained with the help of the VR environment, helps the professional to understand the phenomenon, which corresponds to the acceleration process of the generator and the consequence of this acceleration to both the turbine and the electrical system as a whole. Figure 5b shows the behavior of the frequency of the turbine in a condition of short circuit. There is an acceleration of that and then the return to a steady condition. Such frequency is related with the speed of the generator rotor. A single Francis turbine-generator, with the exciter and governor in a hydropower plant connected to the local load and infinite bus as shown in Figure 6a, is considered for the study. Although the actual power network is more complex, the single machine infinite bus system is, however, a useful starting point for a study of its design and performance [1,14,15]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002453_jst.23-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002453_jst.23-Figure3-1.png", "caption": "Figure 3. Positions of the center of mass and the center of buoyancy of a typical human body in a horizontal, motionless floating position. Buoyant force acts more cranial to the center of mass, generating a leg-sinking moment around the center of mass of the body.", "texts": [ " The stability of a human body in a horizontal, motionless floating position is determined by how the body is configured to form a certain posture and by the composition of the parts of the body. As many of us may know from our own experiences, our body is not generally stable in a horizontal, motionless floating position: the legs tend to sink to a lower position than the initial horizontal position. Studies confirm that the legs in fact tend to sink. This is due to the buoyant force acting more cranial to the CM of the body (Figure 3), generating the moment around the CM that causes the legs to sink [7,8\u201312]. Some of these studies examined the sex differences in the body\u2019s stability and found that women tend to float more horizontally than men [10,11,13], due primarily to women having a greater amount of body adipose tissue stored around the hips and thighs, causing the CB to be located closer to the CM [11]. The state of breathing was found to affect the stability in a horizontal, motionless floating position [10,12]. When air is inhaled, the lung volume increases and the CB shifts cranially to increase the leg-sinking moment" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002816_es2009-90480-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002816_es2009-90480-Figure6-1.png", "caption": "Figure 6 Mock Blade with PZT Patches", "texts": [ " Fig. 4 shows a typical three-blade wind turbine with a large nacelle that encloses the gearbox and generator. As shown, lightning rods are mounted to the top of the nacelle. The architecture for the proposed system is shown schematically in Fig. 5. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/08/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 5 Copyright \u00a9 ASME 2009 Sensing Interface After careful consideration, equally spaced piezoelectric patches (Fig. 6) are selected as the sensors of choice. Piezoelectric sensor/transducers (PZT\u2019s) are used to determine the characteristic impedance signatures of the blade by inducing strain in one of the patches with an electric (ac) signal. The actuating electric signal is chosen as a high frequency waveform (kHz) that imparts energy to the blade which produces a vibration that the other patches can sense. By establishing a characteristic vibration baseline of the blade in an undamaged condition, damage is detected by taking data and comparing it to this baseline" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003641_eej.21132-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003641_eej.21132-Figure1-1.png", "caption": "Fig. 1. Cross section of the PMa-SG.", "texts": [ " For claw-pole alternators used in automobiles, it is known that adding PMs between the adjacent claw poles improves the output power [1, 2]. However, no mention of the existence of the PM effect of reducing the magnetic saturation has been made in the literature. For conventional salient-pole synchronous machines, no other publications discuss an attempt to add PMs between the adjacent field pole shoes, so far as the authors are aware. Investigations of the effect of the PMs in the salient-pole synchronous machines are useful because the rotor structure is different from that of claw-pole alternators. Figure 1 shows a cross section of a PMa-SG. The PMs magnetized circumferentially are fixed between the adjacent field poles. The north and south poles of the PMs face the north and south field poles, respectively. The PMs are located in the leakage flux paths, not in the main flux paths produced by the armature and field mmfs. Therefore, the PMs do not disturb the main fluxes. The effect of the PMs is as follows. Under no-load conditions, the fluxes \u03c6pm are produced by the PMs and the fluxes \u03c6f are produced by the field current If as shown in Fig. 1(a). The directions of \u03c6pm and \u03c6f are opposite in the field poles. Therefore, the PM fluxes \u03c6pm decrease the magnetic saturation in the field pole bodies due to the field fluxes \u03c6f and hence increase the main fluxes by If. Accordingly the armature flux linkage and the terminal voltage increase. Under load conditions, the distribution of the main flux and ff + fa deviates as shown in Fig. 1(b) due to the fluxes \u03c6a produced by the armature current Ia. Magnetic \u00a9 2010 Wiley Periodicals, Inc. Electrical Engineering in Japan, Vol. 174, No. 4, 2011 Translated from Denki Gakkai Ronbunshi, Vol. 129-D, No. 1, January 2009, pp. 109\u2013116 saturation arises in the pole tips in addition to the pole bodies. The PM fluxes \u03c6pm also have a significant effect on pole-tip saturation. The reduction of saturation in the pole tips and the pole bodies by the PMs allows an increase in the terminal voltage", " In order to verify the effect of reducing the magnetic saturation, the terminal characteristics and the magnetic flux distributions in a PMa-SG and a conventional SG of the same dimensions were calculated by FEA [7\u201310]. 3.1 Analyzed machines The analyzed machines were the following conventional SG and PMa-SG. The conventional SG was a usual salient-pole synchronous generator without PMs. The rated values of the frequency, capacity, terminal voltage, armature current, and power factor were 50 Hz, 2.3 MVA, 650 V, 2021 A, and 1.0, respectively. The PMa-SG had additional PMs between the field pole shoes and had the same dimensions as the conventional SG. The configuration is shown in Fig. 1. Table 1 shows the specifications common to the two machines. The stators and rotors of these machines had no skew. 3.2 Method of analysis Two-dimensional transient magnetic field analysis was performed using the FEA software JMAG, Version 9.0 (JSOL). B\u2013H curves for nonoriented steel sheets 50H400 and 50H1000 were used for the stator and rotor cores, respectively. The PMs in the PMa-SG were NEOMAX38H (60 \u00b0C, remanent flux density 1.19 T, coercivity 902 kA/m, recoil relative permeability 1.05). For simplicity, eddy currents in the PMs were neglected", "21 T smaller than the remanent flux density. There is no irreversible demagnetization in the PMs. As an application example of the PM effect, the pole-body width can be made narrow instead of reducing the magnetic saturation in the pole bodies [7, 10]. In this case, the temperature rise in the field windings can be reduced. In Fig. 3(b), the flux lines passing through the PMs go to the stator beyond the air gap. However, the paths of the flux lines do not conflict with the flux paths of \u03c6pm shown in Fig. 1(a) (see Appendix). In conclusion, the PMs in the PMa-SG have allowed reduction of the magnetic saturation in the pole bodies and an increase of the terminal voltage under no-load conditions. 3.4 Short-circuit characteristics Figure 4 shows the computed three-phase short-circuit curves. The short-circuit currents of the two generators are almost the same. There is no effect of the PMs on the short-circuit current because no saturation arises in the field poles. Figure 5 shows the computed magnetic flux lines and flux density amplitude under short-circuit conditions at the rated current", " Experimental verification of a permanent-magnet-assisted salient-pole synchronous generator. Annual Conference of IEEJ Industry Applications Society, Y-103, 2008. (in Japanese) APPENDIX Flux Paths Produced by PMs In this appendix, the flux paths produced by the PMs are considered. In Figs. 3, 5, and 8, the flux lines passing through the PMs go to the stator beyond the air gap. However, the flux lines do not represent the flux paths produced by the PMs. Therefore, the flux lines do not conflict with the flux paths of \u03c6pm shown in Fig. 1(a). This can be explained as follows, taking the unsaturated no-load condition (If = 50 A) as an example. Linear FEA of the PMa-SG is performed under the conditions. The relative permeabilities in the rotor and stator cores are 2000 and 6000, respectively. Figure 12(a) shows the computed flux lines in the PMa-SG. The flux lines passing through the PMs go to the stator beyond the air gap. However, it should be noted that the flux lines in Fig. 12(a) are the result of superposition of the flux lines produced by the PMs and the field current", " By the superposition principle, combining the flux lines shown in Figs. 12(b) and 12(c) yields the flux lines shown in Fig. 12(a). This example illustrates that even when all the fluxes produced by the PMs remain in the rotor and do not go to the stator, combining the flux lines produced by the PMs and the field currents makes the resultant flux lines look as if the fluxes by the PMs pass to the stator. Therefore, the flux lines passing through the PMs and going into the stator in Figs. 3, 5, and 8 do not conflict with the flux paths of \u03c6pm shown in Fig. 1(a). In addition, when the rotor poles are highly saturated, part of \u03c6pm will flow across the air gap into the stator as stated in Section 2. AUTHORS (continued) (from left to right) Takahito Hayamizu (student member) received a B.E. degree in electrical engineering from Kanazawa Institute of Technology in 2008 and is currently working toward an M.E. degree. His current research interests include the design and analysis of electric machines. Kazuo Shima (member) received his B.E. and M.E. degrees from Kyoto University in 1993 and 1995, and D" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003498_s12239-011-0083-z-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003498_s12239-011-0083-z-Figure3-1.png", "caption": "Figure 3. Critical stress location.", "texts": [ " In this paper, the critical stress location of the stub axle was identified by developing a finite element (FE) model based on MSC/PATRAN\u2122 and MSC/NASTRAN\u2122. The inputs to the process are an FE model of the component, a set of cyclic material properties and a set of sinusoidal excitation bending loads ranging from 1000 N to 7000 N. An imported solid model of the stub axle from CATIA\u2122 was meshed using a second order tetrahedral element (TET10) topology, and a linear static analysis of the model was performed using MSC/NASTRAN\u2122. The results were then evaluated using MSC/PATRAN\u2122. The critical stress location of the stub axle is shown in Figure 3. The stub axle material is medium carbon steel JIS S48C. The chemical composition and mechanical properties for the JIS S48C steel are shown in Tables 1 and 2, respectively. A sample of ten stub axle units was subjected to a set of four different levels of cyclic bending fatigue loads. This cyclic loading was achieved by clamping the four mounting points at the base of the spindle with a two-ton clamping mechanism. The cylindrical end of the stub axle was attached to a load arm, which was connected to a motor with an eccentric mass to induce a moment" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002303_1.1660819-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002303_1.1660819-Figure1-1.png", "caption": "FIG. 1. Elliptical-edge disclination with Volterra's rotational angle 03.", "texts": [ "5 On: Mon, 01 Dec 2014 20:38:45 240 COMMUNICATIONS topic about relation between dislocations and disclin\u00b7 ations, brought up lately by deWit,9 is particularly interesting. The calculations of elastic fields and energy of a circular-edge disclination have been done by the au thors. IO They supported the existence of disclina tions in materials. The same procedures can be ex tended\u00b7 to apply in the case of elliptical-edge discli nations. An elliptical-edge disclination is formed by twisting two faces of an elliptical slip surface rela tive to each other through a constant rotational angle (Fig. 1). For an infinitely extended material in which plastic distortion,l:ltJ is prescribed by a func tion of coordinates, the displacement field isll u .. (X) = - (~lllill\", + jJ.lll1. llJI + jJ.llilllJ\") x rOOf 00 roo ~(ll,..xm +ll/mX,,: (3 -4v)ll/gnx, ).00 -ffO}.OO II r3 (1) where flu is Kronecker's delta, ~ and jJ. are Lamll constants, v is Poisson's ratio, and dx' = dx~dx;dx~, ~j=Xj-X~, }'2=Xj X j \u2022 For an elliptical-edge disclination as defined above, the plastic distortion may be written ~tJ = - nJE,k/O\"x;H(1 - xf2/a~ _~a/a=)I5(x~), (2) where a\" is Volterra's rotational vector, nJ is the unit normal of the Slip plane, a1 is the major semi axiS, aa is the minor semiaxis, EjJ\" is the permuta tion tensor, H(x) is the Heaviside function, and fl(x) is the Dirac delta function" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001101_ac50016a047-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001101_ac50016a047-Figure5-1.png", "caption": "Figure 5. Shape of the Gran\u2019s and EFL plots is influenced by the electrode time response and kinetics of the chemical reaction. Data obtained from (A) 6 min and (B) 48 rnin titrations of 0.125 mmol of thiocyanate in 0.1 M KNO, with 0.05 M AgNO,", "texts": [ " However, at slow titration rates, chemical reactions such as AgBr + Br- --f [AgBr,][AgBr,]- + Br- -, [AgBr,l2- are the major contributors to the nonlinear Gran plot. Neither of these phenomena greatly affect the equivalence point estimation by the EFL technique. Not all electrodes display such a marked difference between the equivalence values obtained by the three methods but, in all cases tested, the EFL technique has the highest precision and accuracy. As an example, consider the titration of 0.125 mmol of thiocyanate solution in a 0.1 N KNOB ISA solution -0.05 N silver nitrate. Figure 5 shows titration curves and Gran plots for total titration times of 6 and 48 min. In spite of the great differences in the Gran plots, the EFL values for various titration times between 6 and 48 min were more precise than those obtained via the inflection point. These results indicate that even for titration systems where the inflection point can easily be utilized for quantitative analyses, the EFL technique can result in a further improvement in precision and accuracy. The use of the present linear model to represent the cell output during the time delineated by the minimum in the EFL is a fair-but not concise-representation of the cell output during that time" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003983_j.eswa.2012.08.005-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003983_j.eswa.2012.08.005-Figure1-1.png", "caption": "Fig. 1. Mass on a moving belt system.", "texts": [ " Section 4 proposes an updating rule for tuning the weights of the neuro-fuzzy system according to the Lyapunov stability theory, which is associated with a PD control scheme with a friction compensation term. Section 5 reports simulation results on a one-dimensional motion dynamics of a mass which moves on a surface with friction to illustrate the effectiveness of our proposed neuro-fuzzy system modeling and active control techniques. Section 6 concludes this work. The free body diagram of a block of mass m, placed on a moving belt and constrained by a spring of stiffness k, is shown in Fig. 1. The non-dimensional equation of motion of a single-degreeof-freedom undamped oscillator with the proposed control is governed by the following differential equation (Hinrichs, Oestreich, & Popp, 1998; Zjinjade & Mallik, 2007): mx00\u00f0s\u00de \u00fe kx\u00f0s\u00de \u00bc Ff \u00f0v\u00de \u00fe uc; \u00f01\u00de where m is the mass of the block, x is the displacement of the mass, uc is the control signal, Ff(v) is the friction force, v is the relative velocity. Let t \u00bc x0s; x0 \u00bc ffiffiffi k m q and _x \u00bc dx=dt \u00bc x0=x0; \u20acx \u00bc d\u00f0 _x\u00de=dt \u00bc x00=x2 0: \u00f02\u00de System (1) can be rewritten as: \u20acx\u00f0t\u00de \u00fe x\u00f0t\u00de \u00bc F\u00f0v\u00de \u00fe u; \u00f03\u00de where v \u00bc v0 x0 _x, v0 is the velocity of the belt, F(v) = Ff(v)/k, u = uc/k", " In this paper, we employ a neuro-fuzzy system as the learner model for modeling the friction force in the system (3). Also, we adopt the framework proposed in Wang and Mendel (1992) and Wang (2003) to extract the fuzzy rules of neuro-fuzzy system. Here, model parameter optimization and verification are associated with the closed-loop system performance. The rest of this section details data collection, the neurofuzzy system description and the model parameter initialization. In this paper, we consider the friction force as a function of the velocity. By employing an experimental system in Fig. 1, some data pairs composed of the velocity and the friction values for a motion object were generated. The velocity values were obtained by using the well-known M/T method based on the encoder signals. However, the friction force value corresponding to a specific velocity could not be evaluated directly. Therefore, an indirect method has to be adopted for problem solving. It is well known that a control force is equal to the friction force whilst the object moves at a constant rate. Hence, we acquired the friction force information through the so-called constant velocity testing method (Armstrong & Canudas De Wit, 1994)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000326_1.1897410-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000326_1.1897410-Figure1-1.png", "caption": "Fig. 1 Typical configuration of self-actuating, traction-drive speed reducer", "texts": [ " Some microscale speed reducers have already been manufactured using gears. This new speed reducer is well suited for microelec- JULY 2005, Vol. 127 / 63105 by ASME hx?url=/data/journals/jmdedb/27807/ on 03/23/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F tromechanical systems MEMS because of the simplicity of its geometry in comparison to that of gears and because of its selfactuating capability. A prototype of this invention has been fabricated and assembled. Most of the parts in this embodiment are shown in Fig. 1. Part A is the mounting plate on which the assembly is mounted. Part B is the back plate, which is securely fastened to part A. The back plate holds a bearing, which defines the axis of rotation for the input and output members. This axis is identified as C in the figure. The input shaft is marked D, and the input member, part E, is rigidly attached to D and rotates with it. There are three pairs of intermediate rollers shown in the Fig. 1, but there could be more or less than three sets. A typical inner roller is labeled F, and G is an outer roller. Inner and outer rollers are held in proximity to one another by the roller plate, H. The outer hole, J, through which the outer roller shaft protrudes, is slightly elongated. This allows the inner and outer rollers to press firmly against one another, generating sufficient friction to transmit torque through the unit. The roller plate, along with the inner and outer rollers, is free to rotate about pin, I", " Also, similar units could be cascaded with the one shown to produce as many different speeds as desired. For the members labeled 2\u20137 in Fig. 3, the speed ratio is given by 7 2 = r2r4r6 r3r5r7 1 where rn is the radius of each roller with n=2,3 , . . . ,7. Members 3 and 4 rotate together as do members 5 and 6. In fact, 3 and 4 are likely to be different sections of the same roller with different diameters, and the same would be true of 5 and 6. On the other hand, if the device is to be used for a single speed reduction, as shown in Fig. 1, then the speed ratio would be 5 2 = r2 r5 2 It should be noted that for this configuration, the output member will rotate in the same direction as the input member. It should also be noted that for the configuration shown in all figures, the input member must rotate counterclockwise to drive the output member through the intermediate roller assemblies. If the input member rotates clockwise, the output member will not rotate because the intermediate roller assemblies are inclined to selfactuate for counterclockwise rotation", " This implies that there is no relative spin between contacting elements about an axis normal to the surfaces as would be present for traction-drive CVTs and also for some fixed-ratio designs. Because of this, very high efficiencies are possible both when the device is used as a traction-drive system with a traction fluid or as a friction-drive system without lubrication. 634 / Vol. 127, JULY 2005 rom: http://mechanicaldesign.asmedigitalcollection.asme.org/pdfaccess.as For the configuration shown in Fig. 1, one-third of the power would be transmitted from the input member to the output member along each of the three intermediate roller pairs. The overall efficiency would then be = T5 5 T2 2 23 where T5 and 5 are the output torque and speed, while T2 and 2 are the torque and speed for the input. As in 7 individual efficiencies can be defined for the torques and speeds = T5,th T 5,th v T2 2 = T v 24 Here T5,th and 5,th are the theoretical output torque and speed assuming no power loss in the device, and T and v are the torque efficiency and speed efficiency, respectively", " For his zero-spin design he measured efficiencies in the 91% to 96% range for well-lubricated conditions using a traction-drive fluid and in the 94% to 99% range for starved lubrication conditions. Clearly, this type of zero-spin design holds great promise for applications where significant power levels are transmitted at fixed speed ratios. Furthermore, the design can be used as the basis for a transmission that can be shifted from one distinct speed to another similar to conventional automotive transmissions. For the configuration depicted in Fig. 1, one would typically start with a desired speed ratio. This would lead to selection of reasonable values for r2 and r5 consistent with Eq. 2 . Then suitable values would be established for r3 and r4. The four radii could then be used to determine from Eq. 4 . Eqs. 6 , 8 , and 9 would be used to evaluate R, 1, and 2. Once all the angles are known, Eq. 17 and 22 could then be used to calculate minimum values for the three coefficients of friction. These results would lead to appropriate material selections for the rolling elements" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000308_0094-114x(87)90058-9-Figure12-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000308_0094-114x(87)90058-9-Figure12-1.png", "caption": "Fig. 12. The solution eight-bar at the six precision postions.", "texts": [ " 10, so that the coupler point p of the eight-bar linkage will pass through six prescribed motion positions pj for j -- 1-6 at the specified angular displacements of ground links 1-2, 4-5, and 7-8 which may be geared together to form a one-degree-of-freedom mechanism. The dimensional synthesis of the eight-bar linkage is accomplished by combining the three solution triads 1-2-3-p, 4-5-6-p, and 7-8-9-p for the specified six precision positions which is given in Fig. 11. The configurations of the solution linkage at the six precision positions are given in Fig. 12. The user may pick any section of the tan(p3/2)-/~2 curve and ask the program to show the corresponding Triad-Burmester curve (Fig. 8). Triad-Burmester curves and solution triads Similar to the application of Burmester curves in the dimensional synthesis of a dyad for four precision positions, there exist Triad-Burmester curves for the dimensional synthesis of a triad for six precision positions. Triad-Burrnester curves consist of three curves which correspond to the three traces of the three joints of all solution triads" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001367_20070822-3-za-2920.00086-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001367_20070822-3-za-2920.00086-Figure2-1.png", "caption": "Fig. 2. Illustration of the LOS angle \u03c8dj .", "texts": [ " In this section we present a controller that solves the control problem stated in the previous section. Since we are dealing with a group of vessels, we will use subscript j to denote the jth vessel. The controller consists of two components. The first component is responsible for the yaw control: \u03c4rj := \u2212 F (uj, vj , rj) + \u03c8\u0308dj \u2212 k\u03c8(\u03c8j \u2212 \u03c8dj) \u2212 kr(rj \u2212 \u03c8\u0307dj), (10) where k\u03c8 > 0 and kr > 0 are control gains and \u03c8dj := \u2212 tan\u22121 (ej \u0394 ) , ej := yj \u2212 Dj , (11) is the so-called Line of Sight (LOS) angle with a design parameter\u0394 > 0. The meaning of this angle is illustrated in Figure 2. The LOS angle corresponds to the orientation of the vessel towards the point that lies at the distance\u0394 ahead of the vessel along the desired path. As follows from (1c) and (3c), controller (10) is a feedback linearizing controller that guarantees that the yaw angle of the vessel \u03c8j exponentially tracks the LOS angle \u03c8dj . As will be shown below, under certain assumptions this controller guarantees that the vessel converges to the desired straight line path corresponding to ej = 0. The idea of using such an LOS guidance for achieving path following of straight line paths has been proposed by Pettersen and Lefeber (2001) for a simplified ship model and by Fredriksen and Pettersen (2006) for a full ship model" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001649_9780470061565.hbb002-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001649_9780470061565.hbb002-Figure1-1.png", "caption": "Figure 1. Schematic of Clark oxygen electrode.", "texts": [ " It also demonstrates how much things have changed, commercially speaking, in recent years, dominated by headlines involving huge business deals. The total world market for biosensors was recently estimated to be nearly $7 billion, of which around 85% involved home blood glucose monitors.2 The father of the biosensor concept can be identified clearly as the late Professor Leland C. Clark Jr. In 1956, Clark published his definitive paper3 on the oxygen electrode, a schematic of which is shown in Figure 1. The Clark electrode was a considerable breakthrough and devices based on this design have remained in production ever since. However, Clark was keen to expand the range of analytes that could be measured in the body and realized that this could be achieved, by modifying Handbook of Biosensors and Biochips. Edited by Robert S. Marks, David C. Cullen, Isao Karube, Christopher R. Lowe and Howard H. Weetall. 2007 John Wiley & Sons, Ltd. ISBN 978-0-470-01905-4. the oxygen electrode, with enzymes. He made a landmark address in 1962 at a New York Academy of Sciences symposium, in which he described \u201chow to make electrochemical sensors (pH, polarographic, potentiometric or conductometric) more intelligent\u201d by adding \u201cenzyme transducers as membrane enclosed sandwiches\u201d" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003643_aim.2011.6027000-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003643_aim.2011.6027000-Figure7-1.png", "caption": "Fig. 7. Model of compartmental", "texts": [ " 6 Concerning the number of fiber reduction of the expansion progression angle. Thus, we assume that the required pressure tends to increase as the number of fibers increases, and i should be as small as possible; we set i to 4 in this study. The condition for determining the axial length of the cylindrical tube is that the fold lines should not occur on the rubber part between the carbon fibers, and the inside of the unit should close completely. First, the condition that fold lines do not occur on the rubber part between the carbon fibers is considered. Fig. 7 shows a schematic of the rubber part divided by the fibers. The divided rubber part of the cylindrical tube is clearly a rectangle. According to the expansion experiment we have performed by changing the aspect ratio of the divided rubber part. To prevent the generation of fold lines on the rubber part, the side parallel to the axial direction of this rectangle must be the long side at the time of pressurization. Therefore, this condition is expressed as where l is the axial length of the cylindrical tube and x is the amount of contraction at the time of pressurization" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002603_iros.2009.5354362-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002603_iros.2009.5354362-Figure2-1.png", "caption": "Fig. 2. Configuration of the crane described by the Denavit-Hartenberg convention with parameters given in Table I.", "texts": [ " The joint variables form the vector of generalized coordinates q = [q1, q2, q3, q4] T for this 4 degree-of-freedom system. The forward kinematics can be conveniently expressed using the Denavit-Hartenberg (DH) convention [9], where each link configuration is represented by the homogeneous transformation Ai = Rotz,\u03b8i Transz,di Transx,ai Rotx,\u03b1i , (1) parameterized by joint angle \u03b8i, link offset di, link length ai, and link twist \u03b1i. In Table I the parameters are provided that describe the configuration of the forwarder crane as depicted in Fig. 2. respect to the base frame of the robot is defined by p0 = x y z = [ I3\u00d73 03\u00d71 ] T 0 4 [ 03\u00d71 1 ] , where T 0 4 = A1(q1)A2(q2)A3(q3)A4(q4) . (2) Inverse kinematics from a configuration of the boom tip to the joint variables can be found as a solution of a set of nonlinear trigonometric equations given by T 0 4 in (2). In our case only the boom-tip position shall be specified along some motion such that corresponding joint variables are computed in closed form by a function q = F (p0, q4) , (3) where q4 is the chosen redundant joint variable" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001071_3.58552-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001071_3.58552-Figure7-1.png", "caption": "Fig. 7 Bill of material clearance.", "texts": [ " Loosening of the front snap was recognized initially but the total method by which it contributed to instability was not seen as a problem from any previous experience. New designs should endeavor to maintain one configuration which can then be analyzed and endurance tested with confidence. 2)Determine the clearance sensitivity of the knife-edges before setting the flowpath. A quasi-steady-state analysis is useful in demonstrating the relative sensitivity of various seal combinations to cavity pressure disturbances. In Fig. 7 the variations in the downstream cavity pressures that result from changing the clearance of the first knife-edge are shown. In Fig. 8 the same clearance variations cause only about 20% of the pressure variation when starting from a new baseline with an increased first knife-edge clearance. Obviously the leakage flow across the seal will also increase in the latter case but only by 4% seal flow (0.1% engine flow). This quasi-steadystate analysis does not constitute an aeroelastic stability analysis, but the results are complementary" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001838_0094-114x(75)90072-5-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001838_0094-114x(75)90072-5-Figure6-1.png", "caption": "Figure 6. Involute system for elliptic gears.", "texts": [ " 1). Any two smaller equal ellipses gl and g2, confocal with the rolling ellipses pl and p: and rigidly connected with them, have the property that a crossed transmission belt slung around them does not disturb the original motion (Fig. 5). This fact gaves the possibility for an apparently new gear system for elliptic wheels which represents a natural generalization of the well-known involute system for circular wheels: The tooth profiles are throughout involutes of the \"base\" ellipses g, and g~ (Fig. 6). Corresponding tooth profiles c,, c2 may be considered as the relative paths of a point C on the straight part of the mentioned transmission belt (Fig. 5). The problem of similar developments for other non-circular wheels remains open. [1] DINGELDEY F., Kegelschnitte und Kegelschnittsysteme. Enzykl. math, Wiss. III C1 (1903). [2] HINKLE R. T., Kinematics o/Machines. Prentice-Hall, Englewood Cliffs (1960). [3] KLEIN F. und BLASCHKE W., Vorlesungen iiber h6here Geometrie. Grundlehr. math. Wiss. 22, (1926)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003587_1.3622200-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003587_1.3622200-Figure4-1.png", "caption": "FIG. 4. IN-718 impellers with laser-consolidated blades built on premachined substrate (left one with as-consolidated surface and the right one after sand blasting).", "texts": [ " Laser consolidation can be performed to complete building up a blade and then another one. However, the completed blade(s) will interfere with the laser beam and powder flow when building other blades. In order to avoid the interference issue, laser consolidation was conducted to build one layer at a time for each blade. Using a 5-axis CNC motion system to deal with large tilt-rotation movement, laser consolidation was successfully conducted to build blades on the premachined substrate to form an integrated impeller. Figure 4 shows an IN-718 impeller (left) with its blades directly built up on a premachine substrate using LC process. The surface finish of the as-consolidated IN-718 blades is about Ra\u00bc 1.6 lm along horizontal direction and about Ra\u00bc 2.94 lm along vertical direction. After sand blasting, laser-consolidated blades and premachined substrate show consistent surface finish (Fig. 4, right side). The mechanical properties of laser-consolidated IN-718 are comparable to that of wrought IN-718. After standard heat treatment, the yield and tensile strengths of the LC IN718 are about 1085 and 1238 MPa, and the elongation is TABLE I. Chemical compositions of alloy powders (wt. %). Alloy C Ni Fe Cr Mo Ta\u00feNb V IN-625 0.03 Bal. \u2014 22.0 9.0 3.7 \u2014 IN-718 0.05 Bal. 17.0 19.0 3.0 5.0 \u2014 H13 0.42 \u2014 Bal. 5.04 1.33 \u2014 1.06 This article is copyrighted as indicated in the article. Reuse of AIP content is subject to the terms at: http://scitation" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.17-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.17-1.png", "caption": "Fig. 14.17 Basic construction of a hydraulic excavator (after [14.16])", "texts": [ "2 \u2022 Tracklaying excavators, with crawlers or rubberbelted tracks\u2022 Mobile excavators, which are equipped with a wheel chassis Tracklaying excavators are designed for low driving speeds of up to 6 km/h. Normally, they just move on construction sites and do not participate in normal traffic. Mobile excavators or wheel excavators, however, do take part in the normal traffic. As is the case in agriculture, driving speeds are increasing considerably, with some machines traveling with a maximum speed of 50 km/h. Figure 14.17 shows the basic construction of a hydraulic excavator. It is equipped with a revolving superstructure hinged to the chassis, which may be of chain or wheel type. The revolving superstructure carries the drive, including the oil and fuel containers as well as the cooling system, the filter, the control valve, the actuators, the driver\u2019s cabin, and the counterweight. The arrangement of these components is determined by the demand for balanced mass distribution, visibility, and ergonomics. Another important component is the rotary transmission leadthrough, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001844_20070625-5-fr-2916.00070-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001844_20070625-5-fr-2916.00070-Figure1-1.png", "caption": "Fig. 1. Lay-out of the \u201cHelicopter\u201d.", "texts": [ " Two motors with propellers mounted on the helicopter body can generate a force proportional to the voltage applied to them. The force, generated by the propellers, causes the helicopter body to lift off the ground and/or to rotate about the pitch axis. All electrical signals to and from the arm are transmitted via a slipping with eight contacts. The system is also equipped with a motorized lead screw that can drive a mass along the main arm in order to impose known controllable disturbances. 1 Quanser Consulting, http://www.quanser.com/choice.asp. Following notation is used through the paper (see Fig. 1): \u03b8(t) is pitch angle; \u03b8\u0307(t) is pitch angular rate; \u03b8\u2217(t) is pitch reference signal; vf (t) and vr(t) are control voltages of the (conditionally) \u201cfront\u201d and the \u201crear\u201d motors; u(t) is pitch torque command signal; w(t) is normal force command signal (used for elevation/travel control); ff(t) and fr(t) are tractive forces of front and rear propellers. The control voltages vf (t) and vr(t) are computed from the command signals u(t), w(t) as follows: vf = 0.5(w + u), vr = 0.5(w \u2212 u). (1) The motor control voltages have saturation level 5 V on magnitude" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.9-1.png", "caption": "Fig. 14.9a\u2013c Cross section through a tractor tyre with rim (after [14.10], description in text)", "texts": [ " Above all, it depends on the slip \u03c3 , as shown in Fig. 14.8. As is shown by this curve, the softer the soil, the lower the propulsion that the wheel can transmit. We can also see that the wheels tend to spin faster and more easily for a flat curve profile. Approximately the following formula describes the behavior up to slip of 30% \u03ba \u2248 \u221a \u03c3 . (14.3) The curves shown in Fig. 14.8 are influenced by three features: \u2022 Tyres and stud deformations\u2022 Soil deformations\u2022 Gliding on the contact surface Tyre Structure Figure 14.9 shows the general structure of a tyre. The tyre consists of a casing and a contact surface. The casing is also laterally covered by rubber material. The wire-wound core keeps the tyre inside the rim\u2019s bead. Nowadays, drive wheel tyres are equipped with a hose in most cases, as the driving power transmission would otherwise be restricted. Consisting of several tissue layers, the casing provides the tyre with stability. According to the casing\u2019s construction, different tyre construction varieties can be divided into the following groups: \u2022 Biased ply construction (Fig. 14.9b): The tissue layers run from one bead to the other with 45\u25e6 staggering. This construction variety is relatively easy, making the tyres cheaper.\u2022 Radial ply construction (Fig. 14.9c): In this tyre type, also called belted tyres, the tissue layers run radially from bead to bead. Around this inner layer, on the contact surface, there is an additional contact surface consisting of several tissue layers (the belt). Thus, the contact surface gains greater stability. Due to the soft casing, the tyre exhibits much greater spring deflection and thus a larger footprint than that produced by a diagonal tyre. Additionally, there is: \u2013 Better force transmission \u2013 Reduced soil pressure \u2013 Reduced roll resistance \u2013 Softer spring deflection of the tractor \u2013 Extended life of the contact surface Force Transmission Between Tyre and Soil Various forces are transmitted from the tyre to the soil" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001266_bfb0039280-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001266_bfb0039280-Figure3-1.png", "caption": "Figure 3, A second example. The variable z is to be contzolled.", "texts": [], "surrounding_texts": [ "The inverse kinematics, i.e. q as function of ~:, is derived using the notat ional conventions in (Craig, 1989). Star t e.g. f rom the Cosine-theorem to first yield q~ as \u00a2osq2 and sin q2 from cos q2 i 2 = ~(=1 + =] - 1 - l 2) -- c(=) and slnq2 = 4 - ~ _- s(z). The final result is q, = At=2(=~, ~:d - At~.~.(z-.(~:), 1 + t~(~)) (6) q~ = A t ~ ( , ( ~ ) , ~(~)) Consider the robot in the singular configuration where the robot is fully s tretched along the z~-axis, as seen in Figure 1 b. We will now s tudy motions inward along the zl-axis from this singular configuration. Such a mot ion is obta ined if z2 = 0 i.e. if the joint angles are related as: sinql + l,in(ql + q2) = 0 (7) A n approz i rna te analys is Before developing an explicit expressions for a mot ion along the z r ax i s , an approximate analysis based only on the leading terms in a Taylor expansion of the kinetics (4) is enlightening. The analysis clearly indicates how different t ime functions i.e. different trajectories, (z~(t), z2(t)) may describe possible motions. Close to the singularity (zt,z~) = (1 + t , 0 ) the angles (qx,q2) a r e small. A Taylor expansion of (4) up to order two gives { ~, = 1 - q[/2 + t [1 - (q, + q~)'/2] \u00f7 O(q') \u2022 ~ = q~ + t(q~ + q~) + O(q ~) (8) Now requiring mot ion along the zi-axis and neglecting higher order te rms gives the constraint t q~ = - Y - ~ q~ (9) and { ~ = I + t - ~ d/2 + o(q~) (10) \u2022 ~ o + O(q~) We thus see tha t it is possible to move approximate ly along the zl-axis close to the singularity, since we have ~,,., --* 0, when q3 - , 0. The analysis so far has been without regard to velocities. Assume now motions along the z r a x i s away from the singularity zt = 1 + t - p ( t ) (11) where p(O) = 0 and p(t) >_. O. Note tha t p(t) = ~ q~/2 + o ( d ) (12) We see tha t linear mot ion in z, i.e. p(t) = t implies q~ ~ 4 f => ~2 ~ ~ ~ co as t --. 0. However, quadrat ic motion, p(t) = t~ implies q2 ~ t =~ ~ ~ 1. The conclusion for this part icular example is tha t constant speed, ~1, leads to infinite joint rates, but tha t there exist softer starts that make the motion possible. To be more formal on the latter statement: The kinematics (4) and the motion { , ql = -T'+\"t \"t (13) q2 =~ gives ~-~ ~ 0 as t -* 0, and all derivatives are bounded. A = i Path fo l lowing The approximate analysis gives a motion that locally starts out along the zl-axis. We will now give a motion that really follows a path with z, = o. We will use the inverse kinematics, eq. (6), with z2 = 0. To give such an example, it is enoush to study q~ e (-~r/2, x/2) and hence to use arctan instead of Atzm2. Introduce the time \u00a2o6q2(t) = ~ . Let the morton dependence t = 2 tan~/2, which yields siaq2(t) = ~ and 1-t' 4 \u2022 be such that it is defined over a time-interval and so that the singularity t = 0 is included, i.e. use e.g. t e [0, I]. Introduce this into the inverse kinematics, eq. (6), to obtain It qlC t) = - arctan 1 + t + (1 - t)t2/4 (14) t q2(t) = arctan 1 - t2/--'---'-~ Note that (14), for small t, simplifies to (13). It is straightforward to verify that (14) has all derivatives bounded and hence is a possible motion along the zraxis. S u m m i n g Up The two-rink robot is a standard object of study, and it is usually claimed that the only motions possible in the singular configuration are those perpendicular to the arm (Asada and Slotine, 1986, pp. 65-66; Craig, 1989, pp. 173-174; Spong and Vidyasagar, 1989, pp. 25-26). The motion (14) is a counterexample to this statement. The interpretation of a singularity for control purposes is thus nontrivial. The kinematics is a non-linear function and the fact that the first order term, the Jacobian, loses rank means that degrees of freedom are lost only in some sense. Usually there are higher order terms that determine the behavior. In particular, in the specific example treated here, we find it reasonable to talk about two degrees of freedom for position control, although one of these degrees of freedom cannot be extended outside the reachable space of the robot and has restrictions on the shape of the velocity profile. 3. Controllability: The Differential Geometry Approach We will now investigate the kinematic singularities with respect to robot controllability as defined by the differential geometry approach. Consider the following nonlinear system m =/(~) + ~ 9,(~)~ (15) where ](z), gi(z) E IR\" are smooth functions belonging to C ~, i.e. with continuous partial derivatives of order k. The vector z e ~t\" describes the state vector and ~ e IR are the system inputs. According to the concept of nonlinear controllability as defined by Herman and Krener (1977), the system (15) is said to be locally controllable if the vector fields gl, jr and their consecutive Lie brackets [jr, gd, [jr, [jr, gd],..., span the whole state space. The condition for a system of the form (15) to be controllable is thus: ranka~(z) = rank (.., g~, .., [L 9d .... [L [jr, ~d],...) = n (16) A Second .Ezample These concepts will be studied first using a second simple example. Consider the system shown by Figure 2, where q is the controlled joint angle and r is the applied torque. The dynamic of the robot is simply described as # = r (17) and the mappings between q ~ z, (q, ~) P-~ ~ a n d (q, ~, ~) ~-~ ~, are given as z = cosq ( 1 8 a ) = -~ sin ~ (ISb) = -~ sin q - ~2 cos q (18c) Since (18a) does not provide a bijcctive transformation between the joint coordinates and the Cartesian coordinates we constrain the joint motions to the set Sq = (q E lit : 0 _< q _ ~r/2} for which the transformation (18a) and its inverse are unique. Note that a kinematic singularity is included in Sq. The Jacobian, sin q, is zero at q = 0, which corresponds to the position z = 1. The dynamics expressed in z can now be obtained by combining (17) with (18a-c). This gives ~2 = -~sina q cosq - (sin q)r = ~-~2z - Jl - z ~ r (19) Introducing the following state-space variables z~ = z and z2 = 3, the system (19) can be rewrit ten as = f(=) + g(=)u (20) with r = u and f(z) and g(z) defined as follows: o o Consider first position displacements excluding singular points, that is in S'~, = {zz ~ IR : 0 ___ z~ < I} and define a corresponding space for the full state vector (z~, z2) as N; = 5'-, x IR. Vector fields f(z) and g(z) arc thus analytic functions in 9\" and hence the system (20) becomes a member of the system class described by Equation (15). The controllability test can thus be performed in n\" by inspecting condition (16). Computing the bracket [A 9] gives: [f,9] = ~ . / - ~ .a = o g 2 . o/2 = = ~ (22) Subsequent brackets evaluate to zero for all =. Thus the condition for the matrix t,~(z) to be full rank is: ! \u00b0 \" I and the system (20) is thus controllable in fl'_, This analysis shows that the controUabi]ity properties of the robot model in the joint space are preserved through the transformation (18) for motions in the set fl\" within which transformation (18 a-b) defines a diffcomorphism. Note that in a larger set including the singularity (z~,z2) = (1, 0), i.e., fz= = fr_ u (1, 0) z, the mapping (18 a-b) no more defines a diffeomorphism. Intuition devclopcd in the prcvious section by analyzing thc kincmatic relations suggests that motion might be possible to and from the singularity provided that ccrtain constraints are imposed on the way that motion is performed. The following trajectory in.the z-space, defined for t e [0, vq~X], confirms the result obtaincd from thc kincmatic relationships of the first example zz = cosat 2, z2 = -2atsincd z, $2 = -4a2t2coszt 2 - 2asinzt z (23) which is one solution of the system (20) with the constant input torque r(t) = 2a and with the initial condition, z(0) = (i, 0) T. Actually (20) also has other solutions like z(t) = z(0), Vt, also for r ~ 0, but they have no physical meaning. They arc consequences of the lack of differential bijcctivity of the relation (18) at the considered initial point. Note that the chosen torque yields \"square motion\" in the joint space, i.e., q(Z) = at 2, just as expected from (17). The single link manipulator will thus move from z1(0) = i to m(x/'~/Ta) = 0 with finite torque r = 2of. This simple example thus also shows that motion with finite torque is possible starting in a singular configuration. The trajectory (23) also implies boundedncss of the vector fields f and g, Vz e [0, v/~], and in'particular also at the singularity~ i.e. by inserting the exphcit time dependency (23) into (21) and (22) and thereafter taking the limit we get lirnt~o f(z(t)) = tlmt_og(z(t)) = limt-o[f, g] = (0, 0) r . The rank condition (lfi) can thus be evaluated to give rank (9, lag] , ' \" ) = 0. In conclusion, the controllability criterion (16) loses rank in the limit despite the obvious controllabihty of (i7)." ] }, { "image_filename": "designv11_20_0000674_0954406042369080-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000674_0954406042369080-Figure4-1.png", "caption": "Fig. 4 Definition of the contact length dCONT, ball COM displacement xB and ball deformation rate _dB during impact", "texts": [ " The concept of momentum flux can be illustrated by considering the impact between a thin-walled spherical membrane (no structural stiffness or internal air pressure) and a rigid surface. The initial velocity of this hollow sphere is VB. At the start of the impact, the material in the initial contact region rapidly decelerates from VB to zero. A reaction force, numerically equal to the rate of change in momentum, acts on the surface. During the compression phase of impact, the sphere can be considered as two separate sections: section 1 continues to move towards the surface while section 2 is in contact with the surface, as shown in Fig. 4. For simplicity it is assumed that section 2 is flat and stationary and therefore remains in contact with the C07703 # IMechE 2004 Proc. Instn Mech. Engrs Vol. 218 Part C: J. Mechanical Engineering Science surface during impact. Furthermore, no energy is stored in section 2 (this assumption will be discussed later). It is also assumed that section 1 is undeformed, and therefore all points on this section move towards the surface with the same velocity _dB1. When a segment of section 1 impacts on the surface it is assumed that the velocity changes from _dB1 to zero, and the size (and mass) of section 2 increases", " Furthermore, immediately after initial impact, the edges of block B momentarily rebound off the surface, but the majority of the block remains in contact. This leads to a non-uniform compression of the block, and thus a non-uniform distribution of the contact forces. These forces cause transverse deformations in the block which lead to a significant proportion of the initial kinetic energy being stored in vibration of the block after impact. Conversely, in orientation A the contact force distribution is effectively uniform owing to the geometry of the block, and therefore negligible transverse vibrations are excited. Referring back to Fig. 4, the motion of section 2 of a tennis ball will be analogous to that of a rubber block landing in orientation B. The only main difference is that the tennis ball material is not perfectly elastic and dissipates a large fraction of its energy during impact. The FE analysis has illustrated that the contact time for a block in orientation B is much less than the duration of the compression phase of the impact between a tennis ball and rigid surface which lasts approximately 2ms. Therefore, during this compression phase, section 2 of the tennis ball will rebound off the surface at a fraction of its inbound velocity" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002090_9780470264003-Figure12.7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002090_9780470264003-Figure12.7-1.png", "caption": "Figure 12.7 DFMA optimum design.", "texts": [ "datum, efficiency using DESIGN FOR MANUFACTURE AND DESIGN FOR ASSEMBLY 279 \u03b7assem.,datum assem.,theoretical assem.,actual = \u00d7 = T T 100 12 163 % \u00d7 = 100 7 362 % . % (12.2) SO this is not an assembly-efficient design. 5. Redo step 4 for the optimum design (with the minimum number of parts) after all practical, technical, and economic limitation considerations. Assume that the bushings are integral to the base and that the snap-on plastic cover replaces standoffs, cover, plastic bushing, and six screws, as shown in Figure 12.7. These parts contribute 97.4 seconds in assembly time reduction, which amounts to $0.95 per hour assuming an hourly labor rate of $35 per hour. Other added improvements include using pilot point screws to fix the base, which was redesigned for self-alignment. A worksheet for the optimum design is given in Table 12.3. 280 MEDICAL DEVICE DESIGN FOR X \u2022 Total actual assembly time, Tassem.,actual = 46 seconds from the DFMA database. \u2022 Total theoretical assembly time, Tassem.,theoretical = 12 seconds (the theoretical number of parts is four, with an average of 3 seconds of assembly time)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001679_gt2008-50305-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001679_gt2008-50305-Figure8-1.png", "caption": "Figure 8. 3-D image of turbine showing asymmetric features creating eccentricity of aerodynamic flow, significant asymmetry highlighted, reflecting the part\u2019s eccentricity and angular offset from plane of the image.", "texts": [], "surrounding_texts": [ "The connection between the turbine rotor and the instability was more ambiguous, as it occurred in only ~15% of the engines. The rotordynamic modeling supported the connection between radial turbine flow induced instability and the resulting vibration measurements, but this would occur more frequently if it was part of the design. The investigation team interpreted this to indicate that the problem was not inherent in the design of the engine and its aerodynamic components. Engines that were apparently identical had nonidentical results. The team initiated an extensive review of the turbine rotor fleet and made comparisons of the good and bad turbine rotors. The turbine rotor of the microturbine is similar to other small turbomachinery products, as seen in Figures 5 and 6. There is a Mar-M 247 cast radial turbine wheel with an Inconel 718 tiebolt inertia welded to the backface, visible in Figure 6. The final machining is completed on this compound, welded assembly. The good and bad rotors, with respect to sub-synchronous vibration, were consistent for meeting the existing drawing tolerances, with the critical machined features held to within 5 \u00b5m (~.0002 in.) and the cast features conforming to typical aerospace tolerances for investment castings. It was then noted that the majority of the bad rotors were received with a high initial unbalance. It must be emphasized that the quality of the balance work and the amount of residual unbalance at time of assembly was identical for all rotors, good or bad. As a temporary measure, initial unbalance limits were applied to all incoming rotors. This was effective in protecting the Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/27/2016 T manufacturing process and the end user, but cost prohibitive due to a high scrap rate of incoming parts. The root cause had to be identified and corrected now that the problem was quarantined. It was theorized that the casting was not well centered in the workholding tool during the machining of the critical datum features used in subsequent machining steps, such as blade profile grinding and location of the pilot fit. This would create an effect where the cast aerodynamic features, as seen in Figure 7, were not symmetric about the rotating center of the part. As the casting, namely the massive hub, was eccentric, the initial unbalance of the worst components was quite high, generally proportional to the centering error. But it was still possible, even with badly unbalanced parts, to measure the machined dimensions to be in print relative to the machining datums. This was because the datums themselves were mis-aligned with the casting center of mass. It was clear at this point that there was a deficiency in the component definition that allowed this to occur. In addition to solving the manufacturing problem, the subsequent challenge was how to quantify the casting-tomachining eccentricity or variation, and how to appropriately impose this requirement on the inspection document (i.e. the blueprint). Attempts were made with CMMs of basic and advanced design to measure the cast surface of the rotor\u2019s hubline between blades with respect to the blade outside diameter profile and the machined pilot diameter, which locates the rotor to the bearing shaft. Due to problems, including overwhelming the memory of the controller of the CMM manufacturer\u2019s 4 Copyright \u00a9 2008 by ASME erms of Use: http://www.asme.org/about-asme/terms-of-use machine, no absolute measurements were made to verify the theory. In the interest of continuous improvement goals, an investment was made in a 3-D optical imaging scanner. With this device, the position of the cast surfaces with respect to the center line of the pilot fit could be made with relative ease. Figures 8 and 9 show the results of a typical bad rotor and good rotor, respectively. These images are actual components captured with the 3-D scanner, then aligned with the solid model using the machined datum features as reference points. The colors show the surface error with respect to the theoretically perfect model. The magnitudes and directions of the displacements are shown in the scale on the right side of the figures. A positive value is away from the surface, in the normal direction. This part is made with an investment cast process. Simply put, a sacrificial mold is created in a ceramic material that can endure the pouring temperatures of the metal. After the part is poured and cooled, the mold is forcibly removed from the metal part that has been created inside. The process may include chipping with hammers, blasting with media, or treating in chemical solutions. Due to the mold removal step, the blade tips in Figure 9 were bent with respect to the nominal blade shape, but evenly and symmetrically. The hub line, and therefore the flow volume of the blade pairs, is also even and symmetric about the rotor. For this reason, this part performed Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/27/201 well with respect to rotordynamics, but most likely had a reduction in aerodynamic performance. The bad rotor clearly shows that the cast surfaces are displaced approximately 750 \u00b5m (~.030 in.) radially from the axis defined by the machined diametral datum. The resulting asymmetry in the flow created the Alford\u2019s-type forces, which destabilized the rotor at the engine\u2019s highest power levels. Asymmetric features as small as 250 \u00b5m (~.010 in.) were found to be significant for some engines, although not in all. Some turbine rotors with small displacements of the cast hubline, and thus relatively low initial imbalance values, were still responsible for high sub-synchronous vibration. Further investigation using the 3-D scanner demonstrated that the rotor\u2019s axis of rotation and its casting axis were mis-oriented. Even though the axes were not highly displaced (they could even intersect), the angle between them was relatively large. This could lead to a similar aerodynamic instability without the very large initial imbalance. It should be noted that, being a complex system, there are some other factors outside of this study which affect the final stability of the complete turbomachine. 5 Copyright \u00a9 2008 by ASME 6 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded From: 6 Copyright \u00a9 2008 by ASME http://proceedings.asmedigitalcollection.asme.org/ on 01/27/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 209187. D" ] }, { "image_filename": "designv11_20_0001198_00368790610651521-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001198_00368790610651521-Figure1-1.png", "caption": "Figure 1 Journal bearing and rotating system frame (r, t)", "texts": [ "1108/00368790610651521] D ow nl oa de d by N ew Y or k U ni ve rs ity A t 0 5: 37 1 6 A pr il 20 15 ( PT ) Assuming that the oil flow in the isotropic porous matrix is laminar, under the hypothesis that the H/D ratio is sufficiently small and that the lubricant fluid is Newtonian, the component W0 is determined by Darcy\u2019s equation (D\u2019Agostino, 1984; Kumar, 1998): W 0 \u00bc fH m 1 R2 \u203a2 p \u203au2 \u00fe \u203a2 p \u203a z2 \u00f02\u00de As : W p \u00bc d h dt \u00f03\u00de by substituting equations (2) and (3) in equation (1) the following fundamental equation for isothermal flow and the unsteady lubrication of the porous journal bearings is achieved (Figure 1): 1 R2 \u203a \u203au \u00f0 h3 \u00fe 12fH\u00de \u203a p \u203au \u00fe \u203a \u203a z \u00f0 h3 \u00fe 12fH\u00de \u203a p \u203a z \u00bc 6m v \u203a h \u203au \u00fe 2 dh dt \u00f04\u00de While, in the case of steady lubrication: 1 R2 \u203a \u203au \u00f0 h3 \u00fe 12fH\u00de \u203a p \u203au \u00fe \u203a \u203a z \u00f0 h3 \u00fe 12fH\u00de \u203a p \u203a z \u00bc 6mv \u203a h \u203au \u00f05\u00de Considering the model of infinitely long journal bearings, the term with the axial pressure gradient can be neglected, obtaining the typical one-dimensional form: \u203a p \u203au \u00bc 6mvR2 h 2 hM h3 \u00fe 12fH \u00f06\u00de where with hM is indicated the film thickness at the circumferential coordinate u \u00bc u where the hydrodynamic pressure has the peak value" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000454_b138835-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000454_b138835-Figure7-1.png", "caption": "Fig. 7. Plot of the equilibrium invariant at different times. Time t = 0:1 (top) and t = 0:6 (bottom). Earlier well-balanced scheme (left) and new well-balanced scheme (right). Number of grid points N = 50", "texts": [ " It can be shown easily that f(d) = E(d; z = 0) < E(d; z > 0) for d < dn: (35) Therefore, the starting point of nucleation d = dn obtained in the last section by considering the absolute minimizer is the smallest strain at which there are other local minimizers with the same or lower energies than those of the pure \u201c-\u201d phase. And at d = dn, the phase mixture with z = zn has the same energy as in the \u201c-\u201d phase and all the other phase mixtures have larger energies, i.e., fn := f(dn) = E(dn; z = zn) < E(dn; z 6= 0; zn): (36) Energy penalty, energy barrier and hysteresis in martensitic transformations 95 As shown in Fig. 7a, the phase mixture with z = zn has a lower energy than the \u201c-\u201d phase for d > dn. Therefore, we may expect from considering the local minimizers that the body starts to transform its phase at d = dn and it transforms from z = 0 to z = zn. The stress n0 required for nucleation defined by (28) is essentially the stress needed to eliminate the energy barrier between f(d) andE(d; zn) as made evident in Fig. 7b by consideringE n0d. Once the body has climbed over the energy barrier, it falls into the energy branch E(d; zn) and has a lower stress nzn = E;d (dn; zn) < n0 as defined by (29). Thus, we may call the energy fn at d = dn as defined by (36) the driving force of nucleation. The formation energy of nucleation Ef defined by (30) is just the energy (32) of the local minimizer z = zn at its stress-free configuration. And the 96 Y. Huo, I. M\u00fcller difference between these two energies is fg := fn Em = nzn 2 2\u02db : (37) As discussed in the last section, the body under extension first transforms its phase by nucleation at d = dn and the stress drops from n0 to nzn ", " We consider here a case in which the solution is a small perturbation of a stationary non-static equilibrium. The initial condition is given by Eq. (8), but with q0 = 0:17. The numerical solution for the total height at two different times is reported in Fig. 6, where only 50 grid points are used. The main difference between the earlier wellbalanced scheme and the new one which preserves non-static equilibria is noticeable near the middle of the channel, where the bottom is higher, and the water profile becomes lower. To enhance the effect, in Fig. 7 we report the quantity I(x; t) = q2 2(H B)2 + gH; which is invariant at equilibrium. Notice the spurious effect near the center of the channel in the numerical result obtained using scheme (10), (11). We remark, however, that these effects are rather small, and quickly disappear as the grid is refined. Fig. 8, for example, represents the same quantity with 200 grid points, and the effect is barely noticeable. In all cases, the reference solution is obtained using the earlier well-balanced scheme with 1600 points" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002567_robot.2007.364203-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002567_robot.2007.364203-Figure2-1.png", "caption": "Fig. 2. Force cones as seen on the surface of the unit sphere. The picture does not preserve the linearity. It rather illustrates the are of interest. (a) F3 and \u03a8(F1 \u222aF2). (b) Enlarging of F1 and F2 by the half angle of F3, i.e., the Minkowski\u2019s sum of each cone.", "texts": [ " The question is whether \u2212F3 has non-boundary intersection with \u03a8(F1 \u222a F2). Observed that the question is equivalent to the question whether a vector n3 lies strictly inside \u03a8(F \u2032 1 \u222aF \u2032 2), where F \u2032 i is a the cone Fi the half angle of which is increased by \u03b83, the half angle of F3. Since we concern only the direction of a force cone, let us represent a force cone by its intersection with a unit sphere. When we look at the surface of the sphere, the intersection of \u03a8(F1 \u222a F2) resembles a racetrack (see Fig. 1). Fig. 2 displays the transformation. From Fig. 2, a vector n3 is strictly inside a \u03a8(F \u2032 1 \u222a F \u2032 2) only when n3 lies strictly inside F \u2032 1, F \u2032 2, or the area in between, which is represented by a shaded region in the figure. Checking whether a vector is inside a circular cone is merely checking whether angle between the vector and the cone\u2019s axis is smaller than the half angle of the cone. The remaining problem is to check whether the vector lies strictly inside the area in between. The area in between can be represented by the intersection of four half spaces, each of which has its normal vector pointing inward to \u03a8(F \u2032 1 \u222a F \u2032 2)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002684_225704-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002684_225704-Figure1-1.png", "caption": "Figure 1. (a) Geometrical parameters and (b) characteristic magnetic configurations of a bi-phase micro- and nanotube.", "texts": [ " Besides, we can expect the appearance of new magnetic properties, like the dipolar magnetic bias responsible for the giant magnetoimpedance behaviour of amorphous microwires [12]. The possibility of achieving such a, spin-valve-like, hysteresis loop is very attractive because of potential applications to sense magnetic fields in magnetic recording systems. The purpose of this paper is to investigate the magnetic ordering of bi-phase tubes. Our cylindrical particle, of length L , consists of two magnetic tubes separated by a nonmagnetic spacer. The internal (external) tube is characterized by its outer Ri(Re) and inner ai(ae) radii, as illustrated in figure 1(a). Magnetic configurations of the bi-phase tubes can be identified 0957-4484/07/225704+05$30.00 1 \u00a9 2007 IOP Publishing Ltd Printed in the UK by two indices [ci, ce], where ci, ce = 1 or 2 denote the magnetic configurations of the internal and external magnetic tubes, respectively (see figure 1(b)). We adopt a simplified description of the system, in which the discrete distribution of magnetic moments is replaced by a continuous one characterized by a slowly varying magnetization density M( r). The total energy E [ci,ce] tot is generally given by the sum of three terms, the magnetostatic E [ci,ce] dip , the exchange E [ci,ce] ex , and the anisotropy contributions. Here we are interested in soft or polycrystalline magnetic materials, in which case the anisotropy contribution is usually disregarded [4, 6]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000326_1.1897410-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000326_1.1897410-Figure5-1.png", "caption": "Fig. 5 Geometry and forces for one typical outer intermediate roller", "texts": [ " 14b , and the resulting pair of equations are solved to determine expressions for F3 and F5 F3 = r5 \u2212 r2 R cos 2 + cos 1 + 2 \u2212 1 sin 1 + 2 f 15a F4 = r5 \u2212 r2 R cos 1 \u2212 cos 1 + 2 + 1 sin 1 + 2 f 15b To prevent slip between the input member and the inner roller as well as between the outer roller and the output member, the following relationships must be satisfied: f 3F3 f 4F4 16 where 3 is the coefficient of friction between the input member and the inner roller and 4 is the coefficient of friction between the outer roller and the output member. Combination of these inequalities with both of Eqs. 15 , will establish the following minimum coefficients of friction to avoid slip 3 sin 1 + 2 r5 \u2212 r2 cos 2 + cos 1 + 2 \u2212 1 R JULY 2005, Vol. 127 / 633 hx?url=/data/journals/jmdedb/27807/ on 03/23/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F 4 sin 1 + 2 r5 \u2212 r2 R cos 1 \u2212 cos 1 + 2 + 1 17 Figure 5 shows the forces acting on one intermediate roller. Angles shown in the figure are related by the following expression: = \u2212 1 + 2 18 Once moments are summed about point O in the figure, the friction force between the intermediate rollers is determined f34r4 \u2212 f4r4 = 0 f34 = f4 = f 19 Next, forces are summed along the line of action of F34 F34 \u2212 F4 cos \u2212 f sin = 0 20 To determine F34 Eq. 15b is combined with Eq. 20 F34 = r5 \u2212 r2 R cos 1 + 1 cos \u2212 cos sin 1 + 2 f 21 Then, to prevent slip between the inner and outer rollers, the coefficient of friction between the two must satisfy the following relationship: 34 sin 1 + 2 r5 \u2212 r2 R cos 1 + 1 cos \u2212 1 \u2212 2 \u2212 cos 22 Thus, to prevent slip at all three contact points involving rolling motion between adjacent members, the three inequalities of 17 and 24 must all be satisfied" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002482_robot.2008.4543557-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002482_robot.2008.4543557-Figure1-1.png", "caption": "Fig. 1. World and body fixed coordinate frames", "texts": [ " The velocity and acceleration The authors are with the Department of Mechanical Engineering and Mechatronics, The Paslin Robotics Research Laboratory, Ariel University Center, Ariel 40700, Israel. This work was partly supported by grant 01-99- 08430 of the Israeli Space Agency through the Ministry of Science Culture and Sports of Israel. shiller@ariel.ac.il limits are computed for each plane using a vector-algebraic procedure, and are then combined to yield the total dynamic stability margins. Figure 1 shows the global and vehicle fixed coordinate frames. The vehicle travels along a path parameterized by its arc length s, such that its center of mass along the path is a parametric function of s. The vehicle\u2019s orientation along the path relative to the global frame is described by the rotation matrix R0 B. The vehicle\u2019s angular velocity is expressed by rotations around the vehicle fixed frame, as shown in Figure 1. The vehicle\u2019s angular velocity \u03c9 is computed by differentiating the rotation matrix of its body fixed frame relative to the inertial frame ([15]): S(\u03c9) = R0 B T R\u03070 B (1) where S(\u03c9) is a skew symmetric matrix, consisting of the elements of \u03c9 , and R\u03070 B is approximated by the finite differences of two successive rotation matrices along the path. To facilitate the analysis of the vehicle\u2019s stability, the acceleration of its center of mass, x\u0308, is best described in terms of its speed s\u0307 and acceleration s\u0308 along the path: x\u0308 = s\u0308t+(s\u03072/\u03c1)n (2) where t is a unit vector tangent to the path, n is a unit vector normal to the path and pointing towards the center of path curvature, as shown in Figure 1, and \u03c1 is the radius 978-1-4244-1647-9/08/$25.00 \u00a92008 IEEE. 2301 of curvature. The vehicle\u2019s angular speed and acceleration are expressed as: \u03c9 = \u0398ss\u0307 (3) \u03c9\u0307 = \u0398ss\u0308+\u0398sss\u0307 2 (4) where \u0398s is obtained by differentiating R0 B with respect to s: S(\u0398s) = R0 B T (R0 B)s (5) and \u0398ss is its partial derivative with respect to the arc length s. D. Inverse kinematics Given the terrain map and vehicle\u2019s location and orientation in the the inertial coordinate system, we wish to calculate the vehicle\u2019s orientation and the direction of the ground normals" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000774_1.7143-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000774_1.7143-Figure3-1.png", "caption": "Fig. 3 Stretching of wrinkled membrane.", "texts": [ " This is because the deformation gradient F ignores the out-of-plane deformation of the wrinkled membrane and thus underestimates the length (1 + \u03b2)l of the wrinkled membrane in the w direction (Fig. 2). To recover the uniaxial tension conditions, the deformation gradient F has to be modified according to F\u2032 = (I + \u03b2w \u2297 w) \u00b7 F (12) where I denotes the identity tensor and \u03b2 represents a measure of the amount of wrinkliness. A physical interpretation of the term (I + \u03b2w \u2297 w) is that it stretches the wrinkled membrane surface along to w until its wrinkles just vanish, as shown in Fig. 3. This stretching involves rigid-body movements only because membranes are supposed to possess no bending stiffness in the TF theory. Consequently, strains and stresses in the stretched membrane remain the same as those in the wrinkled membrane. Using the modified deformation gradient F\u2032, the Green\u2013Lagrange strain E\u2032 of the stretched membrane, equivalent to that of the wrinkled membrane, can be obtained as E\u2032 = E \u2032 \u03b1\u03b2G\u03b1 \u2297 G\u03b2 = 1 2 (F\u2032T \u00b7 F\u2032 \u2212 I) = E + EW (13) where EW = 1 2 \u03b2(\u03b2 + 2)w\u0302 \u2297 w\u0302 (14) w\u0302 = w \u00b7 F = ( w\u03b1g\u03b1 ) \u00b7 F = w\u03b1G\u03b1 (15) From Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000401_vetecf.2004.1404946-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000401_vetecf.2004.1404946-Figure3-1.png", "caption": "Fig. 3. Multiple addresses", "texts": [ " 2) Multiple sink associations: It is obvious that since the addressing begins from the sink, the probability of nodes being associated with more than one sink exists. The number of addresses to be stored can be a tunable parameter (depending on the storage and computation constraints) and each node could have more than one address. The advantage of doing so is that each node has multiple options to send its data. This can add robustness to the process of data transfer to the sinks. Consider the scenario shown in Figure 3. The figure shows node 1224 which also has an address 6213. Thus, the node can reach both sink 1 and 6. This could be useful in situations where even though the distance from both sinks is the same, there could be congestion in the direction of one sink and therefore the node could make a decision to send it towards the other sink. 0-7803-8521-7/04/$20.00 (C) 2004 IEEE 45800-7803-8521-7/04/$20.00 \u00a9 2004 IEEE 3) Sending data to the sink: Using the unique addresses leads to a loose routing framework which can enable the transfer of data from source nodes to the sinks" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000688_s00170-003-1853-1-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000688_s00170-003-1853-1-Figure1-1.png", "caption": "Fig. 1a\u2013c. Sizing the depth buffer. a To the tool diameter b To the tool teeth c Zoomed in view of the teeth depth buffers", "texts": [ " Secondly, as a consequence of the first limitation, an extended Z-buffer must store multiple intersection points, whereas a depth buffer need only store a single value. With the aid of the rendering engine, the depth buffer is implemented directly in graphics hardware. While the implementation of the mechanistic model using the depth buffer is discussed extensively in [4], a brief summary is provided here for the sake of completeness. For every position in which the cutting forces are to be determined, a scene is composed of the current tool position, every previous tool position and the stock material. This is shown in Fig. 1a. The reader should note that a rendered tool position consists of the swept sector of each tooth for a given time step. For each scene in the simulation, the depth buffer direction is adaptively set, aligned with the current tool axis. In graphics parlance, this is accomplished by viewing the scene along the tool axis. When a scene is rendered, the graphics engine determines the state of the depth buffer by determining how those objects would be positioned in the scene, relative to the current eye location (the viewing datum)", " Essentially, the difference between the two states determines the in-process chip geometry. 2.1 Improvements in the determination of the metal removal rate, Q The determination of Q is discussed extensively in [4]. However, in that paper, the depth buffer was sized to the tool diameter. While this is an improvement over extended Z-buffer methods, where the Z-buffer is sized according to the stock material, most of the depth buffer is not used for modelling the in process chip geometry. As can be seen in Fig. 1a, most of the buffer holds information about the previous tool positions, leading to an inefficient usage of the depth buffer. An efficient usage of the depth buffer would concentrate a greater majority of the depth buffer elements (sometimes, referred to as dexels) to determining in process chip geometry. In this case, instead of allocating one depth buffer sized to the tool diameter, one could allocate a number of depth buffers equal to the number of tool teeth (Fig. 1b). In this manner, the simulation could effectively \u201czoom in\u201d on the cutting teeth, allowing the simulation to more accurately determine the in process chip geometry (Fig. 1c). Notice the depth buffer sized to the tool diameter and the depth buffers zoomed to the tool teeth are approximately the same size, yet the latter is a more efficient usage. Furthermore, the tool shape used in [4] was a simple flat end mill. This is but one type of tool used in milling. The theoretical 7-parameter (d, e, f , h, r , \u03b1, \u03b2) APT tool is shown in Fig. 2. Through a judicious choice of parameters, this tool may mathematically model flat end, ball nose, bull nose and even more complicated tool shapes" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002515_00423110701810596-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002515_00423110701810596-Figure4-1.png", "caption": "Figure 4. Control volume extracted from the envelope and relative meridian section.", "texts": [ " In Equation (1), T and Tp are the meridian and parallel tensions, respectively, while pi is the inflation pressure. The aforesaid relations can be obtained by solving the two equations system, one differential and the other one algebraical, constituting the local equilibrium conditions. To effect a global analysis, an envelope elementary sector, characterised by amplitude d\u03b8 , is considered. Such sector is isolated around the meridian half-plane associated to \u03b8 and a control volume is isolated from this, as illustrated in Figure 4. According to the sectionising principle, the aforesaid volume is in equilibrium with the forces exercised by the removed part on it. From the first equation in Equation (1), multiplying r d\u03b8 , it is possible to write T r d\u03b8 = \u03c1pi r + r \u2032 2 d\u03b8. (2) From relation (2), it is evident that the force T r d\u03b8 , varying r , balances the force caused by pi pressure acting on the trapezoid, whose height and bases are, respectively, \u03c1, r d\u03b8 and r \u2032 d\u03b8 (Figure 4). Relation (2) gives an immediate interpretation of the T variation with \u03d5: varying \u03d5 the coordinate r changes, and also the area on which the pressure pi acts. Consequently, the force that must be balanced by the tension T changes. Similarly, the Tp expression can be explained by forces of equilibrium considerations regarding the previously described control volume. After a simple passage, the second equation of Equation (1) becomes Tp\u03c1( \u2212 \u03d5) = pi ( \u2212 \u03d5)\u03c12 2 . (3) The Tp\u03c1( \u2212 \u03d5) force balances the pressure force acting on the circle sector (Figure 5) whose radius is \u03c1 and the angle is ( \u2212 \u03d5)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002509_icelmach.2008.4800231-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002509_icelmach.2008.4800231-Figure1-1.png", "caption": "Fig. 1. Computation domain of electromagnetic field", "texts": [ " The 2D finite element models concern: \u2022 the evaluation of rated-load characteristics, of no-load characteristics, of start-up characteristics and of the parameters of the equivalent electrical scheme of the motor corresponding to rated-load operation and to motor start-up; \u2022 the study of instantaneous oscillations of electromagnetic torque dues to the slotting of stator and rotor armatures; \u2022 the study of motor transient and dynamics for no-load startup, for the transition from no-load to rated-load operation and of for the DC braking operation. The geometry of the 2D computation domain of the electromagnetic field, Fig. 1, contains: - the STATOR_CORE, and ROTOR_CORE regions, which are non-conducting (\u03c3 = 0) and magnetic nonlinear (\u03bc(H) > \u03bc0) regions ; - the stranded coil regions U1_PLUS, U2_MINUS, V1_PLUS, V4_PLUS, W2_MINUS, W3_MINUS, which are non-magnetic (\u03bc = \u03bc0) and non-conducting regions, characterizing two groups of coil sides for each phase of the stator winding; - the solid conductor (\u03c3 \u2260 0) non-magnetic regions BAR1, BAR2, \u2026\u2026, BAR10, which represent half of the squirrel cage; - the AIR_GAP and the INSULATION regions, nonmagnetic and non-conducting", " 2, where the voltages of DC sources are V_U = 380 V, V_V =0 V, V_W = 0 V. If the phase stator currents are known, the three voltage sources in Fig. 2 are replaced by current sources. The boundary conditions concerns: \u2022 null value of the local magnetic flux density on the outer and inner contours of the stator and rotor magnetic cores, where the lines of the magnetic field are tangent; \u2022 anti-cyclic periodicity related to the state variable in correspondent points placed on the radial lines that close the computation domain, Fig. 1, because the computation domain contains one of the two machine poles. The initial values of the rotor speed and of the electromagnetic torque are the values of steady state no-load operation for the study of motor dynamics after the rated load apply and the values of steady state rated-load operation, for the study of DC braking dynamics. In the magneto-harmonic models the complex image A(r) of the magnetic vector potential A(r,t) satisfies the equation: curl[(1/\u03bc)curlA] + j2\u03c0f\u03c3A = J1 , (1) where f is the field frequency and J1(r) is the current density in the coil stator regions", " 3: the magnetization inductance Lm, the magnetic losses resistance Rm, the rotor leakage inductance L'\u03c32a and the resistance R'2a corresponding to motor start-up, and L'\u03c32b, R'2b , corresponding to the rated-load motor operation. The magneto-harmonic models of induction motors take indirectly into account the rotor motion, through the input data slip s. The corresponding results do not consider the real variation in time of the rotor position with respect the stator. The results of magneto-harmonic models depend on the relative rotor-stator position in the computation domain, Fig. 1, i.e. on the input data called \u201cinitial position of the rotor\u201d, being more or less different for different values of this data. For the slip value 0.01 and rated voltage and frequency supply, U1n = 380 V, f1n = 50 Hz , the rotor position with respect the reference position from Fig. 1 is parameterized from 0 to 18\u00b0, with the step 0.5\u00b0. The post processing of simulation results gives values of the electromagnetic torque in the range (7.6 \u2026 9.1) Nm. The different values for different rotor \u2013 stator relative positions are the effect of armatures slotting. The average value 8.16 Nm with respect the rotor position corresponds to the rotor shift 2.45\u00b0 with respect the reference position. This result will represents the input data initial rotor position in all magneto-harmonic computations", "0345 corresponds to the value Pn = 7500 W of the motor rated power. Table I contains other rated-load characteristics, the motor rated-speed nn, the rated current I1n, the rated torque Mn and the rated values of the energetic parameters power factor cos \u03d5n and efficiency \u03b7n. The electromagnetic torque \u2013 slip dependence for rated supply (U1n, f1n), Fig. 6, is the result of a parameterized magneto-harmonic analysis for slip s values ranging from 0.001 to 1. The torque values correspond to the computation domain, Fig. 1, they are half of all the machine torque. This electromagnetic torque - slip model offers the values of the starting current I1s and starting electromagnetic torque Mes and the value sm of the critical slip and of the corresponding breakdown torque Mem in Table II. The no-load motor operation corresponds to a slip value much lower than the rated-slip, for example s0 = 0.001. For rated supply (U1n, f1n) this model gives the values I10 = 3.65 A of the no-load current and Pm0 = 85.62 W of the magnetic losses in the stator armature" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000540_b:tril.0000017420.44627.63-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000540_b:tril.0000017420.44627.63-Figure1-1.png", "caption": "Figure 1. Physical configuration of a dynamically loaded journal bearing.", "texts": [ " The effects of circumvolving in the circumferential direction and squeeze film behavior in the axial direction are considered simultaneously, which is deficiency in other research. To take into account the couple stress effects due to the lubricants blended with various additives, the modified Reynolds equation for dynamic loads governing the film pressure is derived using the Stokes constitutive equations. The problem to be considered in this paper is a dynamically loaded journal bearing lubricated with an incompressible couple stress fluid. The physical configuration of the system is shown schematically in figure 1. The momentum equations and the continuity equation of an incompressible fluid with couple stresses are DV Dt \u00bc rp\u00fe B\u00fe 1 2 r C\u00fe r2V r4V: \u00f01\u00de r V \u00bc 0: \u00f02\u00de where the vectors V;B and C represent the velocity, body force per unit mass, and body couple per unit mass, respectively; is the density, p is the pressure, is the classical shear viscosity, and is a new material constant responsible for the couple stress fluid property. It is assumed that the body forces and body moments are neglected, the fluid inertia is small, the fluid film is thin compared with the journal radius, the curvature of the fluid film is neglected" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.40-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.40-1.png", "caption": "Fig. 14.40 Principle of operation of a concrete pump with control valves in the form of flat gates with connecting ports", "texts": [ " The rapid development of concrete pumps started once hydraulic drives were incorporated into their design, supplanting crank power drives. Another breakthrough in the development of concrete pumps was the use of distributing booms, which greatly facilitated the placement of of concrete in the work area by delivering concrete mix directly to the placement area and distributing it there. Currently the most common pump design is a pump with two parallel, alternately operating concrete mix cylinders whose pistons are driven by hydraulic (oil) actuators connected to them in series (Fig. 14.40). The reciprocating motion of the concrete mix pistons is controlled by means of two noncontact limit switches located in the cleaning water tank. Switches of the same type are also used to control the motion of the cutoff valves. An important distinguishing feature of particular piston-type concrete pump designs is the system of valves that controls the flow of concrete mix from the hopper to the cylinders and from the cylinders to the conveying pipe. Besides valves in the form of flat gates, other valve systems, such as the conveying pipe\u2019s swing segment (C- and S-valves), plug valves, etc" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003369_j.mechmachtheory.2010.12.002-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003369_j.mechmachtheory.2010.12.002-Figure2-1.png", "caption": "Fig. 2. Parameters describing position of a torus and a point on its surface.", "texts": [ " The first index (i=1,2) specifies the torus number (this is also the number of the mechanism element), while the second index specifies the number of the coordinate system. A torus is created by rotating Zi2 circle of radius si and moving it away from the axis by ri distance. Point Qi, which is the centre of the torus, was moved away along axis Zi2 by distance qi. Point Si belonging to the torus surface is described by two radii ri and si, and by two angles: \u03c3i\u2014 rotation of main radius ri around axis Zi2 and \u03c6i \u2014 rotation of the radius of torus tube siaround the axis tangent to the main circle of the torus (Fig. 2). Point Ri is a point located on the main circle of the torus. These independent parameters are necessary to determine the toric surface situation in the fixed coordinate system. In the system of auxiliary coordinates QiXi2Yi2Zi2, the components of the vector of point Si located on the surface of the torus will take the following form: Si2 = ri + sisin\u03c6i; 0; sicos\u03c6i; 0\u00bd T \u00f07\u00de The vectors of points Qi and Ri will be respectively: and un Qi2 = 0; 0; 0; 1\u00bd T and Ri2 = ri; 0; 0; 1\u00bd T ; \u00f08\u00de it vector Ni normal to the torus surface at point Si will be: Ni2 = sin\u03c6i; 0; cos\u03c6i; 0\u00bd T : \u00f09\u00de The homogenous coordinate-transformation matrix between Qi2Xi2Yi2Zi2 and Oi1Xi1Yi1Zi1 has the following form: Mi21 = cos\u03c3i \u2212sin \u03c3i 0 pi sin\u03c3i cos \u03b3i cos \u03c3i cos \u03b3i \u2212sin \u03b3i \u2212qi sin \u03b3i sin \u03c3i sin \u03b3i cos\u03c3i sin \u03b3i cos \u03b3i qi cos \u03b3i 0 0 0 1 2 64 3 75 \u00f010\u00de After multiplying the matrix (10) by successive vectors (7)\u2013(9) in the system of coordinates Oi1Xi1Yi1Zi1, the following dependencies were obtained: \u2013 for vector Si of the point located on the torus surface (an equation of the torus in a movable system of coordinates): Si1 = si1x si1y si1z 1 2 664 3 775 = ri + si sin\u03c6i\u00f0 \u00decos \u03c3i + pi ri + si sin\u03c6i\u00f0 \u00desin \u03c3i cos \u03b3i\u2212 si cos\u03c6i + qi\u00f0 \u00desin \u03b3i ri + si sin\u03c6i\u00f0 \u00desin\u03c3i sin \u03b3i + si cos\u03c6i + qi\u00f0 \u00de cos \u03b3i 1 2 664 3 775 \u00f011\u00de \u2013 for vectors of points Qi and Ri respectively: Qi1 = pi \u2212qi sin \u03b3i qi cos \u03b3i 1 2 64 3 75and Ri1 = ri1x ri1y ri1z 1 2 664 3 775 = ri cos \u03c3i + pi ri sin \u03c3i cos \u03b3i\u2212qi sin \u03b3i ri sin\u03c3i sin \u03b3i + qi cos \u03b3i 1 2 64 3 75: \u00f012\u00de Unit vectors will have the following components: Ni1 = ni1x ni1y ni1z 0 2 664 3 775 = sin\u03c6i cos\u03c3i sin\u03c6i sin \u03c3i cos \u03b3i\u2212cos\u03c6i sin \u03b3i sin\u03c6i sin\u03c3i sin \u03b3i + cos\u03c6i cos \u03b3i 0 2 64 3 75; \u00f013\u00de QXi1 = cos \u03c3i sin\u03c3i cos \u03b3i sin \u03c3i sin \u03b3i 0 2 64 3 75;QYi1 = \u2212sin \u03c3i cos \u03c3i cos \u03b3i cos \u03c3i sin \u03b3i 0 2 64 3 75; \u00f014\u00de Ti1 = 0;\u2212sin \u03b3i;cos \u03b3i; 0\u00bd T \u00f015\u00de Unit vectors QXi and QYi constitute the directions of axes which go through point Qi, and unit vector Ti determines the direction of the torus axis. The direction of velocity Vi1 of contact point Si1 (11) in the case of a revolute pair in point Oi1 (Fig. 2) can be determined by calculating the unit vector Ui1 of the direction that goes through the contact point and at the same time is perpendicular to rotation axis Zi1 and to the radius at a given point: Ui1 = ui1x; ui1y; ui1z; 0 h iT = si1xffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi s2i1x + s2i1y q ; si1yffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi s2i1x + s2i1y q ; 0; 0 2 64 3 75 T : \u00f016\u00de The direction of contact point Vi1 velocity unit vector can be calculated as a vector product of unit vector Ui1 (16) and of the unit vector of the direction of rotation axis Zi1 = 0 0 1 0\u00bd T : Vi1 = Ui1 \u00d7 Zi1 = ui1y;\u2212ui1x; 0; 0 h iT \u00f017\u00de If the pair at point Oi1 is a prismatic pair, then the direction of the contact point velocity unit vector will be the same or opposite to the direction of axis Zi1", "2): 1\u2014 dead centre, type A point; 2, 3\u2014 extreme position of an output link, type B points; 4 \u2014 bifurcation point, type C point. Fig. 5c presents the function of link 2 angular velocity depending on the driving link rotation angle \u03b21. The maximum deviation of the given transmission ratio regarding the rotation of the active link 0\u2264\u03b21\u2264180\u00b0 is\u0394k=0.01045. Themaximumpressure angle in the mechanism did not exceed 42\u00b0. As a result, the following values of mechanism synthesis output parameters were obtained: b1\u2212b2=0.75, p1=1.10, q1=1.45, r1=2,55, s1=4.53, \u03b31=39.18\u00b0, p2=14.17, q2=\u22121.18, r2=2.87, s2=2.91, \u03b32=150.21\u00b0 (Fig. 2). A common model was proposed for mechanisms with one degree of freedom, single-loop and spatial mechanisms with a higher pair, and linkage. It can be used to analyse mechanisms taking into account the inaccuracy in the machining of their elements or deformations of their links. However, in the author's opinion, the most important application area of the model presented is for mechanism synthesis. A common and uniform model makes it possible to synthesize any single-loop linkage avoiding the structural analysis stage" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003891_978-94-007-1643-8_22-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003891_978-94-007-1643-8_22-Figure2-1.png", "caption": "Fig. 2 Experimental sit-to-stand transfer trajectories", "texts": [ " In addition, they are independent in order to restore lateral balance. The end effectors consists of handles equipped with a six axis forces/torques sensor that are used in the experiments presented below to evaluate the relevance of the sit-to-stand trajectory. A preliminary analysis of the movement in elderly sit-to-stand clinical trials, with a specific measurement system [11], has allowed to describe the trajectory of the hands during assisted movements. First, the handles must pull slowly the patient to an antepulsion configuration (see Fig. 2). Then, they go from this down position to the up position, used for walking. For each patient, several sit-to-stand transfer trajectories were recorded, some examples of these trajectories are given in Fig. 2. The analysis of these transfer trajectories shows that the global shape of the trajectory is a \u201ds-like\u201d curve and is not directly related to the age or height of the patient but seems to be correlated with its own personal strategy to stand up or sit-down. This seems to reflect invariants of the trajectory generation [14]. The trajectory of the handles has to be similar to the general curve presented in Fig. 3. The term \u201dtrajectory\u201d refers here to Cartesian-space planning of the handle movement" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001215_s00707-005-0298-z-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001215_s00707-005-0298-z-Figure9-1.png", "caption": "Fig. 9. Oriented graph representation of the Stanford/JPL", "texts": [ " The equivalent open-loop chain for the Stanford/JPL finger Therefore A1 \u00bc Ch10 0 Sh10 a1Ch10 Sh10 0 Ch10 a1Sh10 0 1 0 d1 0 0 0 1 2 66664 3 77775; \u00f029\u00de A2 \u00bc Ch21 Sh21 0 a2Ch21 Sh21 Ch21 0 a2Sh21 0 0 1 0 0 0 0 1 2 66664 3 77775; \u00f030\u00de A3 \u00bc Ch32 Sh32 0 a3Ch32 Sh32 Ch32 0 a3Sh32 0 0 1 0 0 0 0 1 2 66664 3 77775: \u00f031\u00de The position and orientation of the end-effector relative to the base frame, the \u00f0x0; y0; z0\u00de coordinate system, can be obtained from the following transformation matrix: T \u00bc A1A2A3: \u00f032\u00de It is known that the transformation matrix T consists of four submatrices: T \u00bc R3 3 p3 1 01 3 1 \" # \u00bc rotation position matrix vector zeromatrix scaling 2 64 3 75: \u00f033\u00de The angular velocity vector of the end-effector x30 relative to the base frame is given as x30 \u00bc x32 Z20 \u00fe x21 Z10 \u00fe x10 Z00; \u00f034\u00de where Z20, Z10, and Z00 are the unit vectors attached to the Z2-, Z1- and Z0-axes and expressed in the base coordinate system shown in Fig. 8. Hence Z00 \u00bc 0 0 1 0 2 66664 3 77775; \u00f035\u00de Z10 \u00bc A1 0 0 1 0 2 66664 3 77775 \u00bc Sh10 Ch10 0 0 2 66664 3 77775 \u00f036\u00de and Z20 \u00bc A1A2 0 0 1 0 2 66664 3 77775 \u00bc Sh10 Ch10 0 0 2 66664 3 77775: \u00f037\u00de In the previous Section the angular velocity of the end-effector, x30, is written in terms of the joint variables x32, x21 and x10, see Eq. (34). In order to express these velocities in terms of the input velocities x40, x50, x60 and x70, an oriented graph is used (Fig. 9). The following fundamental circuit equations can be deduced from the graph: x041 \u00bc x40 x10; x051 \u00bc x50 x10; x061 \u00bc x60 x10; x071 \u00bc x70 x10; x081 \u00bc x0081 \u00bc x81; x021 \u00bc x0021 \u00bc x21; x032 \u00bc x32; x082 \u00bc x81 x21: \u00f038\u00de For the pulley pairs we get the following equations: \u00f04; 2\u00de\u00f01\u00de : x041 \u00bc r24x 0 21; \u00f039\u00de \u00f05; 2\u00de\u00f01\u00de\u00f0 cross-type\u00de : x051 \u00bc r25x 00 21; \u00f040\u00de \u00f06; 8\u00de\u00f01\u00de\u00f0 cross-type\u00de : x061 \u00bc r86x 00 81; \u00f041\u00de \u00f07; 8\u00de\u00f01\u00de : x071 \u00bc r87x 00 81; \u00f042\u00de \u00f08; 3\u00de\u00f02\u00de : x082 \u00bc r38x 00 32: \u00f043\u00de Using the fundamental circuit equations, Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000942_cimsa.2004.1397255-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000942_cimsa.2004.1397255-Figure3-1.png", "caption": "Fig. 3. Single Link Flexible Robotic Manipulator", "texts": [ " By simulating and studying the process it was found that for the range of load torque [0,24], the closed loop system response was best by keeping the range of w, [0,12]. Considering the proposed approach for the following manipulator tip load, the fuzzy mapping is thus used for the generating multiple reference models. IV. SINGLE LINK FLEXIBLE MANIPULATOR MODEL The proposed approach is applied to a manipulator for tracking the angular position based on the desired trajectory for two distinct cases. Figure 3 shows a single link flexible manipulator, which is modeled based on the physical parameters. The model details of the manipulator are discussed in Appendix A. This dynamic model of the manipulator is utilized to track the command signal applied for eight seconds and the tip load is varied arbitrarily at different time instant. The variation in the tip load contributes to the mode swings in thc model. Based on these modes of operation a fuzzy switching scheme is developed using offline design analysis" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001713_crat.19750100903-Figure16-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001713_crat.19750100903-Figure16-1.png", "caption": "Fig. 16. Hypothetical structurc of point singularities in tic schliereii textures. a) s = +l/Z; b) s = -l/Z smec-", "texts": [ " As already has been mentioned, smectic C modifications tend to exhibit schlieren textures with double inversion walls which may be aligned nearly parallel close together. This experimental fact may be considered as a tendency of the C modification to exhibit homogeneous regions with nearly uniform tilt, which are interrupted by changes of the molecule orientation by 360\". As NEHRING and SAUPE conclude from theoretical considerations, points with s = 1/2 only occur at apolar structures as the nematics, when the molecules are aligned parallel to the surface. These points never have been found a t smectic C or B. Figure 16 shows that singularities with s = 112 are not possible because of the quasi polar structure of the tilted smectics. In both points a singularity line or an additional inversion wall would be necessary. This wall would cause 2 dark brushes in addition to the dark brush on the other side of the point, whereas a singularity line should be clearly visible. These optical properties would not be in accordance with the required properties of singularities with s = +1/2. After this discussion of the smectic schlieren textures it is possible to understand that only planar structures of tilted smectics yield schlieren textures and why for instance smectic A or the orthogonal smetic B never occur in a schlieren texture" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000461_iecon.2005.1569118-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000461_iecon.2005.1569118-Figure3-1.png", "caption": "Fig. 3 Block diagram, open loop", "texts": [ " ( ) ( ) ( ) ( ) ( )uu M 1 M 1 * j M,uu M 1 k 0 \u02c6 1P\u0302 u k u k e M \u2212 \u2212 \u2212 \u2126\u03ba \u03ba=\u2212 \u2212 = \u03a6 \u03ba \u2126 = \u22c5 + \u03ba \u22c5\u2211 \u2211 (31) For the cross power density the analogue expression is ( ) ( ) ( ) ( ) ( )uy M 1 M 1 * j M,uy M 1 k 0 \u02c6 1P\u0302 u k y k e M \u2212 \u2212 \u2212 \u2126\u03ba \u03ba=\u2212 \u2212 = \u03a6 \u03ba \u2126 = \u22c5 + \u03ba \u22c5\u2211 \u2211 . (32) The conclusion is that the periodogram-method is based on the same fundamental theory of signal processing as the well known correlogram-method. The application of the method according to Welch, which has been described above, leads to excellent results. The results obtained on a laboratory setup are presented in paragraph IV. Depending on the characteristics of the plant different configurations for accomplishing the measurement of the frequency response are conceivable [1]. Fig. 3 shows the open loop configuration. The PRBS is used as reference value for * qi which is the torque-generating component of the stator current. The first order lag element represents the current control loop. Its output is the actual current iq. \u03c9M is the mechanical speed. Alternatively, the measurement can be accomplished in the closed loop speed control. In fig. 4 the input signal is the multifrequent test function. That means, the PRBS is equal to the reference value for the speed in the control loop", " The cycle time of the control software is 62,5 s.\u00b5 The digital signals are transmitted to the PC via USB-Interface. Since the measurement of the speed has to be carried out with a very high resolution, the analogue signals of a 2048 pulse incremental encoder are used. As an additional information, the resolution reached is approximately 0,5 Mio. inc./rev. Fig. 8 shows the measured frequency response Gmech(j\u03c9). The measurement has been carried out in the open loop control configuration depicted in fig. 3. In this case the tunable mechanics have been adjusted for resf 64Hz.= The cut-off frequencies of the two-masssystem can be identified very exactly. Fig. 9 displays the measurement result for resf 100Hz= in order to point out that the method does not only work in case of one special mechanical system. As the fig. 10 and 11 depict, the measurement of the frequency response can also be accomplished successfully in both of the closed loop measurement configurations. For calculating the physical parameters TM, TL, TC and d of the two-mass-system the least squares algorithm of Levenberg and Marquardt is applied" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000588_1.1800573-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000588_1.1800573-Figure2-1.png", "caption": "Fig. 2 \u201ea\u2026 Schematic of tendon \u201erepresented by the cylinder\u2026 showing the orientation and placement of the imaging planes. The intersection between the slice selected by the p\u00d52 pulse \u201ecoronal plane\u2026 and the slice selected with the p pulse \u201eaxial plane\u2026 results in a volume of excitation shown as a shaded polygon. Signal is collected from this intersection of the two slices and projected onto the axial-plane axis at readout. The axial slice thickness was 2.0 mm for all tendons. The axial slice was taken from the center of the tendon. The coronal slice thickness varied with the diameter of the tendon and was normally \u00c81.0 mm. \u201eb\u2026 A schematic of the position of the coronal imaging slice placed over an image of an unloaded tendon. The position of the slice is chosen such that the upper and lower rim regions are not included in the slice. The imaging parameters for the image are TR\u00c42000 ms, TE\u00c45.5 ms, matrix size \u00c4128\u00c3128, slice thickness\u00c42.0 mm, spectral width\u00c4\u00c116 kHz, and number of signal averages\u00c42.", "texts": [ " mass was maintained on the string in the unloaded state, in order to avoid the tendon going slack and changing its geometry. The long axis of the tendon was positioned horizontally, parallel to the magnet bore and aligned with the main magnetic field. All experiments were performed using a GE CSI-II 2.0T/45 cm imaging spectrometer ~GE NMR Instruments, Fremont, CA! operating at 85.56 MHz for 1H and equipped with a 6280 G/cm gradient insert. The excitation volume used in the generation of the linescan profile was defined by the intersection of two slice-selection planes with differing slice thicknesses @Fig. 2~a!#. Linescan data @12\u201314# were acquired using a DW spin-echo sequence without a phase-encoding gradient in which the p/2 pulse was selective in the coronal plane and the p pulse was selective in the axial plane. The one-dimensional projection was generated by a frequencyencoding gradient. Data were read out along the axis perpendicular to the axial slice ~see Fig. 2!. Because the tendons were relatively uniform along their long axes, a standard 2.0 mm slice thickness was used for the axial slice. In order to determine the placement of the coronal slice, a spin-echo axial scout image of the tendon was first acquired. From this image we determined ~a! the slice offset needed to locate the coronal slice through the center of the tendon and ~b! the slice thickness needed to exclude the top and bottom rim regions. A schematic of the position and thickness of the coronal slice is shown as a white rectangle over the scout image in Fig. 2~b!. By excluding data from the top and bottom rim regions of the tendon, signal from the core region could be obtained without any contamination from rim regions. Once the slice offset and the thickness of the coronal slice were determined, the DW-linescan experiments were performed. Sequence parameters were TR/TE52000/11.0 ms, spectral width 5616 kHz, 128 complex data points, 8 signal averages, FOV 55.0 mm, 9 diffusion gradient amplitudes from 25.0 to 225 G/cm in increments of 25 G/cm, diffusion gradient duration d52 ms and 652 \u00d5 Vol", " The effect of a finite TE value is to reduce the signal from the protons with the shortest spin\u2013 spin relaxation (T2) times, i.e., M 0\u2192M 0e2(TE/T2). Because the TE value was kept constant, this reduction was constant for the data acquired at each diffusion gradient value. Two regions were identified in the M 0 maps, designated as \u2018\u2018core\u2019\u2019 and \u2018\u2018rim\u2019\u2019 and corresponding to the plateau and the peaks, respectively, in Fig. 3. The rim region was chosen as the bright area at the periphery of the tendon @see Fig. 2~b!#. The core region is the dark central part of the tendon in Fig. 2~b!. Example core and rim regions are labeled in Fig. 3~a!. Because the tendon geometry changed slightly under load, separate ranges were used for the two regions for the unloaded and loaded cases. In some tendons there was a layer of material, which appeared to correspond to loose adventitious tissue, along the outside edge of the tendon. This material appeared on the images as well as the M 0 maps. In order to avoid contamination of the data by the Journal of Biomechanical Engineering rom: http://biomechanical" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003988_j.proeng.2012.09.515-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003988_j.proeng.2012.09.515-Figure6-1.png", "caption": "Fig. 6. Double floor of the cabin with the sensor.", "texts": [ " Regarding the provision of the cabin motion there was selected an actuator \u2013 the direct-current motor whereas the transfer of the torque is ensured by the gear with the gear-wheels to the motion screw which in connection with the motion nut bolt transforms the rotational motion to the linear motion of the cabin. The lift cabin is attached with the motion bolt nut and consequently the rotation of the motion screw is resulting in the linear motion of the lift cabin. (fig.4, 5). The function sensing both the passenger presence and the overloading of the cabin is realized through the double floor. The sensing of the weight of the objects inside the cabin is realized by sensing the deformation of the compression spring placed in the subfloor of the cabin (fig. 6). Sensing of the cabin position is realized by the infrared transmissive optical sensor (break-beam arrangement) where the non-transparent element-flag is tightly connected with the cabin and thus, the light flux in the optical sensor through which the cabin with flag currently passes is interrupted during the cabin movement. The didactic model consists of the electronic modules as follows: - Microcontroller modules, - Power transistor switch modules, - Modules of the infrared optical break-beam sensing the lift cabin position, - Modules of the infrared optical break-beam sensing the presence of the persons within the door area, - Modules sensing the cabin occupancy (the weight of the passengers), - Module of the low-pass filter for the cabin smooth (ramp) start by operating PWM (Pulse Width Modulation), - Module of the cabin reverse movement, - Modules of the control panels with the control push buttons, - Module distributing the power supply of the each modules" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000923_tasc.2005.849119-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000923_tasc.2005.849119-Figure8-1.png", "caption": "Fig. 8. Screen shot of tilted helix path with a conductor to be wound (programmed in FeatureCAM).", "texts": [ " Several technologies can be used to manufacture tilted wirewound coils so as to solve the major problem of holding tilted turns in a predetermined position [2], [3]. One of them is to use a tilted helix form that can be easily fabricated by first milling a groove of the correct geometry into a tube. A commercially available numerically controlled machining software package, such as FeatureCAM, may be used to generate the tool cutting paths directly from the path equation (see (1)), with specified parameters of diameter, pitch, tilt angle and number of turns (Fig. 7). Conductor is then wound into the groove (Fig. 8), and any necessary overwrap or fill materials are applied. The wound coil forms are then nested and aligned to one another to form the complete magnet assembly. Spatial position accuracy of 0.1 mm may be achieved by this method. Most probably, the technology used by AML, Inc [3] is similar to the described. Seemingly, it can provide sufficient strength and integrity at the field of some few teslas; in so doing the ultimate in the average current density can be reached if the gap between adjacent turns (see (5)) is held minimal" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000462_978-1-4020-2249-4_38-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000462_978-1-4020-2249-4_38-Figure10-1.png", "caption": "Figure 10 Flat Ring Articulation for the Icosahedral Zig-Zag Linkage", "texts": [ "J6)+ 6(a sincp + scoscp). For the mechanism in Fig. 8 we find F = 6 + 3- 6 x 1 = 3 and for the Do decahedral Zig-Zag Linkage (Fig.9): F = I./; -61 = (20x6+30)-6x31 = -46. For thkosahedral Zig-Zag Linkage we use the Flat Ring Articula tion. With the {a,s,d,t,{3 = arccos[(-3+,j5)/~(20-8,j5) -n/2 - 31, 71\u00b0}and h(cp) = acoscp-ssincp the edge length ofthe icosahedron can be calculated: L( cp) \"\" 2[(1 I 2) + h( cp )cos( n I 5)]/[ cos{3 sine n I 5)] + 6(a sin cp + s cos cp). The mobility of the mechanism shown in Fig.10 (Flat Ring Articulation and 5 Nuremberg scissors) is found by F = ~/; - 6/ .. (10+ 5)- 6x 1 = 9. For the Ikosahedral Zig-Zag Linkage (Figs. 11,12) we obtain with ~/; = j+30= 12xIO+30 and / = 31 from F \"\" ~/; - 6/ = -36, which indicates overconstraintness. This linkage is movable with one degree of freedom which follows from the fact that for any value of the position angle cp the 30 Nuremberg scissors can be built up with the 12 x 5 = 60 gussets. Agraval, S.K., Yim, M.,Suh, W. (2001), Polyhedral Single Degree of Freedom Ex panding Structures, Proceedings of IEEE, Seoul,Korea" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002714_978-3-540-70701-1_1-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002714_978-3-540-70701-1_1-Figure1-1.png", "caption": "Fig. 1. Inertial (XYZ) and body (xbybzb) coordinate systems of the helicopter", "texts": [ " Remotely piloted and autonomous helicopters have been extensively used for applications involving aerial and lateral views including aerial photography, cinematography, inspection and other aerial robotic applications. The maneuverability and hovering ability of helicopters and other VTOL design are main requirements in many of these applications. However, helicopters are more difficult to control than fixed wing aircrafts. In fact, they require critically stabilization loops which are coupled to the displacement behaviors. C. Bonivento et al. (Eds.): Adv. in Control Theory and Applications, LNCIS 353, pp. 1\u201329, 2007. springerlink.com c\u00a9 Springer-Verlag Berlin Heidelberg 2007 Fig. 1 shows the reference systems used in helicopter control. The position and orientation of a helicopter is usually controlled by means of 5 control inputs: the main rotor collective pitch which has a direct effect on the helicopter height (z axis in the X-Y-Z system); the longitudinal cyclic which modifies the helicopter pitch angle (rotation about the yb axis in the xb \u2212 yb \u2212 zb system) and the longitudinal translation; the lateral cyclic, which affects the helicopter roll angle (rotation about the xb axis in xb\u2212yb\u2212zb system) and the lateral translation; the tail rotor which controls the heading (yaw motion) of the helicopter (rotation about the zb axis in xb \u2212 yb \u2212 zb system) and compensates the anti-torque generated by the main rotor; and the throttle control", " Firstly, a general perspective is given. Then, the model described in previous sections is used to design control strategies that illustrate the concepts explained in the mentioned overview. Although a helicopter is a coupled nonlinear multivariable and underactuated system, simplification of some coupling terms leads to a first simplified scheme of main relations between input-ouput variables of Fig. 5, as shown in Table 1. Notice that translational variables are expressed in the body coordinate frame defined in Fig. 1. This set of relationships is the base of the typical control scheme shown in Fig. 10. This control scheme not only takes into account the main relationships in Table 1 but also considers the most important couplings, such us the lateral and longitudinal movement effect on vertical dynamics. Both linear and nonlinear control strategies have been applied to autonomous helicopters. However, even when linear control laws are applied to the inputs defined in Fig. 5, some authors add nonlinear transformations in particular conditions" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001000_tmag.2004.840315-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001000_tmag.2004.840315-Figure2-1.png", "caption": "Fig. 2. Structure of spindle for mobile HDDs.", "texts": [ " 1, the basic challenges for spindle motor manufacturers are currently high-speed performance and durability for high-grade hard disk drives (HDDs); low-cost manufacturing methods for 3.5-in standard HDDs; and compact HDD spindle designs with greater tolerance of temperature and pressure changes for mobile use. This paper reports the basic design of a hydrodynamic bearing with an oil circulation mechanism that has the state function of expelling air and eliminating bubbles inside the bearing cavity, which affects the disruption-free nature of the oil film in the bearing. Fig. 2 shows the structure of a hydrodynamic bearing. The cylindrical sleeve features a bearing hole in which the inserted shaft can rotate freely. The lower end face of the shaft is equipped with a flange, and an approximately 2- m gap is provided between the shaft and sleeve that is filled with oil. The spaces around the upper and lower surfaces of the flange are also filled with oil. While the radial bearing has a pair of herring bone grooves on the inner surface of its sleeve or on the outer surface of the shaft, the thrust bearings also have herringbone grooves, which generate pressure on both sides of the flange [4], [5]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003932_s1000-9361(11)60062-9-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003932_s1000-9361(11)60062-9-Figure1-1.png", "caption": "Fig. 1 Free-floating flexible redundant manipulators.", "texts": [ " The planning problem is transformed into a parameter optimization problem using Gauss pseudospectral method (GPM) [20], and the dynamics constraint is transformed into algebraic equations. GA is first used to locate an approximate solution. This approximate solution is then used as an initial reference solution for direct shooting method (DSM) [21]. The planning problem is also transformed into a parameter optimization problem using DSM, and sequential quadratic programming (SQP) algorithm [22] is applied to locate an accurate solution. The approximate model of system is illustrated in Fig. 1. The manipulator system consists of a rigid space station and a manipulator which is composed of flexible or rigid links connected in the form of an open- loop chain. The rigid space station is named B1. Links of the manipulator are numbered in sequential order from B2 to B4. B2 and B3 are flexible links, and B4 is a rigid link which includes the end-effector and load. As shown in Fig. 1, the inertial reference coordinate frame is 0 whose origin is located at centroid of the whole system C0. The inertial frame and B1 are connected by a virtual joint C1 located at centroid of B1. The centroid of B4 is C5. The joint between Bi 1 and Bi is named Ci. i is the local frame of Bi, and its origin is located at Ci. li is the distance between Ci 1 and Ci. mi is the mass of Bi, Ei the elastic modulus of Bi, and Ii the moment of inertia of Bi. The following assumptions are made to derive a dynamical model: (1) External forces and torques are omitted, and thus momentum conservation and angular momentum conservation strictly hold" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003758_978-90-481-9262-5_31-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003758_978-90-481-9262-5_31-Figure5-1.png", "caption": "Fig. 5 Construction of a 24-PRRP-evolved radially reciprocating motion mechanism.", "texts": [ " The mechanism is a butterfly-shaped overconstrained mechanism with radially reciprocating mechanism in which when vertexes P, P1, P2 and P3 move away from central point O, vertexes V1, V2, V3, V4 move towards O and vertexes V\u2032 1, V\u2032 2, V\u2032 3 and V\u2032 4 away from O and vice versa. It should point out that in the mechanism, the angle \u03b1 between the two prismatic joints in a single PRRP chain is 54.74\u25e6 Considering the kinematic property of this mechanism, it can be used as a multi-functional grasping robot. Similar to construction of the 12-PRRP-evolved mechanism, integrating six of the type II 4-PRRP-combined chains into a cube by aligning edges and merging their vertexes at central point O of the cube, a 24-PRRP-combined chain with a frame is developed in Fig. 5a. Further, by removing the frame of the 24-PRRP-combined chain, a 24-PRRPevolved overconstrained mechanism is obtained in Fig. 5b. This is a mechanism of mobility one and can produce radially reciprocating motion in such a way that vertexes P1, P2, P3, P4, P5 and P6 move outwards point O, vertexes V1, V2, V3, V4, V5, V6, V7 and V8 move towards point O and vice versa. 298 Overconstrained Mechanisms with Radially Reciprocating Motion The 24-PRRP-evolved mechanism is a new overconstrained mechanism based on pure revolute joints that performs radially reciprocating motion. The kinematics of this mechanism can be studied similar to that of the 12-PRRP-evolved mechanism in [3]", " Similarly, as vertexes V1, V2, V3, V4, V5, V6, V7 and V8 are symmetrically arranged around the same spherical, once the position of one of these vertexes pVk are obtained, the position of the other vertexes pVl can be derived by Eq. (6). Further, since the mechanism is evolved from the PRRP chain, the magnitudes of positions of vertexes P1, P2, P3, P4, P5 and P6, and positions of vertexes V1, V2, V3, V4, V5, V6, V7 and V8 must comply with Eq. (4). In addition, the magnitudes of the velocities of vertexes P1, P2, P3, P4, P5 and P6and that of vertexes V1, V2, V3, V4, V5, V6, V7 and V8 must satisfy |p\u0307i|cos\u03b8 = |p\u0307V k|cos(\u03b8 \u2212\u03b1) (7) where \u03b8 is the rotation angle of the link in the mechanism, \u03b1 is the angle as illustrated in Fig. 5a, and i = 1,2, . . . ,6 and k = 1,2, . . . ,8. With the above equation, singularity of the mechanism can be revealed. From the above construction and analysis of the new overconstrained mechanisms obtained, one can find that these mechanisms are characterized by their radially reciprocating motion. This is interesting motion that the mechanisms presented herein can be applied to whereever radial motion is required. Further, from Eq. (4), by considering the kinematic property of a PRRP chain, adding eight vertexes V\u2032 1, V\u2032 2, V\u2032 3, V\u2032 4, V\u2032 5, V\u2032 6, V\u2032 7 and V\u2032 8 and corresponding links to the mechanism to the 12- PRRP-evolved mechanism, a manipulator with eight end-effectors driven by only one actuator can be obtained" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003654_urai.2011.6145931-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003654_urai.2011.6145931-Figure6-1.png", "caption": "Fig. 6. LEGO NXT Robot for autonomous navigation task", "texts": [ " So enhanced NBC/NXC firmware that support multi dimensional array will be used here. It is also important to use float data type on \u03b1 (learning rate) and \u03b3 (discount rate), so their value can be varied between 0 and 1. Experiment data will be saved on NXT brick as text file and it will be transferred to PC after all experiments are finished. Robot used in this research has two ultrasonic sensors (to detect the obstacles), two light sensors (to detect the target) and two servo motors. NXT Brick behaves as \u201cbrain\u201d or controller for this robot. Figure 6. shows the robot. Arena that will be used in experiments have 3 different home positions and 1 target location (by using candle as light source). The general arena is shown in Fig. 7. Beside this arena, some simple structure of some obstacles and target will also be used in order to know characteristics of learning mechanism clearly. 5. Result and Discussion First experiment that will be done is to test robot\u2019s ability in solving autonomous navigation task. Given three different home positions, robot should avoid the obstacles and find the target" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000359_tmag.1986.1064623-FigureI-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000359_tmag.1986.1064623-FigureI-1.png", "caption": "Figure I: Expmentd setup for measuring slider dynamics an a quartz", "texts": [], "surrounding_texts": [ "to 12 kHz with I nm resolutions using a two beam interferometric technique. These experiments used shaker tables and other art8cial stimuli to excite the suspension modes sufkiently to affect the slider, h e r doppler techniques have also been used to measure slider vibrations? These measurements had to be taken at the top of the &der and as such could not directly measure variations in the actual air bearing film thickness or the average flying height of the slider. Presented here is a technique that for the &st time directly measures bath the static flying height and dynamic vibrations of the slider to 100 kHz without any artificial excitation. Measurements of slider vibrations of much less than a nanometer are shown, This accuracy is sufficient to observe both the air bearing and suspension modes as they appear at the &der when dynamic stimulation is only h m disk run out and other sources natural to the rigid disk recording environment* The ability to simultaneously measure both the slider absolute flying height and vibrations dows direct calibration of dynamic data in absolute u n i t s .\nThe slider was flown an a transparent quartz disk, and monochromatic light was focused through the disk onto the slider. The beam reflected off the slider interfered with the beam reflected fiom the surface of the disk, pmducing an intensity that depended on the slider flying height. By recording the intensity of the interference through two fringes the absolute flying height of the slider could be accurately detemined. Dynamic behavior of the slider manifested itself through high frequency variations in the interference intensity and could be measured by a low noise detector. As the flying height changes, the intensity at the detector, I , is given by the equation:\nwhere: Is= intensity of the light reflected from the &der Id= intensity of the light reflected from the disk h = flying height A = wavelength of the dlumhating light % = angle of incidence 9 = phase shift on dection (material dependant)\nThe first argument of the cosine, 41Thcus B / h , is the difference in optical path length between the light reflected o\u20acF the slider and the E&ht reflected off the disk. The second argument, 9, adjusts for the type of media forming the boundaries from which the two reflections take place. A reflection takes place at the boundary between two different media. A reflection at the boundary in the denser medium (higher index of refraction) is phase shifted with respect to a reflection in the less dense medium. For nonabsorhg media t h i s phase shift is rn radians. Thus, if both the slider and disk are made of quartz, the phase shift between reflection of\u20ac the\nExperiment\nA &der was flown on a 14 inch diameter quartz disk at 60 rnls for leading edge measurements and 52 m/s for trailing edge measurements. L i g h t fiom a s tabied mercury arc h p was passed through a monochromator tuned to 545 nm. The intensity of the light interference was measured with a EG&G FND 100 silicon detector and fed into a low noise preamp whose 3dB point fiom DC was at 180 kHz. The preamp output went to an HP 3456A voltmeter and was averaged \u20acor 1.7 seconds to provide information on the average flying height of the slider. For dynamic measurements a Wavetek 5830A FFT analyzer was used to examine the detector signal. Dynamic vibrations of the slider of less than 1 nm peak to peak were measured from 0 to I00 kHz. The noise floor \u20acrum 2 to 100 kHz was below .2 m peak to peak. Figure 1 shows the experimental setup.\ndisk,\nResults\nMeasuring the DC intensity of the trailing edge of the slider for various linear velocities generates the curve shown in Figure 2. As explained above, this curve is used to calibrate the dynamic data in absolute u n i t s . The h e a r velocity at which the dynamic data has the greatest si@ to noise level is where th is curve has the greatest slope. For the trailing edge, the dynamic data was taken at 52 mls. This is near the point of greatest slope. For the leading edge, the optimum point was near 60 mls. As such the leading edge dynamic data was taken at this velocity. The extremely low system noise level allowed both suspension and air bearing resonances to be observed. Suspension resonances can be diskinguished from air bearing resonant modes because they do not change fie-", "1018\nooo Q 0\n0 0\n0\n0 0\n0\n1 -5\n1 .o\n0.5\n0 0\n0 0\n0 1 I J != 0.2 t I 0 1 0\nt 1 10 20 30 40\nFREQUENCY (kHz) LINEAR VELOCITY (m/s)\nFigure 4 Amplitude vs. frequency ptot of the trailing cdge of a slider flying on a quartz disk.\nquency with disk velocity. At lower finear velocities the slider flies lower and the air bearing 3s st8er. This raises the resonant fkquency of the air bearing modes. Similarly, at higher linear velocities t h e resonant frequencies are lower. For the slider measured, all sigdicant resonances were M a w 30 kHz. F i p 3 shows the FFT analyzer output when the detector was lucated to measure the dynamics of the leading edge ofthe slider r d . Bdow 10 kHz are high Q resonances corresponding to suspension modes. At about 15 kHz a low Q air bearing resonance can be seenl This is idenaed\n' as an air bearing mode since its kquency varies with disk velocity.\nless than 0.5 nm codd be measured from DC to 100 kHz. This allowed measurements to be taken of the air bearing formed between a magnetic recording slider and a quartz disk, Air bearing variations were caused by resonant vibrations of the slider and its suspending system. Stimulation of these resonant modes was caused by disk run out and other elements inherent in the magnetic disk recording environment. No extemd stimulation, such as shaker tables, was pruvided.\nThe dynamic behavior of both the leading and trailing edges of a slider ffying over a quark disk was shown. Below 10 kHz, high Q suspension resonances could be seen, The trailing edge air bearing resonance was seen at 25 kHz with an amplitude of 0.5 nrn peak to peak, and the leading edge air bearing resonance was at 15 kHz with an amplitude of 1.5 m peak to peak. These air bearing amplitudes were tau low to be easily measured by other techniques. n 6\nE c W 4\n2\nThe authors would like to thank Pad Ruscitto and Ken Koga for help in developing the test stand hardware and Tim O'Suliivan for many disctlssions involving analysis a\u20ac the data.\n* IBM Magnetic Recording Institute ** Presently at Rockwell Science Center, Thousand Oaks, CA\nI I I r 1\n0 10 20 30 40 FREQUENCY (AHr) Referencrss\n[ 11 C. L,in and R.F. Sullivan, \"An Application of W h i t e Light Interferometry in Thin Film Measurementsx, IBM J Res Develop, pp. 269-276, May 1972. [Z] J.M. Fleischer and C. Lin, \"Mared Laser Interferometer for Measuring Air-bearing Separation\", IBM J Res Develop, pp. 529-533, November 1974.\n131 A. Nigam, \"A Visible Laser Interferometer for Air Bearing Separation Measurement to Submicron Accuracy\", Tram of ASME, 104, pp.\n[4] K. Tanka, Y. Takeuchi, S. Terashima, TI Odaka and Y . Saitoh, *Measurements of Transient Motion of Magnetic Disk Slider\", KEEE Tram Mugn, M A G Z U , 5, pp. 924-926, September f984. [SI Y. Mizoshita, K. Aruga, T. Yamada, pDynamic Characteristics of a Magnetic Head Slider\", IEEE Tram Mugn, MAG-21, 5, pp. 1509- 151 1, September 1985. 141 D.K. Mu, G. Bouchard, D.R Bogy and FIE. Take, cDynamic REsponse of a Winchester Type Slider Measured by Laser Doppler Interferometry\", IEEE T r m Magn, MAG20,5, pp. 927-929, September 1984.\n60-65, J~IIUXY 1984.\nFigure 3: Amplitude YS, frequency ptut of the leading edge of a dider flying on a quartz disk.\nFigure 4 shows the measwed spectrum at the trailing edge of the slider\nrail. Changes in the slider flying height at the trailing edge can effect the magnetic performance of the disk drive, since the readlwrite element is placed at this edge. The suspension resonances are seen again quite clearly as low k q u e n c y (below 10 kHz) fluctuations in the slider flying height. They occur at the same frequencies as the leading edge, but their amplitudes are si@cantly smaller, In addition an air bearing resonance is seen at around 25 kHz. The trailing edge air bearing resonance is at a higher frequency than the leading edge air bearing resonance, but with a lower amplitude. The low amplitude of t h e trailing edge slider fluctuations causes only small effects on the readlwrite process.\nSummary\nA technique for measwing the thickness of an air bearing has been presented. T h i s technique used two beam interferometry. The signal to noise ratio was large enough that, fox the first time, air bearing thickness variations of" ] }, { "image_filename": "designv11_20_0001091_0021-9797(78)90255-2-Figure13-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001091_0021-9797(78)90255-2-Figure13-1.png", "caption": "FIG. 13. Dependence of(AR/R), X 10 s on charge in surface oxidation and reduction ofa PbS electrode (dashed lines with circled points) in comparison with the cathodic and anodic charge-potential relations (heavy lines 2 to 6 derived from Fig. 2a) obtained from the cyclic voltammetry data in 0.1 M aq Na2B4OT. Curves 3, 5, and 6 for the (AR/R), data correspond to curves with the same numbers in the charge-potential profiles. Letters A to G correspond to those in Fig. 2a designating ranges of potential sweeps.", "texts": [ ", to 8/z monolayers , where z is the electron number per site involved in the process. Some possibilities for this reaction are suggested in Section 5, but first the reflectance behavior must be discussed in relation to the i -E profiles obtained in stirred and unstirred solutions. 4 . R E L A T I O N T O T H E R E L A T I V E R E F L E C T A N C E ( A R / R ) B E H A V I O R Figures 2b and 13 show (AR/R)~I as a function of potential and, more usefully, as a function of charge passed to various potentials in the sweep, respectively. Figure 13 also shows the corresponding anodic and cathodic charges Q obtained from the i -E profiles of Fig, 2a. The numerical designation of the curves in Fig. 2 is again used for corresponding sweeps in Fig. 13. The (AR/ R)b~ data for the curve drawn with solid dots are taken from anodic sweeps 3 and 6, and the two drawn with open circles are for corresponding cathodic sweeps. If the assumption is made that the reflectance change is proportional to the amount of a specific species generated on the surface with the optical properties of the substrate remaining unchanged, 1 the region between the two adjacent kinks in the (AR/R)Lh vs Q curve of Fig. 13 can be regarded as characterizing one type of surface species, and therefore probably one surface elec- In the case, e.g., of chemisorption of H or metal atoms on Pt this assumption cannot be made because the optical properties of the top layer of the metal substrate can become changed owing to electron withdrawal or donation by the chemisorbed species (8, 9). However, here the surface films generated are substantially thicker than a monolayer. trode process. On the heavily dotted curve, three different regions can be distinguished fairly clearly, taking into account experimental errors", " The potentials of all the kinks are also indicated on the i -E profiles in Fig. 2a, for reference. It is always useful, in AR/R measurements, to relate changes of AR/R to the Journal of Colloid and Interface Science, Vol. 65. No. 1, June 1, 1978 charge Q passed in producing some change of state of the surface by deposition of a species or by oxidation of the surface (10). It is found for the PbS electrode that the value of the integral slope of the plot of (AR/R), vs Q decreases sharply in the region beyond point D (Fig. 13). As shown on the (AR/R), vs Q curves for sweeps 3 and 6 (Fig. 13), the first few data points for the cathodic sweep, following reversal of the anodic sweep at E (Fig. 2a), fall on an extension of the same line as that for the change of (AR/R)o in the previous anodic sweep. This, of course, is due to continuation of the oxidation process on the cathodic sweep until the current reaches the zero-current line and eventually becomes cathodic in the \"cathodic\" sweep. This type of effect is familiar in studies of the formation and reduction of surface oxide on Pt. No data for the variation of (AR/R)t~ with Q, however, are available from the end of the cathodic sweep G (Fig", " Hence the (AR/R), vs Q curve could be traced back almost to the origin in this case, giving a refiectivity close to that of the initial surface at P. Journal of Colloid and Interface Science, Vol. 65, No. 1, June 1, 1978 The appearance of the cathodic (AR/R)j+ vs Q curve for sweep 3 allows the possible interpretation that at the positive end potential of this sweep the surface contains two types of species corresponding to regions AB and BC in Fig. 2a on the anodic current profile and on the corresponding (AR/R)IL curve as a function of Q in regions AB and BC of Fig. 13 and in the cathodic direction in regions CX and CP of this figure. The species corresponding to the anodicsweep region BC is reduced in an i -E profile which overlaps that for reduction of the species produced in the anodic sweep region AB (see cathodic sweep curves 2, 3, and 4 in Fig. 2a). In fact, the cathodic sweep curve 4 in Fig. 2a begins to show a resolvable second peak. In the (AR/R)jj vs Q lines of Fig. 13, a region CX can, in fact, be distinguished from the remaining region XP cor responding to the two processes referred to above for the anodic sweeps. [(AR/R) changes in cathodic sweeps are designated by double or triple arrows, while those for the anodic sweeps are shown with single arrows in Fig. 13.] In the region PA in the anodic sweeps, there is very little change in (AR/R)~, although a significant charge is passed. This is connected with initial dissolution of the surface region of the PbS indicated by the rotating electrode results. This situation is similar to that found in the initial stages of anodization of Pb metal in CI- or SO42- solutions (11). Beyond peak A, however , a passivation process sets in and (AR/R)~ becomes appreciably more negative with further passage of the charge over region AB in Figs", " Lines cl, fl and dl represent data obtained in the anodic peak (A~) region of sweeps c, f, and d, while cz, f2, and d~ are those in the plateau region of the corresponding three i -V profiles of Fig. 5. The gradual shift of the lines c~, fl, and d~ to a common stable line, c2, f2, and d~, made up from the data of all three runs in the plateau region, is obviously significant and implies that a common species is associated with the change of AR/R over the whole of regions A to D in the anodic sweep. For the region of the anodic sweep beyond D in Fig. 2a, the AR/R results shown in Fig. 13 reveal an interesting behavior of the system: a further increase of -(AR/R)It Journal of Colloid and Interface Science, Vol. 65, No. 1, June 1, 1978 \u00d7 10 3 by about four units occurs over the potential range D - E - F (Fig. 13; cf. the i - V profile for this region in Fig. 2a) but, as shown in Fig. 13, does not increase at the same rate with increasing Q as in the region CD. Also, surprisingly, there is no corresponding change in the optical behavior of the surface over the large cathodic peak in the next sweep which appears to correspond electrochemically to reduction of the species generated in the anodic sweep from D, through E to F at the end of that anodic sweep and the beginning of the next cathodic one (region EF). It seems that the anodic current in region DEF must be made up of a component due to extension from D to H in Fig", "4 V, demonstrate quite clearly that the processes in the A2/C2 region are not diffusion-controlled and must therefore involve a surface process. Comparison of Fig. 10 with Fig. 1 or Fig. 11 shows also that the electrochemical behavior in the Az/Cz region, although modified a little by the species involved in the A1 region, is not primarily due to any preceding process in that region. The quantity of the type of product initially formed in the CD region of Fig. 2a, but possibly over the whole range CDEF, can be estimated by producing the line CD in the AR/R relation as an f(Q) in Fig. 13 to the point H on the cathodic direction of the AR/R curve 6 (Fig. 13). The charge corresponding to the distance FH on the Q axis of Fig. 13 would then be the charge that passes in the formation of the product which does not influence AR/R. It is also this species which is reduced without change of AR/R in the large cathodic peak on the next sweep. The remainder of the anodic charge in the range D to F, i.e., corresponding in Q to the difference in H from D of DH in Fig. 13, is to be attributed to production of the species, which produces apparently very little change in the optical properties of the surface. If only the A2 region is scanned in the reflectance experiment, there are significant changes of (AR/R)Ij in both the anodic and cathodic sweeps (Fig. 10b). The small or zero change of (AR/R)~j in the cathodic sweep, when species generated from both the A1 and A2 regions are reduced, should therefore probably be attributed to the presence of two species whose effects on (AR/ R)L ~ are in opposing directions, rather than to the presence of one species which has no effect on (AR/R)~ I, since the rotating electrode experiment indicates that currents in the A2/ C2 regions are associated mainly with a surface process" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000602_s1526-6125(05)70092-x-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000602_s1526-6125(05)70092-x-Figure4-1.png", "caption": "Figure 4 Mechanical Work Done During Each Cycle of Welding", "texts": [ " The ultrasonic mechanical energy, subsequently dissipated as thermal energy, is determined by the computed local cyclic stress, , and strain, , distributions (in all locations of the welded material volume) as well as by the dry friction shear, x, and component slip, ex, at the interface surface. Heat is generated (a) locally in the foil volume, (Qv(x, y, z), in W/m3), by inelastic hysteresis (due to high loading rate effects) and plastic deformation, and (b) at the interface surface, (Qs(x, y), in W/m2), by friction, during each cycle at the ultrasonic frequency, N, as shown in Figure 4. Q x y z N dv eq eq, ,( ) = ( )\u222b \u03c3 \u03b5 \u03b5 (3) Q x y N de N e des x x x z x x,( ) = ( ) = ( )\u222b \u222b\u03c4 \u03b5 \u00b5\u03c3 (4) As a result of this thermal loading, the foil volume displays conductive heat transfer under the ex- ternal boundary conditions of surface contact conduction to the sonotrode on the top and the anvil on the bottom surface regions. There is also minor convection through the surface exposed to the ambient. Radiation is not considered in this study as the experiment showed a very modest temperature rise", " Aluminum deforms elastically first, until its yield point is reached. Quasi-static simulation was performed in each elastic and plastic phase. The resulting von Mises stress, eq, plot at 0.175 seconds is shown in Figure 5. The local slip is calculated for each pair of interface nodes, on the top and bottom foils surfaces and directly under the probe, by taking the difference of their displacements in the X direction in Eq. (2). The frictional work done is then calculated for each pair of nodes (for a quarter cycle) from the plot of shear stress versus slip in Figure 4. The simulation was run every 0.025 seconds (500 cycles of ultrasonic welding) by using experimentally obtained friction coefficient distribution, \u00b5, by Gao (1999) to capture the mechanical transients in both elastic and plastic phase. Inside the metal volume, the following results are in agreement with the mechanical analysis by Gao (1999). According to the mechanical model, during each weld cycle of 0.5 seconds, the aluminum foils deform initially in the elastic regime (0\u20130.3 seconds) and then yield (0", " The material is led from the elastic deformation to the plastic flow as the input loading (that is, the normal pressure and horizontal shear) increases. When local friction coefficient \u00b5(x, y) < 0.548, the ultrasonic deformation is in the elastic regime; when \u00b5(x, y) > 0.548, the welded part is undergoing plastic deformation (Gao 1999). The timedependent local friction coefficient distribution was Journal of Manufacturing Processes Vol. 7/No. 2 2005 experimentally obtained by Gao. The stress-strain curve is plotted (for a quarter cycle) and used to calculate plastic work done per cycle in Figure 4. The mechanical work obtained from the previous analysis (regarding the source of heat generation) is applied as the volume and surface heat inputs in Eqs. (3) and (4) to a 2-D thermal model to predict the temperature field transient during the ultrasonic welding. The material type and the thickness of both foils are congruent with the mechanical analysis. In this step, the thermal resistance Rth distribution across the weld interface in the unwelded region was experimentally calibrated, as explained in the section below" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002397_2013.23130-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002397_2013.23130-Figure1-1.png", "caption": "Figure 1. Reciprocating single-blade cutter bar: (1) hydraulic connection , (2) hydraulic motor, (3) dap-joint for the tractor chassis, (4) counterbar, (5) blade, (6) crank disc, (7) connecting rod, (8) counterbar teeth, and (9) blade teeth.", "texts": [ " This will allow us to clarify the causes of system disequilibrium and T 756 TRANSACTIONS OF THE ASABE determine the influence of inertia forces and couples, cutting resistance, and motion irregularity on the disequilibrium. Moreover, it will allow us to conduct a sensitivity analysis of the cutter bar parameters, to highlight their influence on the causes of disequilibrium. Finally, the analysis will allow optimization in relation to the material being cut and the geometric and dynamic characteristics of the cutter bar. Figure 1 shows a typical single-blade cutter bar. The carrying frame, integral with the toothed counterbar, is fixed on the tractor chassis and acts as a guide for the motion of a reciprocating steel plate on which the tempered steel trapezoidal teeth are bolted or welded. The counterbar teeth divide the shrub into sheaves to facilitate cutting, and they act as a counterpart to the cutting action of the blade teeth. The drive system, an eccentric crank-conrod mechanism, is connected to a hydraulic motor that takes power from the tractor hydraulic circuit (Kanafojski and Karwowski, 1976)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003381_iecon.2010.5675353-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003381_iecon.2010.5675353-Figure6-1.png", "caption": "Fig. 6. Slave of 16-DOF telesurgical forceps robot", "texts": [ " The 2-DOF haptic forceps robot is mounted to the end effector of YASKAWA Electric Corporation\u2019s 6-axis industrial robot arm MOTOMAN-HP3J. This industrial robot has 6-DOF. 3-DOF positions; x-y-z, and 3-DOF orientations; pitch-yaw-roll. The straight 2-DOF haptic forceps robot is mounted to right arm, and the rotary 2-DOF haptic forceps robot is mounted to left arm. The end effector of 4-DOF can obtain haptic information. The master side of the 16-DOF telesurgical forceps robot is shown in Fig. 5, and the slave side of the 16-DOF telesurgical forceps robot is shown in Fig. 6. Fig. 7. Bilateral control system III. Bilateral control system based on acceleration control In section III-A, the acceleration control based on disturbance is explained. In section III-B, the 4ch bilateral controller is explained. DOB estimates disturbance force f\u0302dis as eq. (3). f\u0302 dis = gdis s + gdis f dis (3) If the cutoff frequency in the DOB is infinite, acceleration response equals to acceleration reference x\u0308re f i (i = m, s). The distrubance observer is also utilized as a reaction force observer to estimate the externed force without force sensor [8]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003284_s12239-010-0071-8-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003284_s12239-010-0071-8-Figure2-1.png", "caption": "Figure 2. Developed vane type hydraulic pump.", "texts": [ " This paper will describe the development of the electric motor-driven pump unit and IPMSM for EHPS system, introduce the equivalent magnetic circuit method to analyze and design the IPMSM, and finally provide experimental results of the EHPS system. 2. DESCRIPTIONS OF EHPS COMPONENTS 2.1. Hydraulic Pump The vane type pump, which has a complicated structure, is used for the EHPS system. This pump has several advantages such as high efficiency, low acoustic noise and small pulsation. The pump has 100 bar of pressure at the relief condition. A suction port and a discharge port are integrated into the pump. The suction port opens when the pressure is too high and all of the oil comes out through the discharge port as shown in Figure 2. Figure 3 displays the pump module integrated with the developed pump motor and its vane pump. The ECU commands a suitable hydraulic pressure based on the actual motor speed, which is obtained from the hall sense in the pump motor, the car speed and the turn speed of the steering wheel. Figure 4 shows the block diagram of the ECU. The pump motor is an IPMSM with a six poled rotor. The speed of the motor is controlled by a current vector control module using a maximum torque-per-current (MTPC) control (Hur, 2008)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001225_iembs.2006.259668-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001225_iembs.2006.259668-Figure2-1.png", "caption": "Fig. 2 Concept design of microrobot", "texts": [ " The paper is organized as follows: In the following section, the locomotive mechanism of the proposed microrobot will be explained. Section III will introduce the fabrication of the proposed microrobot and the control system. Through the various in-vitro and in-vivo experiments in section IV and V, the feasibility of the microrobot was verified and the effects of the design parameters were illustrated. Finally, concluding remarks will be drawn in Section VI. T 1-4244-0033-3/06/$20.00 \u00a92006 IEEE. 2211 II. LOCOMOTION MECHANISM First of all, the concept design of the microrobot is shown in Fig. 2. The proposed microrobot consists of a linear actuator which comprises micro motor and lead screw, an inner cylinder, an outer cylinder, multiple legs and robot outer body. The functions of the above elements are as follows: The linear actuator can move the inner cylinder backward and forward; The inner cylinder has grooves and there is some clearance between the grooves and the legs. Owing to the clearance, the inner cylinder makes the legs rotate and moves with the legs and the outer cylinder; The outer cylinder is connected with the multi-legs by wired-type pin and is moved inside of the robot outer body; The multi-legs are protruded from the robot outer body and are folded in the body" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001554_eej.20585-Figure11-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001554_eej.20585-Figure11-1.png", "caption": "Fig. 11. Flux lines (Uvolt = 3.31%, s12 = \u20130.03).", "texts": [ " Figures 9(a) and 9(b) show the magnetic flux density distributions at the outer and inner air gaps when Uvolt = 1.99% and s12 = \u20130.03, and Fig. 10 shows the changes in Bgo and Bgi with respect to Uvolt. As can be seen in Fig. 10, Bgi is virtually constant regardless of the level of Uvolt. This means that the negative-sequence rotating magnetic field generated in the stator is for the most part eliminated by the secondary magnetomotive force due to the squirrel-cage rotor. Figures 11(a) and 11(b) show the flux lines in the motor when Uvolt = 3.31% and s12 = 0.03. Figure 11(b) is a visualization of only the negative-sequence rotating magnetic field for Fig. 11(a). In this figure, in order to exclude the PM flux, the PM was analyzed by replacing the flux with air at an equivalent magnetic permeability. It is clear from the results that although some of the negative-sequence rotating magnetic field remains in the motor, the iron core inside the squirrel-cage rotor becomes a magnetic pathway, and most of the flux passes through the PM. Given this, the eddy current on the PM surface due to the negative-sequence rotating magnetic field hardly flows even when operating the PMIG with an unbalanced grid voltage, and the accompanying risk of thermal demagnetization can be considered low" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001879_tt.51-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001879_tt.51-Figure1-1.png", "caption": "Figure 1. Two identically sized rollers in contact (M 1-00)", "texts": [ " One of the best advanced fi nite element software packages of those extensively used for structure analysis of contact problems and dynamic analysis is ABAQUS. Copyright \u00a9 2007 John Wiley & Sons, Ltd. Tribotest 2008; 14: 27\u201342 DOI: 10.1002/tt Using ABAQUS, two cylindrical rollers in contact were modelled. Those rollers were subjected to a combined loading of normal and tangential components. They were subjected to normal loading through two analytically rigid zero thickness sheet plates, as shown in Figure 1, in which two solid identically sized rollers are in contact (Model 1). The load was applied through those rigid bodies to ensure uniform distribution of loading on the roller surface and chosen to be analytically rigid so that they do not affect the results in the contact region between the two rollers. A tangential loading component was also applied to the two cylindrical rollers. The value of that tangential loading was one-third of the normal loading. A suffi cient coeffi cient of friction of 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000936_978-1-4020-4941-5_24-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000936_978-1-4020-4941-5_24-Figure2-1.png", "caption": "Figure 2. Singularity loci and singularity-free workspace for \u03c6 \u2208 [120\u25e6, 240\u25e6].", "texts": [ " Indeed, at a Type 1 singularity, a motor can freely rotate without affecting the pose of the platform. Such a pose would be Type 2 singular only for specific combinations of the activejoint variables that correspond to singular legs. We will conclude this section by proposing a range of orientations for which the singularity-free workspace is sufficiently large. Apparently, this range needs to exclude \u03c6 = 0, and if symmetry is to be observed, the best choice would be a range centered at \u03c6 = 180\u25e6. Figure 2 shows our parallel robot with a series of Type 2 singularity circles corresponding to \u03c6 = 120\u25e6 + k10\u25e6 (k = 0, 1, 2, ..., 12). The dots correspond to Type 1 singularities. The singularity-free workspace for the orientation range \u03c6 \u2208 [120\u25e6, 240\u25e6] is the one that excludes the circle-swept region in Fig. 2. Based on this purely algebraic analysis, it is certainly not obvious, but one can verify by using Eq. (5) that for any Type 2 singular configuration for which \u03c6 = 0, lines OiBi intersect at a common point lying on the circumcircle of the base (see the gray-colored configuration in Fig. 2). In order to study the global behavior of the parallel robot at all singularities, we will analyze its direct kinematic model. Indeed, this approach is the most intuitive one and gives a clear geometric interpretation of all singular configurations. It is usually very difficult or even impossible to follow this approach, but in our case the direct kinematic model is particularly simple. Indeed, whatever the active-joint variables, there is always the trivial solution when the base and platform coincide" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003077_asjc.196-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003077_asjc.196-Figure3-1.png", "caption": "Fig. 3. Missile attitude control system.", "texts": [ " When the missile approaches its pre-defined interception point at an altitude of around 150 km, the air density is low and the aerodynamic forces are difficult to exploit. Specifically, the divert control system (DCS) with thrust, as shown in Fig. 2, is assumed to be located near the missile\u2019s center of gravity and aligned with the two axes, bbx and bbz , perpendicular to the longitudinal axis, bby , of the kill vehicle, to generate motion. In contrast, the attitude control system (ACS) with thrusts, shown in Fig. 3, is located and aligned such that only three pure rotational moments about the principal axes are generated. III. ZERO-SLIDING GUIDANCE LAW A guidance system is generally designed to generate suitable motion commands for a missile to adjust its velocity and trajectory to fulfill the tactical goal. Therefore, the guidance law to be investigated in this work is proposed. In Fig. 4, Bi =[bix biy biz] is the inertial coordinate frame whose origin coincides with the missile\u2019s center of gravity, and Bb=[bbx bby bbz] is themissile\u2019s body coordinate frame" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001427_3.60495-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001427_3.60495-Figure1-1.png", "caption": "Fig. 1 Half the maximum natural turn angle corresponding to the given planet and a given velocity at infinity, V}.", "texts": [ "6 04 95 AUGUST 1975 OPTIMAL TRANSFERS BETWEEN HYPERBOLIC ASYMPTOTES 981 The periapsis position of the hyperbola associated with a given c velocity can be moved arbitrarily both in direction and radial distance at negligible fuel cost. This has two important consequences. The problem as stated becomes planar, and the turn angle may be manipulated. The natural turn angle is simply the scattering angle caused by the vehicle's encounter with the planet. The maximum natural turn angle is that obtained by moving the periapsis radial distance to R, with the trajectory then a grazing hyperbola. Half the maximum natural turn angle is designated a; Fig. 1. It is convenient to introduce L, the escape velocity at the surface of the planet (which can replace /x in specifying the planet) and U, the perivelocity at level R on a trajectory of hyperbolic excess velocity V. The optimal deviation angle corresponding to a given Vl and V2 is a} plus a2. The significance of this angle is evident from consideration of the case in which the required turn angle A is equal to the optimal deviation angle. Then the optimal transfer is composed of two half branches of grazing hyperbolas with a common periapse; Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001689_gt2008-51204-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001689_gt2008-51204-Figure1-1.png", "caption": "Fig. 1 A Beam segment and the analogous motion planes", "texts": [ " The rotor is modeled by circular Timoshenko beam finite elements to account for shear deformation and rotary inertia. Analogy between motion planes is constructed to relax unnecessary complexities arising from the use of a unified rotational direction convention [16] and thus all the required details are incorporated using basic 3 symmetric sub-matrices of the order 4x4. The material damping of the rotor is incorporated through proportional damping. The equation of motion of a rotating Timoshenko beam with shear deformation included, in two perpendicular planes of motion XZ and YZ (Fig.1) can be written as follows: Copyright \u00a9 2008 by ASME ms of Use: http://www.asme.org/about-asme/terms-of-use 2 2 ( ) ( , )S x z x xAG p Z t m Z Z t \u03c1 \u03b1\u2202 \u2202 \u2202\u23a1 \u23a4\u2212 + =\u23a2 \u23a5\u2202 \u2202 \u2202\u23a3 \u23a6 (1a) 2 2 ( ) ( , )S y z y yAG p Z t m Z Z t \u03c1 \u03b2\u2202 \u2202 \u2202\u23a1 \u23a4\u2212 + =\u23a2 \u23a5\u2202 \u2202 \u2202\u23a3 \u23a6 (1b) 2 2 ( ) ( )S T r z xEI AG J J Z Z Z t t \u03b1 \u03b1 \u03b2\u03c1 \u03b1 \u03c9\u2202 \u2202 \u2202 \u2202 \u2202 + \u2212 = + \u2202 \u2202 \u2202 \u2202 \u2202 (1c) 2 2 ( ) ( )S T r z yEI AG J J Z Z Z t t \u03b2 \u03b2 \u03b1\u03c1 \u03b2 \u03c9\u2202 \u2202 \u2202 \u2202 \u2202 + \u2212 = \u2212 \u2202 \u2202 \u2202 \u2202 \u2202 (1d) Here the following relations hold \u03c1 is shear form factor ( 0.75 for circular beam section, [17]); as shown in Fig. 1, moments and rotational displacements are in different directions in both XZ and YZ planes. However, the rotations \u03b1 and \u03b2 as well as shear angle in both planes are such that relation (3) is similar for both the planes. Kinetic and Potential energies T and P for the beam segment with a length l are, 2 2 2 2 0 1 2 l z T x yT m J t t t t \u03b1 \u03b2\u23a7 \u23a1 \u23a4 \u23a1 \u23a4\u2202 \u2202 \u2202 \u2202\u23aa \u239b \u239e \u239b \u239e \u239b \u239e \u239b \u239e= + + +\u23a2 \u23a5 \u23a2 \u23a5\u23a8 \u239c \u239f \u239c \u239f \u239c \u239f \u239c \u239f\u2202 \u2202 \u2202 \u2202\u239d \u23a0 \u239d \u23a0 \u239d \u23a0 \u239d \u23a0\u23a2 \u23a5 \u23a2 \u23a5\u23aa \u23a3 \u23a6 \u23a3 \u23a6\u23a9 \u222b r zJ dz t t \u03b1 \u03b2\u03c9 \u03b2 \u03b1\u2202 \u2202 \u23ab\u239b \u239e+ \u2212 \u23ac\u239c \u239f\u2202 \u2202\u239d \u23a0\u23ad 2 2 0 1 2 l P EI Z Z \u03b1 \u03b2\u23a7 \u23a1 \u23a4\u2202 \u2202\u23aa \u239b \u239e \u239b \u239e= +\u23a2 \u23a5\u23a8 \u239c \u239f \u239c \u239f\u2202 \u2202\u239d \u23a0 \u239d \u23a0\u23a2 \u23a5\u23aa \u23a3 \u23a6\u23a9 \u222b 2 2 s x yAG dZ Z Z \u03c1 \u03b1 \u03b2 \u23ab\u23a1 \u23a4\u2202 \u2202 \u23aa\u239b \u239e \u239b \u239e+ \u2212 + \u2212\u23a2 \u23a5\u23ac\u239c \u239f \u239c \u239f\u2202 \u2202\u239d \u23a0 \u239d \u23a0\u23a2 \u23a5\u23aa\u23a3 \u23a6\u23ad The solution of partial differential equations (1) can be carried out by the finite element technique over the space dimension Z rendering a set of ordinary differential equations, with respect to time only" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001426_017-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001426_017-Figure1-1.png", "caption": "Figure 1. BOD apparatus for marine pollution monitoring. (a) BOD apparatus, (b) inside configuration of the apparatus, (c) configuration of the detection cell.", "texts": [ " The precursors were stirred at 60 \u25e6C for 1 h, and then 200 \u00b5l of a culture containing Bacillus licheniformis, Dietzia maris and Marinobacter marinus (1:2:1, w/w) at a concentration of 2.9 \u00d7 109 cell ml\u22121, was mixed with 200 \u00b5l 8% (w/w) PVA and 200 \u00b5l of the prepared ORMOSIL. The mixture was then spread onto the oxygen-sensing layer to build up the BOD sensing film. The BOD sensing film was dried at room temperature for 24 h and stored in a 100 mg l\u22121 GGA solution. An on-line roboticized apparatus (figure 1(a)) was designed and developed for the BOD determination of seawater. The optical scheme and system principles of the apparatus have previously been described [25]. An optic-chemical BOD sensing film was used to measure changes in fluorescence intensity. The size of the apparatus was 60 (L) \u00d7 40 (W) \u00d7 22 (H) cm3, and it was composed of (1) the sampling part, including peristaltic pump and valves to control the pumping of the recovery solution, the standard BOD solutions for calibration, and the seawater samples; (2) the controlling part, including an engineering manipulative computer and software; (3) the measurement part, including a sensing detection cell and a constant temperature controller; and (4) the light to current conversion part, including an optical fibre and a PMT (figure 1(b)). Figures 1(b) and (c) show the inside configuration of the BOD apparatus. The controlling electronic circuit was composed of a computer and a driving electronic circuit. The routes of the sample and standard BOD solutions were controlled by the sampling peristaltic pump and valves. The BOD sensing film was placed in the middle of the detection cell. A pre-warming cell with a volume of 20 ml was designed and set into the bottom of the detection cell. As soon as the solution was pre-warmed to a constant temperature, the detection cell was automatically switched on by an electromotor, and the BOD measurement started" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000955_iros.2003.1250739-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000955_iros.2003.1250739-Figure1-1.png", "caption": "Fig . 1. The made1 of the b i a d robot", "texts": [ " In this application the robot must be able to perform movements which are similar to those of the human, which could range widely. The number of possible movements here is vast and an offline precomputation for even a single robot is problematic. Furthermore., by switching from one precomputed trajectory to another, the movement can become unstable because the distance between two trajectories could exceed the operation region of the local controller. 11. PROBLEM FORMULATION A. Modef of the biped Robot The biped robot is modelled as a chain of four rigid bodies: foot, shank, thigh and mnk, as shown in Fig. 1. The dynamical equations were derived using Kane's formalism [61, and have the following form: M (q)G = f ( q , w ) + T 0-7803-7860-1/03/$17.M)$17.00 2003 IEEE 074 where q = ( q l , q 2 , q3)T is the vector of generalized coordinates, which are angles in ankle, knee, hip joints, and w = ( w I , w ~ , w ~ ) ~ is the vector of the corresponding generalized velocities. The matrix function M(q) rakes into account the mass distribution and the vector function f (q, w ) describes the influence of both, the inertial forces and the gravity. The elements of the vector T are generalized forces applied to the system. For the model considered, these are the torques in the joints. The dot denotes the time derivate in a Newtonian reference frame. The kinematical equations for the model are obviously: q = W (2) The proposed control algorithm will be described using the model shown in Fig. 1. This means that only the movements in the sagittal plane of the robot (XY-plane) will be considered. As will be explained in sec. III, the algorithm can be easily extended for more complicated 3D models. B. Stabilify Condition For the formulation of the stability condition, the concept of the Zero Moment Point (ZMP) [7] , [SI is used. The original definition of the ZMP has been slightly modified, so as to incorporate the definition of the imaginary ZMP 171: Definition 1: The ZMP is an imaginary point where the resulting ground reaction force should be applied, so that the resulting torque imposed on the foot becomes zero", " the center of mass is moving with too high acceleration downward). 111. CONTROL ALGORITHM FOR STABLE MOVEMENT A. The Idea The scheme of the control algorithm is shown in Fig. 3. The movement goal is specified as x,y coordinates of a reference point, which can be chosen arbitrarily. In this study, the pelvis was taken as a reference point (see Fig. I). The block kinematical transfomtion computes the joint coordinates q which correspond to the specified position of the reference point. For the model in Fig. 1, this position is uniquely defined by the angles q1 and 42. To define the angle 43 - trunk orientation - the equation for the position of the center of mass (CM) is used: 5 c m = 9 w ( 5 ) 875 integrators. This modification is performed by projecting the vector 3 onto two planes, which are denoted by the authors as the ZMP- and CM-plane. The notions of these two planes After insening the explicit formulas for F,, and A i , , eq. (4) ~ f 0 r m a t i . m are presented below. can be rewritten as follows: Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002643_978-0-387-25842-3_6-Figure6.12-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002643_978-0-387-25842-3_6-Figure6.12-1.png", "caption": "FIGURE 6.12. (a) Schematic of fabrication process. (b) SEM image of the resulting hexagonally ordered array of carbon nanotubes (from Reference 85).", "texts": [ " The nanotube electrode thus prepared shows a sigmoidal voltammetric response to Ru(NH3)6 3+ solution, characteristic of direct electron transfer with a long cylindrical ultra-microelectrode. The use of ordered (aligned/patterned) carbon nanotubes as biosensing electrodes should provide additional advantages than the aforementioned electrodes based on the nonaligned, randomly entangled carbon nanotubes. Many research groups have used porous membranes (e.g. mesoporous silica, alumina nanoholes) as the template for preparing wellaligned carbon nanotubes with uniform diameters and lengths [25], as exemplified by Figure 6.12 [85]. Without using a template, Dai and co-workers [86, 87] have also prepared large-scale multiwalled carbon nanotubes aligned perpendicular to quartz substrates by pyrolysis of iron (II) phthalocyanine, FePc, under Ar/H2 at 800 \u00b1 1100 \u25e6C. As can be seen in Figure 6.13a, the constituent carbon nanotubes have a fairly uniform length and diameter. The same group has also developed microfabrication methods for patterning the aligned carbon nanotubes with a sub-micrometer resolution (Figure 6.13b) [25, 86, 88]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002558_acc.2009.5160468-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002558_acc.2009.5160468-Figure4-1.png", "caption": "Fig. 4. Quadrotor Visualization. Blue curve is the commanded trajectory, and the red curve is the trajectory followed by the simulated quadrotor. The green bars/arrows at the center for the propeller show the force generated at the actuator.", "texts": [ " The collision response consists of the following information: points of intersection of quadrotor geometry with other objects, penetration depth at those points, the and reaction force vectors. This information is then used to update the dynamics of the simulated 6DoF quadrotor. Figure 5 shows the disturbance rejection of baseline and adaptive controller to a vertical (up) impact. C. Visualization System The visualization system is built on OpenGL graphics libraries. It allows real-time rendering of the quadrotor with trajectory and force parameters overlayed on top, as shown in Figure 4. Real-time visualization is an invaluable tool during hardware-in-the-loop testing. While the physical quadrotor is bolted down to the test stand, the visualization system displays the path of the virtual quadrotor as the simulation runs. This visualization gives the control designer the tools necessary to closely examine the behavior of the controller, allowing them to make adjustments and perform tuning quickly and efficiently. The states of the quadrotor are sent to the visualization module via the TCP/IP layer, which is used to update the graphics at a rate of 30 frames per second" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001915_j.triboint.2008.09.003-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001915_j.triboint.2008.09.003-Figure3-1.png", "caption": "Fig. 3. Seal specimen (NSTCs, model TCL 36*52*10).", "texts": [ " A new seal with no discernible lip abrasion was used for each run to minimize the pumping effect of micro-asperities. From the concept proposed by Kawahara and Hirabayashi [6], the ribbed helix lip seal (the right one in Fig. 2) was inversely assembled on the shaft with the air side of seal in front of the oil supply chamber. A ribbed helix lip seal with 70 ribs slanted at 301 to the horizontal, designed and manufactured by NSTCs (Model TCL 36*52*10) was adopted as the test specimen (as shown in Fig. 3). The seal was installed on a shaft of 36 mm diameter in the test rig. Note that such a test rig is particularly suitable for testing a plain lip seal (without ribs) in which pumping is produced primarily by phenomena occurring within the contact width. In the actual application of ribbed seals, the \u2018\u2018air-side\u2019\u2019 is primarily occupied by air, not oil. When testing a ribbed seal in this test rig, the ribs are totally immersed in oil, and therefore are very effective in pumping oil from the \u2018\u2018air-side\u2019\u2019 toward the oil side" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000459_peds.2003.1283156-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000459_peds.2003.1283156-Figure9-1.png", "caption": "Fig. 9. Phasar diagram of an induction motor with one-pulsc mode.", "texts": [ " As the carrier Gequency of IGBT can be set up to ahout 1 H z , the asynchronous region of IGBT inverter is wider than that of GTO inverter, whose carrier frequency is about 450Hz. Fig. 8. shows the pulse mode of the three-level IGBT inverter, which is used with 700-series Sinkansen. At the one-pulse mode, the value of the output voltage of inverters is not adjustable and some considerations are required. In case of the one-pulse mode, only the phase angle of the inverter voltage can be controlled to adjust the torque. Fig. 9 shows the phasor diagram of an induction motor with one-pulse mode. By reducing Id and the secondary flux (z with large a , large IV can flow and large torque operation is realized, as shown in Fig. 9 (b). Although the torque is the product between the secondary flux and I,, the increase rate of I, is bigger than the reduction rate of the secondary flux. Using components V, and Vq of V I , the value of the voltage =m and a are calculated and assigned. In some practices, these setting are decided by feed forward calculation. In other practice, the reduction of the flux is automated with feedback signal of the comparison between calculated IvII and maximum output voltage of the inverter [SI. I so in0 ISO 200 250 Train Speed (kmih) Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003863_b978-0-12-417049-0.00002-x-Figure2.10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003863_b978-0-12-417049-0.00002-x-Figure2.10-1.png", "caption": "Figure 2.10 Geometry of the goal tracking problem.", "texts": [ "\u201d The Dubins car is obtained when the reverse motion is not allowed in the Reeds-Shepp car, that is, the value v1 52 1 is excluded, in which case v1 5 f0; 1g. Example 2.3 It is desired to find the steering angle \u03c8 which is required for a rear-wheel driven car-like WMR to go from its present position Q\u00f0xQ; yQ\u00de to a given goal F\u00f0xf ; yf \u00de. The available data, which are obtained via proper sensors, are the distance L between \u00f0xf ; yf \u00de and \u00f0xQ; yQ\u00de and the angle \u03b5 of the vector ~QF with respect to the current vehicle orientation. Solution We will work with the geometry of Figure 2.10 [19]. The kinematic equations of the WMR are given by Eq. (2.52). The WMR will go from the position Q to the goal F following a circular path with curvature: 1 R1 5 1 D tg\u03c8 determined using the bicycle equivalent model, that combines the two front wheels and the two rear wheels (Figure 2.10, left). On the other hand, the curvature 1=R2 of the circular path that passes through the goal, is obtained from the relation (Figure 2.10, right): L=25 R2sin\u00f0\u03b5\u00de that is: 1 R2 5 2 L sin\u00f0\u03b5\u00de \u00f02:57a\u00de To meet the goal tracking requirement the above two curvatures 1=R1 and 1=R2 must be the same, that is: 1 D tg\u03c85 2 L sin\u00f0\u03b5\u00de Therefore: \u03c85 tg21 2D L sin\u00f0\u03b5\u00de \u00f02:57b\u00de Equation (2.57b) gives the steering angle \u03c8 in terms of the data L and \u03b5, and can be used to pursuit tracking of goals (targets) that are moving along given trajectories. In these cases the goal F lies at the intersection of the goal trajectory and the look-ahead circle. To get a better interpretation of (2.57a), we use the lateral distance d between the vehicles orientation (heading) vector and the goal point, which is given by (Figure 2.10): d5 L sin\u00f0\u03b5\u00de Then, the curvature 1=R2 in Eq. (2.57a) is given by: 1 R2 5 2d L2 This indicates that the curvature 1=R1 of the path resulting from the steering angle \u03c8 should be: 1 R1 5 2 L2 d \u00f02:57c\u00de Equation (2.57c) is a \u201cproportional control law\u201d and shows that the curvature 1=R1 of the robot\u2019s path should be proportional to the cross track error d some look-ahead distance in front of the WMR with a gain 2=L2. The general 2-input n-dimensional chain model (briefly \u00f02; n\u00de-chain model) is: _x1 5 u1 _x2 5 u2 _x3 5 x2u1 ^ _xn 5 xn21u1 \u00f02:58\u00de The Brockett (single) integrator model is: _x1 5 u1 _x2 5 u2 _x3 5 x1u2 2 x2u1 \u00f02:59\u00de and the double integrator model is: _x1 5 u1 _x2 5 u2 _x3 5 x1 _x2 2 x2 _x1 \u00f02:60\u00de The nonholonomic WMR kinematic models can be transformed to the above models" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003381_iecon.2010.5675353-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003381_iecon.2010.5675353-Figure4-1.png", "caption": "Fig. 4. 6-DOF YASKAWA robot arm", "texts": [ " Because of the thrust wire is flexible, the constraint on the placement of the end effector and the linear motor is relaxed. On the other hand, the flexible actuator achieves weight saving. By using the flexible actuator drive, the movement of linear motor is transformed into open-close movement of forceps. In the flexible actuator, the position and the force is assumed as (1) and (2). x = x\u2032 (1) f ext = f ext\u2032 (2) C. 16-DOF Telesurgical Forceps Robot with Haptics Developed 16-DOF telesurgical forceps robot has 8 DOF in each arm. Fig. 4 shows the YASKAWA robot arm. The 2-DOF haptic forceps robot is mounted to the end effector of YASKAWA Electric Corporation\u2019s 6-axis industrial robot arm MOTOMAN-HP3J. This industrial robot has 6-DOF. 3-DOF positions; x-y-z, and 3-DOF orientations; pitch-yaw-roll. The straight 2-DOF haptic forceps robot is mounted to right arm, and the rotary 2-DOF haptic forceps robot is mounted to left arm. The end effector of 4-DOF can obtain haptic information. The master side of the 16-DOF telesurgical forceps robot is shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003863_b978-0-12-417049-0.00002-x-Figure2.12-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003863_b978-0-12-417049-0.00002-x-Figure2.12-1.png", "caption": "Figure 2.12 (A) Velocity vector of wheel i. The velocity vh is the robot vehicle velocity due to the wheel motion, (B) An example of a 3-wheel setup. Source: http://deviceguru.com/files/rovio-3.jpg.", "texts": [ "52) the kinematic equations of the N-trailer are found to be: _xQ 5 v1 cos \u03c60 _yQ 5 v1 sin \u03c60 _\u03c60 5 \u00f01=D\u00dev1tg\u03c8 _\u03c85 v2 _\u03c61 5 1 L1 sin\u00f0\u03c60 2\u03c61\u00de _\u03c62 5 1 L2 cos\u00f0\u03c60 2\u03c61\u00desin\u00f0\u03c61 2\u03c62\u00de ^ _\u03c6i 5 1 Li L i21 j51 cos\u00f0\u03c6j21 2\u03c6j\u00desin\u00f0\u03c6i21 2\u03c6i\u00de ^ _\u03c6N 5 1 LN L N21 j51 cos\u00f0\u03c6j21 2\u03c6j\u00desin\u00f0\u03c6N21 2\u03c6N\u00de \u00f02:70\u00de which, obviously, represent a driftless affine system with two inputs u1 5 v1 and u2 5 v2 and N1 4, states: _x5 g1\u00f0x\u00deu1 1 g2\u00f0x\u00deu2 We observe that the first four lines of the fields g1 and g2 represent the (pow- ered) car-like WMR itself. The following WMRs will be considered [2,4,11,12,16]: Multiwheel omnidirectional WMR with orthogonal (universal) wheels Four-wheel omnidirectional WMR with mecanum wheels that have a roller angle 6 45 . The geometric structure of a multiwheel omnirobot is shown in Figure 2.12A. Each wheel has three velocity components [16]: Its own velocity vi 5 r _\u03b8 i, where r is the common wheel radius and _\u03b8 i its own angular velocity An induced velocity vi;roller which is due to the free rollers (here assumed of the universal type; roller angle 6 90 ) A velocity component v\u03c6 which is due to the rotation of the robotic platform about its center of gravity Q, that is, v\u03c6 5D _\u03c6, where _\u03c6 is the angular velocity of the platform and D is the distance of the wheel from Q. Here, the roller angle is 6 90 , and so: v2h 5 v2i 1 v2i;roller \u00f02:71a\u00de where vi 5 vh cos\u00f0\u03b4\u00de 5 vh cos\u00f0\u03b32\u03b2\u00de 5 vh\u00f0cos \u03b3 cos \u03b21 sin \u03b3 sin \u03b2\u00de \u00f02:71b\u00de Thus the total velocity of the wheel i is: vi 5 vh\u00f0cos \u03b3 cos \u03b21 sin \u03b3 sin \u03b2\u00de1D _\u03c6 5 vhxcos \u03b21 vhy sin \u03b21D _\u03c6 \u00f02:72\u00de where vhx and vhy are the x; y components of vh, that is: vhx 5 vhcos \u03b3; vhy 5 vhsin \u03b3 Equation (2" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002891_1.4000647-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002891_1.4000647-Figure2-1.png", "caption": "Fig. 2 Normal and tangential components of curve", "texts": [ " Considering a phase diference of /z between orbiting and fixed scrolls, the characterstic angle, c, is defined for a scroll with z wraps as follows: Contributed by the Fluids Engineering Division of ASME for publication in the OURNAL OF FLUIDS ENGINEERING. Manuscript received April 5, 2008; final manuscript eceived October 25, 2009; published online December 15, 2009. Assoc. Editor: hunill Hah. ournal of Fluids Engineering Copyright \u00a9 20 https://fluidsengineering.asmedigitalcollection.asme.org on 05/27/2019 Terms of c = z for z = 1,2,3, . . . 1 3 General Geometrical Model of Scroll Compression Chamber 3.1 Scroll Wrap With General Profile. As shown in Fig. 2, for a point on a continuous curve, assume that its normal angle is the positive angle between the x-axis and the normal to the curve at this point and has a property of superposition. Therefore, an arbitrary point on this curve can be expressed as a complex number by the normal component, Rn, and tangential component, Rt, of its radius-vector as follows: P = Rn exp j + Rt exp j + 2 2 Rt = dRn d According to the general profile theory, there are three types of general profile, they are Rn = c0 + c1 + c2 for type-I of scroll profile Rn = c0 + c1 cos + c2 for type-II of scroll profile 3 Rn = c0 + c1 + c2 2 + c3 3 for type-III of scroll profile As shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000094_12.596317-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000094_12.596317-Figure1-1.png", "caption": "Figure 1 \u2014 Selective Laser Melting: (a) process (b) hatching pitch", "texts": [ " In order to increase the wear resistance of titanium models, surface alloying via laser-gas-nitriding (LGN) is applied. * The 5th International Symposium on Laser Precision Microfabrication, May 1 ]fh]4th, 2004, Nara, Japan Fifth International Symposium on Laser Precision Microfabrication, Edited by I. Miyamoto, H. Helvajian, K. Itoh, K. F. Kobayashi, A. Ostendorf, K. Sugioka, Proc. of SPIE Vol. 5662 (SPIE, Bellingham, WA, 2004) \u00b7 0277-786X/04/$15 \u00b7 doi: 10.1117/12.596317 268 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 05/25/2015 Terms of Use: http://spiedl.org/terms Figure 1 shows the experimental apparatus for laser forming by SLM. The SLM system uses an Nd-yttritrium-aluminum-garnet (YAG) pulsed laser of maximum peak power of 3 kW, pulse width varying from 0. 1 to 10 ms and maximum average power of5O W. The spot diameter at focus position is ofO.75 mm. In the SLM system, the laser head is attached to an x-y table controlled by a computer. The laser beam moves onto the titanium powder bed deposited on a steel substrate to form solidified layers. The substrate is attached to a piston which goes downward of one layer thickness of 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003573_s12239-011-0008-x-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003573_s12239-011-0008-x-Figure9-1.png", "caption": "Figure 9. Conventional transmission error tester.", "texts": [ " This test is inappropriate to determine the cause analysis of hypoid gear whine noise for the following reasons. First, the torque condition is low, and the deflection under the torque is not applied due to the structural limitations of the equipment. In fact, it is impossible to evaluate the transmission error under high torque because the equipment being tested is fixed to each axis by frictional jigs, not by bolts. Second, it is impossible to identify the cause of the gear whine noise because the deflection under the input torque is not applied. Figure 9 shows the engagement error meter that utilizes the revolution angle measurement. The transmission error is calculated from the revolution data of the revolution angle meter installed on the extension line of the pinion and gear axes. To prevent motor-side disturbance, the driving method with couplings and belts is utilized, but the measured value tends to change due to the revolutionary vibration of the transmission error. To make up for this disadvantage, a method of measuring the transmission error using laser sensors was suggested, as shown in Figure 10" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002299_j.compstruc.2006.08.080-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002299_j.compstruc.2006.08.080-Figure3-1.png", "caption": "Fig. 3. Example 2, problem definition.", "texts": [ " It is noted that the established linear system for the case with Nt = 1430 is of order 8586, but that the matrix\u2014in the present implementation\u2014only contains 80,086 non-zero components, or 0.11% filling. Substantially more efficient procedures than an LU factorization of the sparse matrix could be developed, [19]. Further studies of a completely similar system, but with higher numbers of degrees of freedom, showed that the proposed method becomes less favorable for larger systems. Qualitatively, it seemed that the computational work grows as about 3(Nt) 1(Nd)3/2, whereas the other methods demand 3(Nt) 1(Nd)1. A mechanical model was studied, Fig. 3, introducing two degrees of freedom ui, (i = 1,2), cf. [12]. The problem sought the variation of the single control force c1(t), 0 6 t 6 T in the horizontal direction, which would invert the position of the stick from an initial state hanging vertically \u00f0u2\u00f00\u00de \u00bc p; u1\u00f00\u00de \u00bc u01\u00f00\u00de \u00bc u02\u00f00\u00de \u00bc 0\u00de to a target of standing upright \u00f0u2\u00f0T \u00de \u00bc 2p; u1\u00f0T \u00de \u00bc u01\u00f0T \u00de \u00bc u02\u00f0T \u00de \u00bc 0\u00de. The time for the inversion was fixed to T = 0.4 [s], and the data were fixed as m1 = 2 [kg], m2 = 1 [kg], L = 0.2 [m], g = 9.81 [m s 2], which seem to be the values used in [12]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003437_j.1538-7305.1970.tb01816.x-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003437_j.1538-7305.1970.tb01816.x-Figure3-1.png", "caption": "Fig. 3\u2014An example illustrating the global minima and saddlepointe of Eg.", "texts": [ " Consider a transversal equalizer with only one tap Co (such a single tap serves as an automatic gain control and the problem is to jointly set the automatic gain control and carrier phase to minimize the mean-square error). For simplicity, suppose that the term tan\"' [Im (tf'y\"'v)]/[Re (|t'y\"'i\u00bb)] in equation (45) turned out to be zero. Then from the proposition has global minima at \u03c1 = 0, \u00b1 ir \u00b127\u0393 , \u00b7 \u00b7 \u00b7 , and 6o has saddlepoints at \u03c1 = \u00b1 \u03c0 / 2 , \u00b13 i r /2 , - \u00b7 \u00b7. The global minima at \u03c1 = 0 and ir are illustrated by points 1 and 3 in Fig. 3 and the saddlepoint at \u03c1 = v/2 is illustrated by point 2. 1082 THE BELL SYSTEM TECHNICAL JOURNAL, JtH^Y-AUGUST 1970 The curved surface in the figure illustrates the variation of 8o with \u03c1 and \u03b2 \u03b1 . For instance, when \u03c1 is fixed at v, &o varies with \u03b2\u03bf as shown by the convex curve passing through points 4, 3, and 5. (For fixed p, So is a convex function of the tap gains e-s to ey .) As can be seen, point 2 is a saddlepoint because varying \u03b2\u03bf away from point 2 with \u03c1 constant increases So , while varying \u03c1 away from point 2 with \u03b2\u03bf constant decreases So " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002891_1.4000647-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002891_1.4000647-Figure1-1.png", "caption": "Fig. 1 Scroll wraps of multiwrap scroll", "texts": [], "surrounding_texts": [ "G C S\nQ S L 2 7 e\nF w c a l s s o T o o c s c\nK s\n1\ns e t s f s g t fi p t g\n2\nc o i\n1 2 p f f i\nJ r C\nJ\nDownloaded From:\neneral Geometrical Model of Scroll ompression Chamber for croll Fluid Machine\niang Jianguo chool of Mechano-Electronic Engineering, anzhou University of Technology, 87 Langongping Road, 30050 Lanzhou, China -mail: qiangjianguo@lut.cn\nor a scroll fluid machine with an arbitrary number of scroll raps on its individual scroll, the phase difference between adjaent scroll profiles on one scroll is defined as the characteristic ngle of the scroll. Based on general profile theory, a scroll pitch ine is defined as of A-type with which the outer profile of orbiting croll and the inner profile of fixed scroll can be formed. Another croll pitch line is defined as B-type with which the inner profile of rbiting scroll and the outer profile of fixed scroll can be formed. he scroll compression chambers corresponding to the pitch lines f A-type and that of B-type are defined as compression chambers f A-series and B-series, respectively. Then the general geometrial model of scroll compression chamber is set up. It can demontrate the deforming, opening, and vanishing process of all scroll hambers. DOI: 10.1115/1.4000647\neywords: scroll fluid machine, general profile, scroll compresion chamber, general geometrical model\nIntroduction The working principle of scroll fluid machine has been known ince 1905 when it was invented by Leon Creux 1 , a French ngineer. For a scroll fluid machine, the meshing and geometry heories are its theoretical basis and, certainly, the most important tudy object. Zhenquan et al. 2,3 , and Bush and Beagle 4 perormed effective studies of meshing theory and profiles of the croll fluid machine, Bush and co-workers 4,5 put forward a eneral profile geometry theory, Gagene and Nieter 6 simulated he working processes of a scroll compressor with a general prole. In the working process of a scroll fluid machine, its basic erformance is determined by the geometrical characteristics of he scroll compression chamber. So it is very critical to establish a eneral geometrical model of compression chamber.\nCharacteristic Angle of Scroll In the working process of scroll fluid machine, all compression hambers formed by one profile of an orbiting scroll inner profile r outer one and one profile of a fixed scroll outer profile or nner one can be called a set of compression chambers.\nFor a scroll fluid machine with z scroll wraps, as shown in Fig. , there are z identical scroll wraps with a phase difference of /z on either orbiting or fixed scroll, respectively. In its working rocess, there is 2z sets of scroll compression chambers that are ormed with a phase difference of /z. Considering a phase diference of /z between orbiting and fixed scrolls, the characterstic angle, c, is defined for a scroll with z wraps as follows:\nContributed by the Fluids Engineering Division of ASME for publication in the OURNAL OF FLUIDS ENGINEERING. Manuscript received April 5, 2008; final manuscript eceived October 25, 2009; published online December 15, 2009. Assoc. Editor:\nhunill Hah.\nournal of Fluids Engineering Copyright \u00a9 20\nhttps://fluidsengineering.asmedigitalcollection.asme.org on 05/27/2019 Terms of\nc = z for z = 1,2,3, . . . 1\n3 General Geometrical Model of Scroll Compression Chamber\n3.1 Scroll Wrap With General Profile. As shown in Fig. 2, for a point on a continuous curve, assume that its normal angle is the positive angle between the x-axis and the normal to the curve at this point and has a property of superposition. Therefore, an arbitrary point on this curve can be expressed as a complex number by the normal component, Rn, and tangential component, Rt, of its radius-vector as follows:\nP = Rn exp j + Rt exp j +\n2\n2\nRt = dRn\nd\nAccording to the general profile theory, there are three types of general profile, they are\nRn = c0 + c1 + c2 for type-I of scroll profile\nRn = c0 + c1 cos + c2 for type-II of scroll profile 3\nRn = c0 + c1 + c2 2 + c3 3 for type-III of scroll profile\nAs shown in Fig. 3, the inner and outer profiles of scroll wraps are formed by shifting their pitch lines by a distance of Ror /2 in the normal direction of the pitch lines. The pitch line forming outer profile of orbiting scroll and inner profile of fixed scroll is defined as A-type pitch line and similarly, the pitch line forming inner profile of orbiting scroll and outer profile of fixed scroll is defined as the B-type one.\nFor a scroll fluid machine, A-type pitch line or B-type one may have z identical pitch lines in number. By using a subscript l l =a or b , the type of pitch line can be indicated, a subscript m m=in or out can be used to indicate the outer or inner profile of scroll wrap, a subscript u u=o or f can be used to indicate the orbiting or fixed scroll, and a subscript i i is positive integer and 1 i z is used to indicate the ordering number of pitch line. Taking the first pitch line i=1 of A-type l=a as a reference and assuming b,i,s a,i,s, then the relationship between the initial normal angles of A-type and B-type pitch lines is\nb,i,s = a,i,s + c 4 The relationship between the phases of two pitch lines of identical type is\nl,i,s = l,1,s + 2 c i \u2212 1 for i = 1,2, . . . ,z 5 The relationship between the normal components of radius-\nvector on two pitch lines of identical type can be expressed as\nRn,l,1 = Rn,l,i + 2 c i \u2212 1 for l,1,s l,1,e 6 The pitch lines and scroll profiles can be expressed as\nPl,i = Rn,l,i exp + 2 c i \u2212 1\n+ Rt,l,i exp j + 2 c i \u2212 1\nPl,i,m = Pl,i Ror 2 exp j + 2 c i \u2212 1 for l,1,s l,1,e\n7\nwhere the positive sign \u201c+\u201d is used for m=in and negative sign \u201c\u2212\u201d for m=out.\nEffective turns of every scroll pitch line is the same as\nnl,i = l,i,e \u2212 l,i,s = l,1,e \u2212 l,1,s = n 8\n2 2\nJANUARY 2010, Vol. 132 / 014501-110 by ASME\nUse: http://www.asme.org/about-asme/terms-of-use", "C f w t o b p d f s m t p\n0\nDownloaded From:\n3.2 General Geometrical Model of Scroll Compression hamber. As shown in Fig. 4, every compression chamber is ormed by moving the outer or inner profile of the orbiting scroll rap in the direction of the profile normal at the meshing point up o the inner or outer profile of the fixed scroll wrap by a distance f Ror. The normal angle region of one scroll compression chamer is l,i , l,i+2 , where l,i is the normal angle at the meshing oint. The compression chamber formed by A-type pitch lines is efined as A-series scroll compression chamber, and the chamber ormed by B-type pitch lines is defined as B-series scroll compresion chamber. A-series or B-series scroll compression chamber ay have z sets of compression chambers in number. Between wo scroll compression chambers of different series, there is a hase difference of c. On the other hand, between the scroll\n14501-2 / Vol. 132, JANUARY 2010\nhttps://fluidsengineering.asmedigitalcollection.asme.org on 05/27/2019 Terms of\ncompression chambers of identical series but different set, there will be a phase difference of 2 c.\nFor an arbitrary scroll compression chamber, the distance between its opposed profiles along the normal of its pitch line is defined as its width and denoted by Bl,i l,i , l,i , where l,i l,i , l,i+2 . The length of its pitch line is defined as its length and denoted by Lcl,i l,i . The product of its width and height of scroll wrap is defined as its normal sectional area and denoted by Anl,i l,i , l,i , that is\nAnl,i l,i, l,i = HBl,i l,i, l,i , l,i,s l,i l,i,e \u2212 2\n9 l,i l,i l,i + 2 , i = 1,2, . . . ,z\nwhere H is the height of scroll wrap. The normal sectional area and length of a scroll compression chamber can be used to demonstrate its size and shape, and they reveal the nature of the orbiting movement.\nFor arbitrary l,i l,i,s , l,i,s\u22122 and l,i l,i , l,i+2 , the length of scroll compression chamber of A-series and B-series can be calculated by\nLcl,i l,i = Ll,i,in l,i + Ll,i,out l,i\n2 for i = 1,2, . . . ,z 10\nIn Eq. 10 , Ll,i,in l,i and Ll,i,out l,i can be calculated as follows:\nll,i,m l,i = l,i\nl,i+2\nRn,l,i l,i d l,i Ror + Rt,l,i l,i + 2\n\u2212 Rt,l,i l,i 11\nwhere the positive sign + is used for m=in and negative sign \u2212 for m=out.\nWith the range of 0 l,i l,i , l,i /4, as shown in Fig. 5,\nTransactions of the ASME\nUse: http://www.asme.org/about-asme/terms-of-use", "t\nT b\nw i\nB\nT\nT\nT\nI n\ne p w\nJ\nDownloaded From:\nhere is l,i l,i , l,i =0 for l,i= l,i , l,i+ , . . . , l,i+k , otherwise l,i and l,i l,i , l,i satisfy Eq. 12\nsin l,i \u2212 l,i + l,i l,i, l,i tan l,i l,i, l,i + cos l,i \u2212 l,i + l,i l,i, l,i\n\u2212 sin l,i \u2212 l,i\nsin l,i l,i, l,i = 0 12\nhen the width of scroll chamber Bl,i l,i , l,i can be calculated y using the following formula:\nBl,i l,i, l,i = Rn,l,i l,i, l,i + Ror\n2 \u2212 Ror sin l,i \u2212 l,i + l,i l,i, l,i sin l,i l,i, l,i\nC2 , + D2 , \u2212 R2 13\nl,i l,i l,i l,i l,i l,i t,l,i l,i\nournal of Fluids Engineering\nhttps://fluidsengineering.asmedigitalcollection.asme.org on 05/27/2019 Terms of\nwhere Cl,i l,i , l,i =sin l,i\u2212 l,i /sin l,i l,i , l,i +Rn,l,i l,i\nl,i l,i , l,i , and Dl,i l,i , l,i =Rt,l,i l,i l,i l,i , l,i , the positive sign + is used for l,i l,i , l,i+ and otherwise the sign \u2212 is used. Obviously, Bl,i l,i , l,i =0 for l,i= l,i and l,i= l,i+2 , and Bl,i l,i , l,i =Ror for l,i= l,i+ .\nFor an arbitrary scroll compression chamber changing from the instance of its formation to that of its vanishing, its living period Tl,i is related to the effective turns of its pitch line in the following form:\nTl,i = l,i,e \u2212 l,i,s \u2212 2\n14\nwhere is the angular speed of orbiting motion. Thus, if the orbiting angle is an arbitrary l,i 0,2 , then the orbiting angle l,i and the number of living chambers N can be expressed as\nl,i l,i\nl,i = l,i \u2212 2 \u00b7 int l,i\n2\nNl,i l,i = int nl,i \u2212 1 + sign nl,i \u2212 int nl,i + int l,i,s l,i\nfor l,i,s l,i l,i,e \u2212 2 , i = 1,2, . . . z 15\nObviously, l,i changes with a period of 2 . If l,i= l,i,e\u22122 , then the suction process would be completed and a scroll chamber ould be formed. With a subsequent reduction in l,i, the volume of the scroll compression chamber would gradually reduce and reach ts minimum when l,i= l,i,e. With further orbiting movement, the scroll compression chamber would be opened. Taking the compression chamber composed of first scroll profile of A-type as a reference, the phase relationship between A-series and\n-series scroll compression chambers is\nb,i = a,i + c for i = 1,2, . . . ,z 16 he phase relationship between two scroll compression chambers of the same series is\nl,i = l,1 + 2 c i \u2212 1 for i = 1,2, . . . ,z 17 herefore, the relationship between all living scroll compression chambers and the first one of A-series can be expressed as\n= a,1 = a,1 \u2212 2 \u00b7 int a,1\n2\nNl,i = int nl,i \u2212 1 + sign nl,i \u2212 int nl,i + int a,1,s + 2 c i \u2212 1 + c c 2\na,1 + 2 c i \u2212 1 + c c 2 for a,1,s a,1 a,1,e \u2212 2 , i = 1,2, . . . ,z\n18 he relationship between normal sectional area and length of all scroll compression chambers can be expressed as\nAnl,i l,i, l,i = Ana,1 a,1 + a,1 + 2 c i \u2212 1 + c c\n2 , a,1 + 2 c i \u2212 1 +\nc c\n2\nLl,i l,i = La,1 a,1 + 2 c i \u2212 1 + c c\n2\nfor a,1,s a,1 a,1,e \u2212 2 , a,1 a,1 a,1 + 2 , i = 1,2, . . . ,z 19\nn Eqs. 18 and 19 , the positive sign + is used for l=b and egative sign \u2212 for l=a.\nSummarizing the above analysis, it is recognized that the genral geometrical model of scroll compression chamber is comosed of combined Eqs. 7 , 9 , 10 , 14 , and 15 , among hich Eq. 7 characterizes its profile and orbiting movement, Eqs. 9 and 10 characterize its shape and deformation, Eq. 14 characterizes its living period, and finally Eq. 15 characterizes the number of scroll compression chambers in the working process. The relationship between an arbitrary scroll compression chamber and the first one of A-series can be expressed by Eqs. 18 and 19 .\nJANUARY 2010, Vol. 132 / 014501-3\nUse: http://www.asme.org/about-asme/terms-of-use" ] }, { "image_filename": "designv11_20_0003711_pime_conf_1967_182_341_02-Figure29.3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003711_pime_conf_1967_182_341_02-Figure29.3-1.png", "caption": "Fig. 29.3. Principle drawings of waviness measurement of ring and ball", "texts": [ " In more sophisticated versions they are transformed to digital numbers which can be used in data processing. The pick-up signal can also be used for controlling a rejection device in a quality inspection unir. voi 192 rt K at UNIV CALIFORNIA SANTA BARBARA on September 4, 2015pcp.sagepub.comDownloaded from Frequently the pick-up signal is also fed into a loudspeaker to permit the detection of signal properties that are drowned in the wide band outputs; thus, for instance, small scratches and indentations are best observed by means of a loudspeaker. Fig. 29.3 illustrates the VKR and VKK equipment, and a cross-sectional view of the pick-up is shown in Fig. 29.4. Vibration measuring equipment, which is known as VKL, is frequently mentioned in this paper. It contains essentially the same mechanical and electronic parts as the VKR and VKK equipment. In particular the spindle, pick-up, and preamplifier are the same and, generally, the same type of filters are used. As previously mentioned, the VKL equipment contains a device for obtaining a purely axial load over the bearing, and the pick-up measures the radial velocity at a point at the outer perimeter of the non-rotating outer ring" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002358_jahs.53.282-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002358_jahs.53.282-Figure1-1.png", "caption": "Fig. 1. NASA Glenn spiral bevel gear/face gear test facility.", "texts": [ " Stresses are presented in terms of American Gear Manufacturing Association (AGMA) stress indices based on spur gear approximations, rather than those determined by finite element analysis. Although not as accurate, the AGMA stress indices can be computed rather quickly, which in turn, are extremely useful during initial sizing and trade studies in a preliminary design. Apparatus The experiments reported in this report were tested in the NASA Glenn spiral bevel gear/face gear test facility. An overview sketch of the facility is shown in Fig. 1(a), and a schematic of the power loop is shown in Fig. 1(b). The facility operates in a closed-loop arrangement. A spur pinion drives a face gear in the test (left) section. The face gear drives a set of helical gears, which in turn, drive a face gear and spur pinion in the slave (right) section. The pinions of the slave and test sections are connected by a cross shaft, thereby closing the loop. Torque is supplied in the loop by physically twisting and locking a torque in the preload coupling on the slave section shaft. Additional torque is applied through a thrust piston (supplied with high pressure nitrogen gas), which exerts an axial force on one of the helical gears" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003887_19346182.2012.663534-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003887_19346182.2012.663534-Figure6-1.png", "caption": "Figure 6. A ball can be represented by a mass m and two springs. If the racquet rotates then so does the ball since the top spring compresses more than the bottom spring and exerts a greater force on the ball. Part (a) shows the geometry just prior to the collision and (b) shows the geometry during the collision.", "texts": [ " The same effect would occur if the ball was glued to the strings since the ball and the racquet would both rotate 108. The ball is not glued to the strings but it is squashed against the strings. The top end of the racquet is rotating faster than the bottom end, so the top side of the ball is pushed harder towards the net than the bottom side. As a result, the ball rotates with topspin, even if the racquet head is not rising when it strikes the ball. Ten degrees of forward rotation in 0.004 s corresponds to about 400 rpm of topspin, which will contributesignificantly to the amount of topspin in a kick serve. Figure 6 shows a simple model of the rotating racquet effect. The ball can be represented by a mass m plus two springs of spring constant k separated by a D ow nl oa de d by [ U ni ve rs ity o f D el aw ar e] a t 1 7: 40 2 7 Ju ly 2 01 3 distance D. More generally, k represents the combined stiffness of the ball and the strings. If the racquet translates in a straight line towards the ball, both springs compress equally and the ball accelerates in a straight line without rotating. If the racquet is rotating then the top spring compresses more than the bottom spring so the force F2 is greater than F1 and the ball will rotate in the same direction as the racquet. The situation shown in Figure 6 can be solved numerically. We assume that the racquet rotates about a fixed axis in the handle located at distance R from the bottom spring. The angular velocity of the racquet, just prior to the collision, is v0. During the collision, suppose that the left end of the bottom spring is displaced by a horizontal distance x1 and the right end is displaced by a distance x2. The horizontal force F1 on the ball is then given by F1 \u00bc k\u00f0x1 2 x2\u00de. If the left end of the top spring is displaced by x3 and the right end by x4 then F2 \u00bc k\u00f0x3 2 x4\u00de", " The analysis given in the previous section was simplified by ignoring the fact that the normal reaction force on the ball rotates in direction as the racquet rotates. Such an approximation can be justified on the basis that a small direction change in the force on the ball will have a negligible effect on the resulting torque on the ball. A more significant result is that the change in direction of the normal force will affect the outgoing angle of the ball. The ball will not emerge in a direction parallel to the initial velocity of the racquet head, as suggested by the result shown in Figure 6(b), unless the rotation angle of the racquet during the impact is very small. The racquet head is rotating rapidly towards the net when the ball is struck, at an average angular velocity of about 40 rad s21 during the impact in a typical kick serve. The racquet head therefore rotates through an angle of about 108 during the impact, typically about 4 ms in duration. If the ball is struck when the head is vertical, then the ball will come off the strings when the head is tilted forward by about 108" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001942_1.2982025-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001942_1.2982025-Figure2-1.png", "caption": "Figure 2. Exploded view of a single-cell test prototype fuel cell used in this study.", "texts": [ " Ammonium-type anion-exchange membranes from Tokuyama and Asahi were in the Cl- form. The membranes are composed of tetra-alkyl ammonium as the fixed cation groups. The Cl- form membrane was converted to the hydroxyl form (OH-) by conditioning the membranes in 1 M KOH for 24 hours. Experimental evidence of substantially higher efficiencies using an AAEM fuel cell was demonstrated in our laboratory. Assembly of Fuel Cells The fuel cells were assembled using single-cell test prototypes. An exploded representation of a cell is shown as Figure 2. The catalyst loadings were identical for PEM and AAEM fuel cells (0.5 mg/cm2). The membrane electrode assemblies used carbon-supported catalysts coated on carbon paper. The anodes were Pt-Ru/C and the cathodes were Pt/C. The electrodes were fabricated in the same manner so that a comparison of methanol and glycerol in a PEM fuel cell could be compared to glycerol and crude glycerin in an AAEM fuel cell. Fuel Cell Tests Polarization and power density curves were obtained by placing a known resistive load across the cell and measuring the resultant cell voltage" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.12-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.12-1.png", "caption": "Fig. 14.12 General construction of a chain track", "texts": [ " Another feature of the chain chassis is its good stability, which is of considerable importance for hydraulic excavators. Rubber belt tracks combine the advantages of both systems. They are applied in mini-excavators and road pavers. In the area of agriculture, rubber belt tracks are used in tractors equipped with high engine powers designed to work on the soil. Rubber belt tracks suffer from less vibrations compared with steel chains, while almost the same driving speeds are possible as with standard tractors. Moreover, they cause less damage to the road surface [14.12]. Figure 14.12 shows the general construction of a chain track. It consists of the chain, the base plate, the track rollers, the support rolls, the driving wheel, the guide wheel, the tensioning device, and the track frame. A chain chassis consists of at least two chain tracks. The guide wheel usually has only one guide profile and, using the chain tensioning device, can be shifted in order to tighten the chain. The smaller track rollers on the bottom serve to support the machine against the chain. The track rollers are designed to transmit even lateral forces emerging from steering movements or slope forces" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003573_s12239-011-0008-x-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003573_s12239-011-0008-x-Figure4-1.png", "caption": "Figure 4. Displacement indicator measuring point and indicator anchorage point.", "texts": [ " pattern was measured by applying an appropriate volume of a gear-marking compound to the gears before the two sets of gear teeth were engaged with each other to turn and by copying the imprint on the gears with tape after turning. The tooth contact pattern was then checked. Figure 2 shows the measured tooth contact pattern. 2.2. Positions of the Deflection Measurement and the Axle Noise Section of Input Torque Conditions on the Vehicle The Hypoid gear rotation direction and torque direction are shown in Figure 3. The measurement position in the direction test is shown in Figure 4, and the axes (E, P, G, and \u03b1) of the hypoid gears are shown in Figure13. Except for the upper/lower and right/left torsions of the pinion and ring gears and the displacement, the real deflections were measured with a trigonometric function (Coleman, 1975). This test result also included tooth deflection. Moreover, each displacement sensor was fixed and measured based on the inner bearing assembly of the pinion gear. Because the input revolution in the direction test was 5~10 rpm and because oil was not used, the test was performed with the low revolution to prevent the teeth from being damaged by the heat generated in the surface of the gear under the torque condition" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003573_s12239-011-0008-x-Figure13-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003573_s12239-011-0008-x-Figure13-1.png", "caption": "Figure 13. Axial coordinate of the hypoid gear.", "texts": [ " pattern was measured by applying an appropriate volume of a gear-marking compound to the gears before the two sets of gear teeth were engaged with each other to turn and by copying the imprint on the gears with tape after turning. The tooth contact pattern was then checked. Figure 2 shows the measured tooth contact pattern. 2.2. Positions of the Deflection Measurement and the Axle Noise Section of Input Torque Conditions on the Vehicle The Hypoid gear rotation direction and torque direction are shown in Figure 3. The measurement position in the direction test is shown in Figure 4, and the axes (E, P, G, and \u03b1) of the hypoid gears are shown in Figure13. Except for the upper/lower and right/left torsions of the pinion and ring gears and the displacement, the real deflections were measured with a trigonometric function (Coleman, 1975). This test result also included tooth deflection. Moreover, each displacement sensor was fixed and measured based on the inner bearing assembly of the pinion gear. Because the input revolution in the direction test was 5~10 rpm and because oil was not used, the test was performed with the low revolution to prevent the teeth from being damaged by the heat generated in the surface of the gear under the torque condition", " Figure 11 shows the pro- cesses of configuring and computing the triaxial transmission error, and Figure 12 is a diagram of the actual triaxial transmission error measurement experiments. Three ring encoders (ERM280) of HEIDENHAIN with three Rotec EDR counter boards were used on the input, LH output, and RH output of the axle. Each encoder rotation signal was calculated with RAS software by Rotec. Because a revolutionary angle meter has an angle revolution power of 20480 per revolution, it was possible to measure the transmission error with a high precision: 200 or more times the number of gear teeth (30~80) to be measured. 4.1. Deflection under Input Torque of the Automobile Figure 13 shows the coordinate system of the hypoid gear. Figure 14 shows the results of the deflection test under the axle input torque of the E (offset direction), P (pinion direction), and G (gear direction) axis, and Figure 15 shows the \u03b1 angle. Table 2 shows the result of the gear tooth contact test under the input torque. The variation of the tooth contact shows spreads out toward the heel as the torque increases in the typical Toe Bias In pattern. According to the input torque, the P axis tends to show a + direction on the drive side of the gear contact tooth when accelerating a vehicle, and a \u2212 direction on the coast side of the contact gear when decelerating a vehicle" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001925_0954406jmes476-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001925_0954406jmes476-Figure4-1.png", "caption": "Fig. 4 The model for temperature computing in the ball bearing", "texts": [ " Hence, the raceways are actually heated by moving heat sources, and the temperature in the rings can be obtained using a step-by-step computation of the heating and cooling on the contact ellipses. To save CPU time, the model must be simplified. Transient temperature in the ball bearing is assumed axisymmetrical because the rotational speed is very high. Then the model can be simplified from three dimensions to two dimensions and a lot of CPU time could be saved. The model with the shaft, ball, and bearing chock is shown in Fig. 4. The black thick lines in the raceways are the contact lines between the balls and the raceways. The heat flux and operating time on the contact lines are presented in equation (16). The value of dt can be determined from the width of the contact ellipse, the shaft speed, and the revolution speed of the ball; ti+1 \u2212 ti can be determined from the ring speed, the revolution speed, and the number of the ball. The heat generated is distributed on the contact surfaces according to a coefficient k that can be determined through the heat conduction coefficient of the material, the convective heat transfer coefficient, and the moving velocities of the contact surfaces. During contact, ti < t < (ti + dt), q could be a large number. Conversely, during no contact ti + dt < t < ti+1, q could be zero q = \u23a7\u23aa\u23a8 \u23aa\u23a9 kH A ti < t < ti + dt 0 ti + dt < t < ti+1 (16) Figure 4 also shows convection boundary conditions on the exposed surfaces of the unit components. It is the forced convection of the lubricant in the compartment of the bearing. The heat transfer coefficients can be defined by equation (17) [14]. Many factors affect the heat transfer coefficient in the bearing compartment. Among these, the primary factors are the velocity and quantity of lubricant in the bearing. Due to the centrifugal force, the quantity of the lubricant in the outer raceway is more than that in the inner JMES476 \u00a9 IMechE 2007 Proc" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003526_iccrd.2011.5763939-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003526_iccrd.2011.5763939-Figure1-1.png", "caption": "Figure 1. Intercept Scenario for UAV and Target", "texts": [ " The construction of this paper is as follows, in the second section, an integrated guidance/autopilot model is formulated, in the third section, the control-guidance law is deduced and by using an auxiliary control based on a sliding mode estimator, the control will be more accurate and the robust performance will be enhanced. A simulation is made and the simulation results are presented in the fourth section, finally, a conclusion is made. II. THE MODEL OF INTEGRATED GUIDANCE/AUTOPILOT FOR UAV ON THE SEA When a UAV flies on the sea at a height of no more six meters, the duty of hitting the target can be changed into a problem of a standard two-dimensional scenario shown In Fig 1. where R -the distance between UAV and target; u T,\u03be \u03be - flight-path angle of UAV and target; U T,V V - velocity of UAV and target; Q -LOS. The kinematical equations between UAV and target can be shown as (1)-(4) T T U Ucos cosR V V\u03bc \u03bc= \u2212 (1) T T U Ucos cosR V V\u03bc \u03bc= \u2212 (2) T TQ \u03be \u03bc= + (3) ___________________________________ 978-1-61284-840-2/11/$26.00 \u00a92011 IEEE U UQ \u03be \u03bc= + (4) Equations (1) and (2) can be changed into the following form of (5-6) T T U Ucos( ) cos( )R V Q V Q\u03be \u03be= \u2212 \u2212 \u2212 (5) T T U Usin( ) sin( )RQ V Q V Q\u03be \u03be= \u2212 \u2212 \u2212 (6) derivate (6), and substitute (5) into it, after some mathematical manipulations, it can obtain that U U T T T T T U U U 2 sin( ) sin( ) cos( ) cos( ) RQ RQ V Q V Q V Q V Q \u03be \u03be \u03be \u03be \u03be \u03be + = \u2212 \u2212 + \u2212 + \u2212 \u2212 \u2212 (7) because the overloads of UAV and target can be written as T T Tn V \u03be= \u2212 , U U Un V \u03be= \u2212 , so(7)can be written as U U T T T T U U 2 sin( ) sin( ) cos( ) cos( ) RQ RQ V Q V Q n Q n Q \u03be \u03be \u03be \u03be + = \u2212 \u2212 + \u2212 \u2212 \u2212 + \u2212 (8) Let RQ U= , then U U T T T T U U U U T T T T U U U T U T U T sin( ) sin( ) cos( ) cos( ) sin( ) sin( ) cos( ) cos( ) ( , , , , , , ) U V Q V Q n Q n Q RQ RV Q V Q n Q n Q U R Rf V V n n Q U R \u03be \u03be \u03be \u03be \u03be \u03be \u03be \u03be \u03be \u03be =\u2212 \u2212 + \u2212 \u2212 \u2212 + \u2212 \u2212 =\u2212 \u2212 + \u2212 \u2212 \u2212 + \u2212 \u2212 = \u2212 (9) In addition , the mathematical model of UAV shown in (6-8) is shown as", "3 ) y y y y y y y y y y y M y y y yd y y y y y y y y M y y y yd y qSLV k m J V qsL qsLm m J J qSL qsLk m m J V J qsL m J \u03c9 \u03b4\u03b2 \u03c9 \u03b2 \u03b4 \u03b2\u03b2 \u03c9 \u03c9 \u03b2 \u03c9 \u03b2 \u03c9 \u03c9 \u03b2 \u03b4 \u03c9 \u03b2\u03c9 \u03b2 \u03c9 \u03c9 \u03b2 \u03b4 \u03c9 = + = \u2212 + + + \u2212 = \u2212 + + + \u2212 (29) Let 2 4 ( 57.3 57.3 ) y y y y y y y Uy y yd y y J qSL m J VqsLm qsL m k J \u03c9 \u03b4 \u03b2 \u03b4 \u03b2 \u03c9 \u03b2 \u03c9 \u03c9 = \u2212 + + \u2212 + (30) Then it can obtain that 2 2 4 3 2yV k k bV\u03c9 \u03b2= \u2212 \u2212 \u2264 (31) where 3 4min( , )b k k= , from (31) we can see that the exponential stability can be got. In fact, the value of TV , Tn and T\u03c8 in (13) can not be measured directly, so it is necessary to estimate them, and the estimating method will be given as follow. Refer to Figure 1, there are the following relations T Ux x x= \u2212 (32) T Uz z z= \u2212 (33) 2 2R x z= + (34) ( ) /R xx zz R= + (35) arctan zQ x = \u2212 (36) 2 zx zxQ R \u2212= (37) After some mathematical manipulations, it can obtain that T Ucos sinx R Q RQ Q x= \u2212 + (38) T Usin cosz R Q RQ Q z= \u2212 \u2212 + (39) then the estimation of the velocity of the target can be expressed as 2 2 1/ 2 T T T \u02c6 ( )V x z= + (40) and we can get the estimation of T\u03c8 T T T \u02c6 arctan z x \u03be = \u2212 (41) By using a low pass filter, T\u0302\u03be and T\u0302V can be got. Then it can obtain that T T T \u02c6 \u02c6\u02c6 Vn g \u03be= \u2212 (42) so that (22) can be rewritten as Mc1 U U M 2 1 T T T T 1 sin( ) cos( ) \u02c6 \u02c6\u02c6 \u02c6sin( ) cos( ) 2 ( ) Rn A V Q Q R V Q n Q A sign A\u03c7\u03c7 \u03be \u03be \u03be \u03be \u03c9 \u2212\u2212 =\u2212 \u2212 \u2212 \u2212 + \u2212 \u2212 \u2212 \u2212 + (43) In practice, the estimation error of the motion information of the target can not be neglected, now an auxiliary control based on a sliding mode estimator is used" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003132_jmes_jour_1969_011_008_02-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003132_jmes_jour_1969_011_008_02-Figure7-1.png", "caption": "Fig. 7", "texts": [ " Although the bearings have radial symmetry, unstable bands are in evidence not only at the shaft natural frequencies, but also at half sums of these. The form of the half-sum instabilities is indicated in Fig. 5. Single-mass shaft on massive and jlexible foundations The idealized system, Fig. 6, is considered; the shaft principal stiffnesses on rigid bearings are A( 1 +a) and With /3 = 10, k, = 100 and k, = 500 the natural A( 1 - u). frequencies of the generating system are found to be wl,, = 0.9552, wcl = 0.9952, wy2 = 3.252 wZ2 = 7.0952 where The corresponding modal shapes, shown in Fig. 7, yield the inertia coefficients 1 (32) ayl = 1-098m, a;, = 1.005m, ayz = 846m a,, = 24OOOm The coefficients p, v are now determined from equations (5), including the 'Stieltje' terms due to shear at the shaft bearings, 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 K C E 5 Vol I1 No 1 I969 at DEAKIN UNIV LIBRARY on August 12, 2015jms.sagepub.comDownloaded from 64 H. F. BLACK similar expressions applying to the other coefficient types. The complete coefficient array is then given by 1 0", "',,] = - [ If the close juxtaposition of wYl and w~~ is left aside, 10 bands of parametric instability are present according to equations (6) and (7). 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 bands may coalesce, the asymptotic solutions, equations (13), (14) and (15), being applicable. The stability map is shown in Fig. 8, where the form of the mode coupled instabilities is also indicated. It is seen that the instability bands accociated with the higher modes are negligibly thin. This is to be expected as these modes (+uz and +2z in Fig. 7) consist principally of base vibrations and therefore the asymmetry of the shaft has little effect. UniformIy asymmetric shaft, with asymmetric 'perfect' boundary conditions The shaft of Fig. 9 has a generating system consisting of pin-pin shaft modes in the Oxy plane, and free-pin shaft at DEAKIN UNIV LIBRARY on August 12, 2015jms.sagepub.comDownloaded from PARAMETRICALLY EXCITED LATERAL VIBRATIONS OF AN ASYMMETRIC SLENDER SHAFT 65 modes in the 0x2 plane. In evaluating the coefficients v, p the notation and numerical data from (12) can be utilized directly" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002858_s12239-009-0081-6-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002858_s12239-009-0081-6-Figure1-1.png", "caption": "Figure 1. Full vehicle FEA model is illustrated in the IDEAS software environment. Some plates are omitted from the figure to display the interior components of the vehicle.", "texts": [ " Acceleration of the mid-point of the driver\u2019s seat rail was used as a measure of the transmitted vibration. Rubber engine mount stiffness, supporting plate (sub-frame member) thick- ness, and the stiffness of the connection of the supporting member to the body were considered as optimization parameters. Various shell elements, rigid links, lumped mass elements, and spring elements were used in the present full vehicle finite element model. All the subsystems including suspension components, steering system, cooling system, power train, exhaust system, etc., were modeled (Figure 1). All spot weld joints are modeled as well. Hypermesh software was used to create and optimize the FEA model, and NASTRAN software was used to analyze the model. To ensure a C1 continuity, Hermite-type elements were preferred (Reddy, 2005). Since such elements are not defined in well-known engineering software such as NASTRAN, ANSYS, I-DEAS, ADINA, and ABAQUS, second-order quadratic shell elements were used for modeling the plates of the body. It is known that for identical number of nodes, rectangular elements provide more accurate results (in comparison with the triangular elements)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001664_978-3-540-77608-6_20-Figure20.19-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001664_978-3-540-77608-6_20-Figure20.19-1.png", "caption": "Fig. 20.19. Single-level attachment system with oriented cylindrical cantilever beams with spherical tip. In this figure, l is the length of fibers; \u03b8 is the fiber orientation, R is the fiber radius; Rt is the tip radius; S is the spacing between fibers; and h is distance between base of model and mean line of the rough profile [57]", "texts": [ " The design variables for an attachment system are as follows: fiber geometry (radius and aspect ratio of fibers, tip radius), fiber material, fiber density and fiber orientation. The optimal values for the design variables to achieve the desired properties should be selected for fabrication of a biomimetic attachment system. The fiber model of Kim and Bhushan [57] consists of a simple idealized fibrillar structure consisting of a single-level array of micro/nano beams protruding from a backing as shown in Fig. 20.19. The fibers are modeled as oriented cylindrical cantilever beams with spherical tips. In Fig. 20.18, l is the length of fibers; \u03b8 is the fiber orientation; R is the fiber radius; Rt is the tip radius; S is the spacing between fibers; and h is distance between upper spring base of each model and mean line of the rough profile. The end terminal of the fibers is assumed to be a spherical tip with a constant radius and a constant adhesion force. Kim and Bhushan [57] modeled an individual fiber as a beam oriented at an angle \u03b8 to the substrate and the contact load F is aligned normal to the substrate" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002565_j.tws.2008.08.010-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002565_j.tws.2008.08.010-Figure1-1.png", "caption": "Fig. 1. View of the clutch. The clutch disc is crushed between the mechanism pressure plate and the flywheel to transmit the torque from the combustion engine to the transmission.", "texts": [ " Car manufacturers requirements for greater power transmission, lightweight, low cost design, smaller design space, high comfort and high effectiveness lead to develop new clutches mechanisms such as ultrasonic clutch [9] or piezoelectric clutch [10]. Nevertheless, today\u2019s passenger cars and light trucks are now almost exclusively equipped with conventional friction clutches (i.e. diaphragm spring clutches). In this work, we only consider dry friction clutches. It is composed ll rights reserved. s Group, Product Advanced , France. . of the clutch disc, the flywheel, and the mechanism, which is itself composed of a cover, a diaphragm spring, and a pressure plate (Fig. 1). When the engine rotates freely (no gear engaged) no torque is transmitted because the clutch disc is not in contact with the mechanism and rotates freely (Fig. 1). When a gear is engaged, the clutch disc is compressed by the diaphragm spring between the pressure plate of the mechanism and the flywheel transmitting the rotary motion and the torque to the mechanism and thus to the wheels. In the field of clutch systems a lot of research has been done for a better understanding of mechanical phenomena. During the clutch engagement manoeuvre, sliding contact occurs between the pair of clutch facings mounted on the clutch disc and the counter faces belonging to the flywheel and the pressure plate" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001370_s11044-007-9038-6-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001370_s11044-007-9038-6-Figure9-1.png", "caption": "Fig. 9 Multilink suspension model", "texts": [ " Figure 8 shows the solution differences between GCP and AFM with 1, 2, and 4 ms integration step-sizes. They are almost identical and reasonably small. Thus, the solutions corresponding up to the 4 ms integration step-size can be acceptable for simulations such as HILS. Springer In order to see whether the proposed AFM is applicable for HILS, the full vehicle simulation has been carried out. The full vehicle model consists of McPherson Strut front suspensions and Multilink rear suspensions as described in Figure 1. Figure 9 shows the Multilink suspension model. It consists of a LCA (lower control arm), a UCA (upper control arm), and a suspension swing arm. The LCA is connected to the chassis frame with a universal joint. The UCA is also linked to the chassis frame with a universal joint. The suspension swing arm is connected to the chassis frame, the LCA, and the UCA with spherical joints. Spring and damper elements are used. Bump and rebound stoppers are also modeled as nonlinear springs. A magical formula is used for the tire model" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000286_00021369.1984.10866431-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000286_00021369.1984.10866431-Figure7-1.png", "caption": "FIG. 7. pH Dependence of Is and v.", "texts": [ " PARAMETERS OF THE ENZYMATIC REACTION OF GOD WITH BENZOQUINONE AS AN ELECTRON ACCEPTOR (pH 5.0, 25\u00b0C) K1 K2 I s max 10mM 10- 1 mM /lA Immobilized 6.2 1.1 5.3 (on graphite electrode) (ko= 2.3 x 102 / S) Solubilized 7.3 4.2 (in solution) vmax 1.8 (ko = 3.6 x 102 / s) 1.0 '\\0 ::: 0.5 ~\\.t> p It) N It) :;- N ~ > c 0 iy ~ E - 0.5 0 \u2022 3.2 FIG. 6. Arrhenius Plots of Is and v. The pH dependence of Is and V was exam ined in buffer solutions of various pHs, all containing 20 mM n-glucose and 0.5 mM BQ. The results are shown in Fig. 7, where the magnitudes of Is and V are expressed as the relative values to the values at pH 5.0. Is and V showed a similar pH dependence, with the same optimum pH of 6.5, though at pHs higher than 5.5, the relative value of immobi lized GOD was high compared with that of solubilized GOD. All results given above indicate that the kinetics of the immobilized GOD are essen tially the same as those of the solubilized GOD, at least at pHs below 5.0. as given in Table II. The agreement of the two ko values is good as is that between the two K t values" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003397_s12289-010-0686-3-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003397_s12289-010-0686-3-Figure1-1.png", "caption": "Fig. 1 Schematic illustration of the roller expansion process from reference [12]; showing a the initial clearance between tube and tubesheet, b the mandrel in the center driving the expander rollers when the tube contacts the tubesheet wall, and c the interfacial pressure developed after removing the roller expander", "texts": [ "sa In this paper an axisymmetric FE model will be used to investigate the combined effects of the friction between tube and tubesheet, initial clearance, and strain hardening on the rolled joint strength represented by the interfacial pressure. The single hole equivalent sleeve model is used to simulate the tube\u2013tubesheet configuration. Roller expansion is a method for the expansion of heat exchanger tubes into tubesheet bores using a three-roll expander tool during a heat exchanger fabrication or re-tubing process. Heat exchanger re-tubing process uses torque controlled tube rolling to expand tubes into tubesheet holes to a specified tube wall reduction based upon empirical qualification results. Figure 1 illustrates roller expansion of a tube in a tubesheet hole to form a joint using Airetool\\ practice, which is the same procedure employed in most heat exchanger maintenance shops [11]. Industrial practice indicates that satisfactory joints are obtainable with steel tubes in heat exchangers by expanding the tubes to wall reductions of 5% to 10% after metal to metal contact of the tube with the tubesheet hole [11]. Light, normal, and heavy roll expansion corresponds to a reduction in the initial tube thickness of 2%, 5%, and 8%" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001149_s002211207800261x-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001149_s002211207800261x-Figure4-1.png", "caption": "FIGURE 4. Taylor instability mechanism: the Coriclis coupling. Owing to rotation, a radial velocity fluctuation u induces a tangential Coriolis force, the sign of which depends on that of the local vorticity 2A. Finally, the coupling is stabilizing for A > 0. Indeed, a fluctuation v > 0 induces a force 3\u2019: > 0 which in turn induces a velocity fluctuation u > 0 ( b ) ; then the tangential force 3\u2019; tends to damp the initial fluctuation v. Conversely, if A < 0, this is destabilizing.", "texts": [ " The simplest description which contains the essentials of the Taylor instability is obtained within the framework of a one-dimensional model which neglects the radial dependence of the fluctuations (a thorough account may be found in Chandrasekhar 1961). The hydrodynamic equations then reduce to Let us assume a tangential velocity fluctuation v. It induces a radial Coriolis force F,\" = 2 p ~ , v . Couette flow in nematic liquid crystals 281 This force tends to create a velocity component u. given by au at p - = 2pw0v. Now, with u is associated a tangential Coriolis force (figure 4a, b ) which tends to modify the fluctuation v assumed at the beginning. As explained in the caption of figure 4, the mechanism sketched above is locally stabilizing as long as Aw, > 0. (4.3) When Am, < 0 the mechanism is potentially destabilizing but the instability can take place only when it is strong enough to overcome the stabilizing effect of the diffusion of velocity fluctuations (viscous damping). As to the inviscid fluid, one can see that the fluid is stable when the two cylinders rotate in the same direction (so that w, does not change its sign) - say the positive one - and when which corresponds to the Rayleigh criterion applied to the unperturbed velocity profile (3" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001028_bfb0042540-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001028_bfb0042540-Figure3-1.png", "caption": "Figure 3: Hayati coordinates, and intermediate x-axis xxj.", "texts": [ " Locate a reference (base) coordinate frame coincident with the last joint of the thumb; then number the joints proceeding from that distal joint to the pMm, and then back out to the tip of the finger, as in Figure 1. Let the 4 x 4 homogeneous transformation I A~ = Rot(z,O j)Trans(z,sj)Trans(z,aj)Rot(z,a~) (1) where the notation Rob(z, \u00a2) indicates a rotation about an axis z by ~b and Trans(z, a) indicates a translation along an axis z by a. Since the D-H 8 parameter is not uniquely defined for consecutive parallel axes the following Hayati convention[7] (see Figure 3) is employed for such axes (i.e. for joints 0, 1, 3, 5 and 6): Aj = Ro Cz,0'j)Tra..C ,.j) o C=,aj)Ro Cy, (2) The position of the last llnk is computed by a sequence of these homogeneous transformations: T\u00a2. - - A I A ~ . . . A , , ! (3) The goal of kinematic calibration is to identify the geometric parameters sj, ~j and a~, and also any non-geometric parameters that may be included in the kinematic model. The non-geometric effects on the kinematic model potentially include bearing play and joint angle sensor error", " In total, the endpoint translation due to all of the parameter variations is given by: tt z i b~A0j + z~ aAs t + x$ x b$+lAa t + j=l and angular variation given by: r t . . E Z;-l~k0J \"~- K;Ao~j (15) $=1 Comparing these to (11) it is seen that the columns of each of the four Jacobians are and zj_l. x bj co15 - - , colt = = . j+ l (17) co15 , cols = The Jacobian columns for parameters of the alternate Hayati convention are found analogously to be: and Z I i ] a~ [ xxj_~ ] (18) , colj ~ = 0 [ , , ] xx~ x bj+~ (19) colj ~ ---- xx~ __ y~ \u2022 , COlj ~ = y j X hi+ 1 where xxj stands for the local x-axis just prior to the last rotation about the f lh y-axls by % (see Figure 3). In both parameter conventions the columns of the Jacobian with respect to the joint scale factors are: colj ~ --- Oj\" z)_1 2 . 7 D a t a c o l l e c t i o n a n d p a r a m e t e r e s t i m a t i o n AX__ = OA~+~ (21) where [clio I \u00b0'1] o= _ , A X = . A:\u00a2 n An estimate of the parameter errors is provided by minimizing (22) which yields A~ = (CTC)-~CTaX__ (24) Finally, the guess at the parameters is updated as and the iteration continues until A X ~ 0. In practice, during iteration the matrix uTu may become singular at an intermediate parameter set, even though the final parameter set does not have a singularity" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000309_s10569-004-1508-z-Figure12-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000309_s10569-004-1508-z-Figure12-1.png", "caption": "Figure 12. 3D L1 halo to L2 halo transfer.", "texts": [ " The transfers between two given halo orbits principally does not differ from the transfer between two given positions (see Section 5.1). Below the halo orbit will be specified by parameters Dx0; z0; v0;u0 giving the initial halo state vector fxL Dx0; 0; z0; 0; v0 cos u0; v0 sin u0g \u00f020\u00de where xL is given by (2). The position of the first halo exit and second halo entry (i.e. the terminal transfer positions) will be given by the time of flight t along the halo starting at the position (20). The L1 halo to L2 halo transfer of the LL-6-1 type is shown in three projections in Figure 12. The transfer time is 220 days. The halo orbit parameters are the following: Dx0 \u00bc 0:1 106 km, z0 \u00bc 0:1 106 km, u0 \u00bc 140 for the L1 halo and Dx0 \u00bc 0:2 106 km, z0 \u00bc 0:1 106 km, u0 \u00bc 30 for the L2 halo. The velocity v0 is selected to provide periodicity of the halo and is equal to 155.1m/s for the L1 halo and 254.3m/s for the L2 halo. The positions of the L1 halo exit and L2 halo entry are given by the times t \u00bc 80 day and t \u00bc 100 day respectively. Three projections of the 70-day transfer between two L1 halo orbits are shown in Figure 13" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000980_1-4020-3559-4-Figure19-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000980_1-4020-3559-4-Figure19-1.png", "caption": "Figure 19. Contact configuration during curve negotiation.", "texts": [ " Another aspect to note is that flange contact is detected with all creep force models. Even when running at the speed of 10 m/s, where the centrifugal forces effect is balanced by the track cant, flange contact occurs. Lateral flange forces develop on the wheels of both wheelsets of the front bogie as presented in Fig. 18 for a vehicle forward velocity of 10 m/s and using the Polach creep force model. During curve negotiation, the outer wheel of the leading wheelset and the inner wheel of the rear wheelset have permanent flange contact. Referring to Fig. 19, for the velocity of 10 m/s, the flange contact occurs on the outer and in the inner wheels of the vehicle. For the velocity of 20 m/s, only the outer wheels have flange contact. This is explained by the fact that, when running at 20 m/s, the vehicle negotiates the curve with a velocity higher than the balanced speed. Many applications of multibody dynamics require the description of the flexibility of its components. For structural crashworthiness it is La te ra l F la ng e F or c e [ N ] often unfeasible to use large nonlinear finite element models" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003906_978-3-319-00636-9-Figure2.10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003906_978-3-319-00636-9-Figure2.10-1.png", "caption": "Fig. 2.10 Dependency of the energy state splitting on the slipped cofacial molecular arrangement. This figure illustrates the dependence of the excited state splitting DE, expressed by the orientation factor j, b, on the change in the slipped cofacial arrangement, expressed by the angle H a. The arrows represent the orientation of the transition dipole moments of two molecules in a dimer with r as the center-to-center distance between them. While the molecules slip cofacially from the face-to-face orientation into the heat-to-tail one, the blue-shift goes over into a red-shift at H = 54.7 , the point at which the excitation energy state is not split. [33]", "texts": [ " In intermediate positions, by cofacially slipping of the molecules from one extrema to the other, E\u2019 continuously goes over to E\u2019\u2019 and vice versa, through an orientation with no energy splitting. DE can be calculated by the following formula [33]: D E \u00bc 2jM1M2jk r3 with k2 \u00bc \u00f01 3 cos2 h\u00de2 where M1 and M2 are the transition dipole moments of both molecules, which are the same in the case of identical molecules, r is the center-to-center distance between both molecules, j the orientation factor and h is the angle between both molecules (transition dipole moments) like shown in Fig. 2.10. The curve in Fig. 2.10b shows the orientation factor j over the angle h, which indicates the arrangement of both molecules in the dimer, starting from the face- 18 2 Theory and Literature Survey: Biomimetic Light-Harvesting to-face arrangement at h = 90 and ending at the head-to-tail arrangement at h = 0 . With slipping the molecules from the face-to-face arrangement into the head-to-tail one, the blue-shift continuously decreases and turns at h = 54.7 into the red-shift, which increases steadily to reach the maximal red-shift at the pure head-to-tail arrangement" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003891_978-94-007-1643-8_22-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003891_978-94-007-1643-8_22-Figure3-1.png", "caption": "Fig. 3 The handles trajectory in cartesian space. This kind of curves look like a third order polynom in the (Pi,ePiPf ,e \u22a5 PiPf ) frame. The input of these curves are the initial and final handles positions (Pi and Pf ) and the curvatures of both pieces of the curve (dev1 and dev2).", "texts": [ " For each patient, several sit-to-stand transfer trajectories were recorded, some examples of these trajectories are given in Fig. 2. The analysis of these transfer trajectories shows that the global shape of the trajectory is a \u201ds-like\u201d curve and is not directly related to the age or height of the patient but seems to be correlated with its own personal strategy to stand up or sit-down. This seems to reflect invariants of the trajectory generation [14]. The trajectory of the handles has to be similar to the general curve presented in Fig. 3. The term \u201dtrajectory\u201d refers here to Cartesian-space planning of the handle movement. A \u201dnatural\u201d trajectory is requested for comfortable human movement assisted by robotic devices. From our point of view, \u201dnatural\u201d means that the trajectory path must be compatible with hand movements when the sit-to-stand transfer is assisted by someone else. It must also be smooth and generate a continuous motion of the hand. As proposed in [15], smoothness can be quantified as a function of jerk, which is the time derivative of the acceleration", " Let s(t) represents the distance that the handles have moved along the curve at instant t. The curvilinear abscissa s(t) defined by Eq. 1 (see [15]), guarantees the smoothness of the handles trajectory. s(t) = s(Ti)+ (s(Tf )\u2212 s(Ti))(10( t Tf \u2212Ti )3 \u221215( t Tf \u2212Ti )4 + 6( t Tf \u2212Ti )5) (1) Where Ti is the initial time and Tf is the final one. The geometrical path describing the hand are not time dependent and may be expressed in terms of Euclidean coordinates. An assisted sit-to-stand transfer trajectory follows a path similar to the one shown in Fig. 3. This kind of curve will be defined by a third order polynomial in the (Pi,ePiPf ,e \u22a5 PiPf ) plane (with : e\u22a5PiPf = z\u2227 ePiPf ). Considering the point Pj on the handle path curve: PiPj = UjePiPf +Vje\u22a5PiPf , with : Vj = \u22113 i=0 \u03b1iUi j. The knowledge of the coordinates of the points Pk (k = 0, . . . ,4, Fig.3) leads to the equations below: \u23a7 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa \u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 \u03b10 = 0 \u03b11U1 + \u03b12U2 1 + \u03b13U3 1 = V1 (= dev1) (A) \u03b11U2 + \u03b12U2 2 + \u03b13U3 2 = 0 (B) \u03b11U3 + \u03b12U2 3 + \u03b13U3 3 = V3 (= dev2) (C) \u03b11U4 + \u03b12U2 4 + \u03b13U3 4 = 0 (where U4 = PiPf ) (D)( dV dU ) U1 = \u03b11 + 2\u03b12U1 + 3\u03b13U2 1 = 0 (E) ( dV dU ) U3 = \u03b11 + 2\u03b12U3 + 3\u03b13U2 3 = 0 (F) A solution of this system is obtained in two steps. First, solving the linear system {(A),(B),(C)}, \u03b1i coefficients can be expressed in term of U1,U2,U3. In a second step, U1,U2,U3 values are the solutions of the minimisation problem based on equations {(D),(E),(F)} defined as: min((D)2 +(E)2 +(F)2) under \u23a7 \u23aa\u23aa\u23a8 \u23aa\u23aa\u23a9 \u2212U1 < 0 U1 \u2212U2 < 0 U2 \u2212U3 < 0 U3 \u2212U4 < 0 (2) The starting point of the optimization procedure, is taken such that: \u23a7 \u23aa \u23a8 \u23aa\u23aa\u23a9 \u2212U1 < 2% of the length curve (s(Tf )\u2212 s(Ti)) U1 \u2212U2 < 2% of the length curve U2 \u2212U3 < 2% of the length curve U3 \u2212U4 < 2% of the length curve (3) We have observed that the resulting polynomial is always very close to the experimental curve" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003435_j.mechatronics.2010.04.007-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003435_j.mechatronics.2010.04.007-Figure1-1.png", "caption": "Fig. 1. Conceptual diagram of magnet wheel.", "texts": [ " force whose magnitude and direction can be controlled without influencing on the normal force. The method is verified from the noncontact conveyance experiment of a conductive plate. Differently from an induction motor, a time-varying traveling field in the magnet wheel is generated by mechanical rotation of PMs instead of AC driving. In this chapter, we review the existing principles of thrust force generation using the magnet wheel and suggest how to obtain the thrust force with partial magnetic shield. As mentioned before, when PMs with the direction of pole shown in Fig. 1 rotate below a conductive plate, a repulsive force FN and a traction torque T are generated. However, for the fixed plate, the torque acts as a rotating load of the wheel. A thrust force from the magnet wheel can be generated by various methods such as a tilt of the wheel axis, a partial overlapping, and a use of Halbach array. Firstly, tilting an axis of the wheel frame including PMs, a vector component in a thrust direction of normal force can act as the thrust force [6]. But, the small air\u2013gap length above the magnet wheel, resulting in the small slope angle, limits a magnitude of the thrust force", " However, it is very difficult to apply to the transport path with a small radius of curvature. Fig. 2b shows a vertical magnet wheel adopting PMs arranged in the Halbach pattern [8,9]. Although two axial forces are generated simultaneously for the rotating wheel as in the figure, a magnetically interacting region is limited within the dotted line. Therefore, its forces are very small compared with the axial magnet wheel with the same radius. Furthermore, unlikely with Fig. 2a, it is impossible to control two forces independently. The traction torque acting on a conductive plate in Fig. 1 can be expressed as a sum of tangential forces into the circumferential direction. Therefore, if the rotating magnetic field from PMs is shielded partially, a linear thrust force can be generated in the unshielded region. This concept is described in Fig. 3. In the figure, a magnetic shield plate opened as much as 120 is inserted between PMs and a conductive plate. When PMs rotate in a counterclockwise direction, a linear thrust force is generated into the direction of block arrow. The direction of the thrust force can be changed varying a rotating position of the open area" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001249_09544100g01805-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001249_09544100g01805-Figure1-1.png", "caption": "Fig. 1 Cross-section through ball and races to show forces and angles", "texts": [ " In the general case of bearings with a combined axial and radial loads, each ball has a different load; therefore, it is necessary to carry out a vector sum of these loads to give the total axial and radial loads on the bearing. It is necessary to adjust the axial and radial inner race displacements G01805 # IMechE 2006 Proc. IMechE Vol. 220 Part G: J. Aerospace Engineering at UNIV OF PITTSBURGH on June 18, 2015pig.sagepub.comDownloaded from in order to achieve the required axial and radial loads. This is done in an outer loop iteration, with the individual ball equilibrium being calculated in an inner loop. The analysis and equations are given in reference [6]. The main features are shown in Fig. 1. The transverse radius of the tracks is typically between 1.015 and 1.05 times the radius of the ball. This ratio is known as conformity. Alternatively, it can be defined as a percentage of the ball diameter. The assembled clearance of the bearing is usually chosen to give a static contact angle of 308under light thrust load. Under high-speed running conditions, the ball inertial force Fc has to be reacted by the contact forces Fo and Fi, which means that Fo is greater than Fi and ao is less than ai" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001447_bf00999915-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001447_bf00999915-Figure1-1.png", "caption": "Fig. 1. The construction of the force platform. For explanation see text", "texts": [ " Before mounting the rings on the steel plate the sensit ivi ty to pressure of each ring was adjusted in such a way tha t they were all similar. This was done by applying a standard load to each of the four rings separately and then trying out across which diameter the load would cause identical deflections on the ink-writer. During this procedure the strain gauges remained connected as a bridge. The appropriate diameter was marked and a hole was drilled through the ring. Four t0 x 50 mm bolts were used for mounting the rings under the steel plate as shown in Fig. 1. The rings with the strain gauges placed in the same way were placed in diagonally opposite corners of the platform. I f a vertical pressure was applied to the platform, e. g. by letting a subject stand on one corner of the platform, only one ring carried the load and the registration system showed a deflection equal to his body weight. Standing on the middle did not change the response of the system, because in this position each of the rings carried one quarter of his body weight. By having the subject move around on the platform it could be shown tha t the sensitivity of the platform was independent of where the force was applied within the surface" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001554_eej.20585-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001554_eej.20585-Figure4-1.png", "caption": "Fig. 4. Space vector diagrams at power generation (s12 < 0).", "texts": [ " Consequently, the position of the coil when the primary and secondary current for each symmetric component is at a maximum and the position of the rotating magnetic field can be estimated for any instant in which the PMIG is rotating at a constant speed. For the space vector diagrams for the positive-sequence and negative-sequence components, if attention is given to a particular instant during operation even though these components are rotating in opposing directions, then Figs. 4(a) and 4(b) can be superimposed to produce Fig. 4(c). Here, if we consider the instant at which I . u reaches a maximum, then based on assumption (1) in Section 3.2, the PM rotor is rotating in synchronization with the positivesequence component rotating magnetic field. As a result, the position of the PM rotor can be estimated based on the position of \u03c6 . gp. Moreover, the position of the squirrel-cage rotor can be estimated based on the position at which I . 2p reaches a maximum. 3.4.2 Calculation of the current distribution The spatial distribution of the secondary current gives the sinusoidal current distribution centered on the coil position at which I . 2p and I . 2n reach a maximum, and combining these can be taken to represent the spatial distribution for the secondary current. As a result, the coil position at which the secondary current reaches a maximum at the instant that I . u reaches a maximum is the position deter- mined based on the combination of I . 2p and I . 2n, as shown in Fig. 4(c). 3.4.3 Performing 2D-FEA Based on the above input information (specifically the position of the squirrel-cage rotor and the PM rotor, and the spatial distribution of the primary and secondary current), the analytical model is created and 2D-FEA is performed as a static magnetic field. Then, the internal magnetic field in the PMIG can be visualized and quantified when the voltage is unbalanced. In this section the authors first demonstrate the validity of the analytical method given in the previous section using experiments involving a test PMIG [2\u20135] at 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure10-1.png", "caption": "Fig. 10. Hyperboloidal-type normal circular-arc gear drive.", "texts": [ " (15) yields the equations of tooth flanks for hyperboloidal-type drive in respective coordinate systems: R\u00f0j\u00deb : R\u00f0j\u00de \u00bc rje\u00f0kj\u00de \u00fe rjkj sin dj\u00f0 sin dje1\u00f0kj\u00de \u00fe cos djkj\u00de \u00fe rj\u00f0e\u00f0j\u00de2 cos h\u00fe e \u00f0j\u00de 3 sin h\u00de \u00f0j \u00bc 1; 2\u00de \u00f055\u00de In engineering application, it must be pointed out that the directrixes of conjugate tooth flanks R\u00f0j\u00deb \u00f0j \u00bc 1; 2\u00de mentioned above are determined by the line C0 of action, which satisfies noninterference condition, but another line C\u00f0 \u00de0 of action does not satisfy noninterference condition and brings gear tooth interference, only a half of two datum surfaces R\u00f01\u00dep and R\u00f02\u00dep can be adopted (shown in Fig. 10). Gear 1 chooses lefthalf of R\u00f01\u00dep and gear 2 chooses down-half of R\u00f02\u00dep , which is similar to bevel gear for crossed-axis drive. In Fig. 10, R\u00f01\u00deb and R\u00f02\u00deb are the sketch of a pair of conjugate tooth flanks, Cm is an actual line of action, and semi-circle corresponding to the conjugate point P is an instantaneous contact line. When two gears rotate at constant speed ratio, instantaneous contact line moves along the line of action from one gear face to another at constant speed. Such a property of instantaneous contact line has been discussed for parallel axis and intersected-axis drive, which shows that this property is common to normal circular-arc gear drive" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001470_app.1979.070240608-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001470_app.1979.070240608-Figure4-1.png", "caption": "Fig. 4. Load-extension curves for the nylon models: (a) no twist; (2) 0.10; (3) 0.21; (4) 0:31; (5) 0.38; (6) and 0.41 radian/mm.", "texts": [ " Several of these models with equal S and Z twists were made but with each having a different amount of twist. This gives models with different convolution angles, which would correspond to different varieties of convoluted cotton fibers, each having a different convolution angle. Other strips were twisted and set with more than one reversal, with unequal lengths of S and Z twist, and with a different convolution angle between the two types of twist. The models were then extended with an Instron tensile tester. The results are shown in Figure 4, and as can be seen from the graph, the strip with the greatest amount of twist has the lowest initial modulus and is more extensible. There is thus an analogy between the twisted nylon strips and cotton fibers. Cotton fibers having a high convolution angle will have a lower initial modulus but will be more extensible, while those with a low convolution angle will have a high initial modulus and be less extensible. If we look back at the x-ray work, in the light of Hebert\u2019s5 and Ingram\u2019# findings, we now can explain the correlations which have been established" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001027_icit.2005.1600634-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001027_icit.2005.1600634-Figure2-1.png", "caption": "Fig. 2. Representation of the skew", "texts": [ " Defining F the MMF distribution in the air-gap due to the current iAi flowing in an arbitrary coil Ai, the elementary flux corresponding in the air-gap is measured in comparison to an elementary volume of section ds and length go such as do= poFg0-lds (9) (13) I LBA (Xr) = JLA(Z xr)dz 0 In the same way as [1], and according to the manner of connections of the coils translated by the sign in (14), this inductance can be obtained by summing all mutual inductances between the q and p coils of winding A and B respectively, such as q p LBA(Xr) = I \u00b1LBjAi(Xr) i=l j=l (14) B. Bars skewing Figure 2 depicts the crossing of a rotor loop rj under the field of a stator coil Ai. The skew is written thanks to the definition of z(x) (10) which will be a function describing the uniform skew, or particularly, the case of spiral skew. We can notice that the pitch aAi of the coil Ai is defined in comparison to its sides placed at xli = r f ji and 12i = r (02i and that the effect of linear rise of MMF across the slot is note represented in this figure. C. Slot opening Let us examine the case of coil Ai with WAi turns placed in slots which can present an opening of the width Q according to the configuration considered" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000513_2004-01-1058-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000513_2004-01-1058-Figure5-1.png", "caption": "Figure 5: Electric Motor Controlled Type TMD", "texts": [ " This ball and ramp mechanism acts on the main clutch pack thus providing the required locking torque. Other systems in the market are activated hydraulically. Typically these systems contain a pump, which creates hydraulic pressure, under speed difference, that acts on a conventional wet clutch pack. By electronically altering this pressure via solenoids or other valves, the locking torque of the hydraulically controlled type TMD is controlled. Another system is the electric motor controlled type TMD. An example for this is shown in Figure 5. Here the torque supplied by an electric motor is transformed into a longitudinal force via a ball and ramp or cam mechanism, which activates a wet clutch pack. This means that the locking torque of the TMD is controlled by altering the electric motor current. Additional information with respect to this type of electronically controlled device can be found in [1]. Other systems like torque-vectoring devices will not be explained in this paper. On surfaces with different friction coefficients \u00b5 on the left and the right side of the vehicle (split-\u00b5) open axle differentials have the disadvantage that they can only transmit the same (low) amount of torque to the high-\u00b5 side which is transmitted to the low-\u00b5 side" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002171_j.triboint.2009.05.009-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002171_j.triboint.2009.05.009-Figure3-1.png", "caption": "Fig. 3. Journal bearing geometry. Left: front view; right: side view\u2014unwrapped bearing.", "texts": [ " The equations of motion may be written in a matrix form as in (7), where the vector b includes the unknown reaction forces, reaction moments and accelerations of the system: A b \u00bc c (7) where b \u00bc ff Bx ; f By ; f Bz ;Ny;Nz; f Ax ; f Ay ; f Az ; f Cz ; \u20acy; \u20acxB; \u20aca; \u20acxc; \u20acycg T The equations that govern the dynamics of the fluid films in a dynamically loaded journal bearing, for the conventional hydrodynamic lubrication case and for the active lubrication case, are presented. The main geometrical relations of a fluid film journal bearing are shown in Fig. 3. The governing equation for the pressure distribution of the oil film in journal bearings can be obtained from the standard reduced form of the Reynolds equation for elastohydrodynamic lubrication [15], given by @ @x rh3 12m @p @x ! \u00fe @ @z rh3 12m @p @z ! \u00bc @ @x rh\u00f0Uj Ub\u00de 2 \u00fe @\u00f0rh\u00de @t (8) Eq. (8) describes the flow in the journal bearing in the domain 0 x 2prb and lb=2 z lb=2. For the upper bearing of the compressor: Uj \u00bc Orb and Ub \u00bc 0. The fluid film thickness measured from j \u00bc 0 can be calculated using: h \u00bc cb\u00f01\u00fe cosj\u00de, where j is the angle measured from the location of the maximum fluid film thickness, as shown in Fig. 3. Thus, using the coordinate transformation x \u00bc rbj in Eq. (8) and assuming that the fluid is isoviscous, Newtonian and incompressible, the Reynolds equation can be rewritten as @ @j h3 m @p @j ! \u00fe r2 b @ @z h3 m @p @z ! \u00bc 12cbr2 b _ cosj\u00fe sinj _f O 2 (9) In dynamically loaded bearings the eccentricity and attitude angle varies through the loading cycle. Therefore, as seen in Eq. (9), the pressure distribution may be only determined if the normal squeeze velocity (_ ) and the rotational velocities ( _f and O) are known at any time during each cycle" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000544_j.cad.2005.02.009-Figure11-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000544_j.cad.2005.02.009-Figure11-1.png", "caption": "Fig. 11. Innovated ball bearing.", "texts": [ " The fitness of a chromosome is a function of these outputs, and can be expressed as follows: F Z k1A Ck2B Ck3C (1) where k1, k2, and k3 are constants and used to adjust the weights of A\u2013C. But, their sum should be equal to 100%. Genetic Algorithms were applied to evolve the algorithmic chromosome. After many generations, the algorithms converge to the best algorithmic chromosome. According to the best-evolved algorithmic chromosome, the chromosome of a new ball bearing can be obtained, based on which a new ball bearing was redesigned as shown in Fig. 11. This innovative ball bearing has one row of balls, a retaining cage, an outer ring, and a pair of separable inner rings. Since there is no need to use second row of balls or another ball bearing to carry the thrust load in another direction, its cost is lower. The cross sections of the grooves on both inner and outer rings have two symmetrical arcs, the centers of which are located on two diagonal contact lines, respectively, as shown in Fig. 11, and the radii of the arcs are larger than the radii of balls. The clearances between raceway and balls can be adjusted accurately by grinding the inner face in one of the two inner rings, so that the contact angle between balls and raceway will not be changed much when load is varying, and the bearing can thus adapt to a varying working load. Since the inner rings can be separated while the balls are filled, this ball bearing can thus contain more balls so that it can carries not only larger radical load but also larger bi-directional thrust loads" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002027_978-1-4020-4535-6_27-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002027_978-1-4020-4535-6_27-Figure1-1.png", "caption": "Figure 1. Axial flux surface mounted PM machine for field weakening application.", "texts": [ " These analyses demonstrate that airgap flux can be commanded with an appropriate armature reaction control. As a result speed range can be increased without significant requirement of the stator d-axis current. In addition a procedure to estimate the d-q parameters based on armature reaction waveform analysis is used. Geometry and iron to magnet ratio define the machine reactances. In the following sections description of the proposed machine, FEA for no-load and load conditions, and parameters procedure calculation are shown. Description The AFPM machine topology proposed is shown in Fig. 1. This machine is composed of two rotors and one central stator. Rotors are north-north (NN) PM surface mounted type containing the excitation poles. Each of these poles is assembled by two parts: PM and iron piece. PM section is a magnet part axially magnetized which provides excitation to the machine. On the other hand an Iron section which offers an easy path for the stator current armature reaction. Due to the short airgap length, the total flux per pole can be considered as two components: one associated to the magnets (high reluctance), the other associated to iron (low reluctance)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001925_0954406jmes476-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001925_0954406jmes476-Figure3-1.png", "caption": "Fig. 3 Scheme of the contact ellipse in the raceway", "texts": [ " Total heat generated in the ball bearing is Ht = z\u2211 j=1 (H1mj + H2mj + Hmsj + Hdj + Hcj) + HCL (15) Heat generated by sliding due to gyroscopic motion, which has been reduced greatly in the modern ball bearing, is ignored. Heat generated by ball sliding in the cage pocket is also omitted because Qcj is very small [12]. A code based on the quasi-static method is developed with FORTRAN. Figure 2 shows the program flow chart of the code. When the geometry, load, material, and lubricant parameters are read-in, the kinematics result and the heat generated can be obtained as output. Figure 3 shows how the heat source moves in the raceway. The contact ellipse between the ball and Proc. IMechE Vol. 221 Part C: J. Mechanical Engineering Science JMES476 \u00a9 IMechE 2007 at University of Bath - The Library on June 11, 2015pic.sagepub.comDownloaded from the raceway is displayed in Fig. 3(a), and the process in which the contact ellipse moves in the raceway is displayed in Fig. 3(b). The ball is on the dashed ellipse when time is ti. After a time interval ti+1 \u2212 ti, the ball moves onto the real line ellipse and the next ball moves onto the dashed ellipse. Simultaneously, the heat generated on the contact ellipse becomes high. Hence, the raceways are actually heated by moving heat sources, and the temperature in the rings can be obtained using a step-by-step computation of the heating and cooling on the contact ellipses. To save CPU time, the model must be simplified. Transient temperature in the ball bearing is assumed axisymmetrical because the rotational speed is very high" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002061_6.2007-6525-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002061_6.2007-6525-Figure1-1.png", "caption": "Figure 1. Wind tunnel model", "texts": [ " Model Description A semi-span model of a SensorCraft concept wing was originally tested in the NASA Langley Transonic Dynamics Tunnel (TDT) in November 20042. The current model and subject of this paper is very similar to the one described in Ref. 2 except for the way it is mounted in the wind-tunnel. The model consists of a rigid inboard spar with fiberglass skin and a flexible main aluminum spar shaped to reflect the structural bending characteristic of a full-scale concept wing. The wing, pictured in figure 1, has four evenly spaced trailing edge control surfaces and one leading edge surface near the tip. The model has an eleven foot span and is attached to the wall of the TDT test section using the rail based \u201cPitch And Plunge Apparatus (PAPA),\u201d depicted in figure 2, that affords 2 DOF motion: pitch and plunge. The wind-tunnel model is instrumented with accelerometers along the spar, strain gauges at the root and mid-spar, a rate gyro at the wing tip, a gust sensor vane in front of the wing, and a rate gyro and accelerometers at the tunnel attachment point" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002015_13506501jet310-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002015_13506501jet310-Figure1-1.png", "caption": "Fig. 1 Cross-section configuration of a short journal bearing at the mid-plane z\u2217 = 0", "texts": [ " [22], the effects of isotropic surface roughness on the weakly non-linear bifurcation behaviour in the vicinity of the linear stability boundary of short journal bearings is analysed in the present study. To show the effects of isotropic surface roughness on the system, bearing characteristics such as the steady-state performance, linear dynamic characteristics, sub-critical, and super-critical bifurcation regions, limit cycles and phase plane portraits are presented with various values of the surface roughness stress parameter and the system parameter. Figure 1 shows the physical configuration of the crosssection at the mid-plane z\u2217 = 0 of a short journal bearing considering the effects of surface roughness. The journal of radius R\u2217 is rotating with angular velocity \u03c9\u2217 within the bearing shell. Assume that the lubricant flow in the bearing is isothermal, incompressible, and laminar, and the thin film lubrication theory is applicable. Then the dynamic Reynolds-type equation governing the local film pressure P\u2217 is obtained from Pinkus and Sternlicht [1] \u2202 \u2202z\u2217 ( H \u22173 \u2202P\u2217 \u2202z\u2217 ) = 6\u03bc ( \u03c9\u2217 \u2212 2 d\u03d5 dt\u2217 ) \u2202H \u2217 \u2202\u03b8 + 12\u03bc \u2202H \u2217 \u2202t\u2217 (1) where z\u2217 is the axial coordinate, \u03b8 the circumferential coordinate, t\u2217 the time, \u03d5 the attitude angle, and \u03bc the lubricant viscosity" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003388_1.4001621-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003388_1.4001621-Figure1-1.png", "caption": "Fig. 1 Model schematic", "texts": [ " More recently, Nassar and Yang 11 developed a nonlinear mathematical model to predict the selfloosening of threaded fasteners. Their formulation took into account the three dimensional 3D geometry of the threads, and various pressure distribution scenarios on the contact surfaces under a hexagonal bolt head and between engaged threads. This work extends the hexagonal head model presented in Ref. 11 in order to investigate the loosening behavior of bolts with countersunk-head configuration. 2 Model Formulation Figure 1 shows a schematic of the investigated bolted joint model, where the countersunk bolt threads are engaged into a tapped hole in the bottom block that is constrained in the lateral direction. The model is based on force and moment, as well as the kinematic relationships during the transverse motion of the upper plate. The angular equation of motion is numerically integrated with a sufficiently small time step to provide the incremental loosening rotation of the bolt, which is, in turn, correlated with the incremental drop in the bolt tension using a MATLAB code", " A similar procedure is ollowed for calculating the thread frictional torque Tt component 11 . The pitch torque Tp component is given by 3 Experimental Procedure and Test Setup In this section, an experimental procedure and test setup are established in order to validate the analytical results obtained from the proposed self-loosening model. A self-loosening testing machine that is similar to the Junker machine 9 is used. The machine consists of a motor that drives a set of pulleys with an eccentric mechanism that reciprocates the upper plate shown in the bolted joint schematic in Fig. 1 in order to cause self-loosening. The tests are displacement controlled to apply a predetermined transverse excitation of amplitude 0.03 and frequency of 10 Hz. An embedded load cell is used to monitor the loss in the joint clamp load/bolt tension due to loosening in real time during the test. The loosening data are curve-fitted using a second order curve in the range Finitial F Fmin. A minimum load Fmin is chosen to be well outside the measurement error of the load cell e.g., Fmin=2 kN, which is 10\u201320% of Fmax ", " 2 h height of the countersunk fastener head m ratio of the change ri / re due to bolt head sliding n\u0302 unit vector perpendicular to the contact surface under the bolt head p thread pitch q1 contact pressure acting on the lower portion of the bolt head surface q2 contact pressure acting on the upper portion of the bolt head surface qto average thread pressure qt increment pressure amplitude caused by the bending effect q contact pressure variation as a function of the radial location re outer contact radius under the bolt head Fig. 1 ri inner contact radius under the bolt head Fig. 1 rmaj maximum thread contact radius rmin minimum thread contact radius rx variable radius of the bolt head; can stand for ri or re vtx velocity of any point on the thread surface along the x-direction with respect to the joint surface v general velocity vector of any point under the bolt head with respect to the joint surface vO velocity of point O with respect to the joint surface vOQ relative velocity of point Q with respect to reference point O h height of a small area dS upon which the under head pressure acts xA displacement of point A in the x-direction due to bolt head sliding ournal of Tribology om: http://tribology", "org/ on 01/28/2016 Terms xB displacement of point B in the x-direction due to bolt head sliding xO displacement of point O in the x-direction due to bolt head sliding zO displacement of point O in the z-direction due to bolt head sliding r loosening angle of the bolt head half the thread profile angle lead helix angle t ratio of major-to-minor thread radii T transverse displacement of the joint upper plate b ratio of the translational to rotational velocity of the bolt head t ratio of the translational-to-rotational velocity of the thread surface angular location bending stiffness of the bolt head b coefficient of bearing friction t coefficient of thread friction bending angle of the bolt head due to external excitation half of the bolt under head complementary cone angle, as shown in Fig. 1 a constant angle that depend on the bolt head geometry 1 relative angular velocity of the bolt under head surface with respect to the joint bearing surface about the y-axis b relative angular velocity of the bolt under head surface with respect to the joint bearing surface about the z-axis t relative angular velocity of the thread surface with respect to the joint thread surface \u0307 self-loosening angular acceleration of the bolt References 1 Junker, G. H., 1969, \u201cNew Criteria for Self-Loosening of Fasteners Under Vibration,\u201d SAE Trans" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.51-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.51-1.png", "caption": "Fig. 14.51 Built-in motor-type electric immersion vibrator", "texts": [ " Pendulum-type immersion vibrators are driven by an electric motor or a combustion engine and a flexible shaft to which a vibration generator rolling on a raceway is attached. Part B 1 4 .3 The diameters of the vibration heads of the vibrators usually range from 25 to 70 mm and the frequencies generated are 300\u2013200 Hz, respectively. Modern pendulum-type immersion vibrators with a flexible shaft are driven by single-phase commutator motors; under load they can generate frequencies as high as 200 Hz (Fig. 14.50). The diameters of vibration heads driven by commutator motors are usually 25\u201365 mm. Built-in motor type electric immersion vibrators (Fig. 14.51) usually operate in conjunction with a voltage and frequency converter supplying a voltage of 42 V at 200 Hz. They can also be supplied from generating sets with an appropriate rated frequency. Because of the considerable permissible length of the power lead from the generator to the vibration head (about 15 m) these vibrators are suitable for compacting high elements. The vibrator\u2019s vibration heads are usually 35\u201385 mm in diameter and the vibration frequency is 200 Hz. Pneumatic immersion vibrators are made with vibration heads that are 25\u2013140 mm in diameter" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002821_j.sna.2010.07.002-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002821_j.sna.2010.07.002-Figure7-1.png", "caption": "Fig. 7. Tangential forces acting on a mass element of the frictional layer in down portion of roller.", "texts": [ " 11) Ndown = b \u222b L \u2212L p(x) dx (11) here FNdown is the normal force acting on down portion of the oller. The tangential velocity vs(x) of stator surface in the spatial xed frame (x\u0303, z\u0303) is obtained as follows: s(x) = U\u0307x(x) = \u2212aA ( 2 )2 vw cos ( 2 x ) (12) n the slip zone, the tangential stress depends on the frictional layer elocity vr(x) and the stator tangential velocity vs(x) according to ollowing relation: (x) = \u2212sign(vr(x) \u2212 vs(x)) p(x) (13) Horizontal motion of frictional layer Uk x (x) is considered in movng frame (x, z). According to Fig. 7 Newton\u2019s second low can be pplied for tangential element of frictional layer in down portion f roller dxU\u0308k x = (x)bdx \u2212 cT dxUk x (x) \u2212 dT dxU\u0307k x (x) (14) Rearranging Eq. (14) results: vw 2Uk\u2032\u2032 x (x) \u2212 dT vwUk\u2032 x + cT Uk x (x) = b (x) (15) For tangential displacement, ()\u2032 denotes derivation with respect o x. The solution of Eq. (15) is obtained as follows: k x (x) = Uk x h(x) + Uk x p(x) (16) k x h(x) = C1e \u00af\u0328 x cos( \u00af\u030c x) + C2e \u00af\u0328 x sin( \u00af\u030c x) (17) k x (x) = B1 cos ( 2 x ) + B2 sin ( 2 x ) + B3 cos ( 2 L ) (18) The tangential frictional layer velocity vr(x) within the spatial xed frame reference (x\u0303, z\u0303) is obtained as follow: r(x) = vR \u2212 vwUk\u2032 x (x) (19) here VR is linear velocity of the roller in x direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002155_j.msea.2009.10.058-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002155_j.msea.2009.10.058-Figure10-1.png", "caption": "Fig. 10. Simulated transient thermal stress (Pa) profiles for group R1, (a) X component, (b) Y component, (c) Z component.", "texts": [ " With higher laser sintering temperature, lower bending strength of the final ceramic samples has been achieved. 4.2. Stress distribution 4.2.1. Transient thermal stress The first constraint condition, element constraint in one of the bottom corners, has been used to simulate the transient thermal 1 d Engineering A 527 (2010) 1695\u20131703 s s c i Y t s d X a t v 2 i t a i t t i d d a s 4 p o b t c p a d r u r t b t l t P s a i s y d g t o T 2 i i 1 t e t m s d 700 X. Tian et al. / Materials Science an tress. The X and Y components of transient thermal stress are hown in Fig. 10. The maximum compressive stresses for three omponents simultaneously appear in the center of laser sinterng zone. The maximum value of the compressive stresses for X, and Z components are 308 MPa, 332 MPa, and 249 MPa, respecively. Stress distributions for X and Y components in the cross ection are unsymmetrical due to the asymmetry of temperature istribution as well as temperature gradient. Tensile stresses for and Y component are in the bottom area of the cross section, s shown in Fig. 10a and b, which are produced by balancing the hermal expansion of material in the sintering area. The maximum alues for tensile stress of X and Y components are 26.1 MPa and 1.0 MPa, respectively. The distribution of the Z component stress s different from X and Y component. Tensile stress in Z direcion is present under the laser sintering zone, as shown in Fig. 10c nd the maximum is 87.7 MPa which is much higher than those n x and y directions. Transient thermal stresses are mainly conrolled by the temperature gradient. It can be observed in Fig. 8 hat the temperature gradient in z direction is larger than those n x and y directions. Moreover, the temperature gradient in x irection is larger than y direction which is the laser scanning irection because higher residual heat in this direction causes small temperature gradient, and consequently small thermal tress" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002904_1077546309104878-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002904_1077546309104878-Figure5-1.png", "caption": "Figure 5. Oscillation of mass mo.", "texts": [ "comDownloaded from Pimp 1 mo cos (23) where is a coefficient that expresses the energy remaining after the impact, is the velocity of the mass mo at the instant of impact, and is the angle at which the mass mo strikes the plate. The above velocity is the geometrical resultant of the following three ones: 1. The velocity o of the mass at the end of its flight at t t x x . 2. We observe that the mass M follows the orbit determined above in Section 2.1.2, while the mass mo, due to the system of the spring/damper, oscillates with respect to M (see Figure 5). The above-mentioned oscillation is describedby the following equation: mo m ko m co m 0 (24) which has the solution m t e ot D1 sin ot D2 cos ot (25) where o o and o are given by equations 5 and 6. At instant t (loss of touch), the normal distance between M and mo changes to m and the following equation is valid (see Figure 6): z m o or, finally, m o z (26) Thus, the initial conditions for the oscillation of mo are m 0 x t o x z t m 0 x t z t (27) which finally give at SIMON FRASER LIBRARY on November 17, 2014jvc" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001920_1.2988480-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001920_1.2988480-Figure2-1.png", "caption": "Fig. 2 Tooth surface of the quasi-complementary crown gear", "texts": [ " Figure 1 shows three coordinate systems: O-xyz, -xgygzg, and O-XYZ. O-xyz and O-xgygzg are attached to the uasi-complementary crown gear and the generated gear, respecively. O-XYZ is fixed in space. Origin O is the machine center. Y plane is the cradle plane. z is the cradle axis, which is equivaent to the axis of the quasi-complementary crown gear. zg is the xis of the generated gear and is inclined at an angle r to the Y xis. yg axis is omitted. O-xyz coincides with O-XYZ when the otation angle of the quasi-complementary crown gear is zero. As shown in Fig. 2, the left tooth surface x of the quasiomplementary crown gear and its unit normal n in O-xyz are epresented by x u,v = \u2212 x u u + v \u2212 sin 0 0 cos 11001-2 / Vol. 131, JANUARY 2009 om: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/17/201 n = cos 0 sin 1 where is the pressure angle of the tool, is the deviation of the tooth trace, namely, the tooth angle error, and u and v are the parameters expressing the tooth surface. x u is involved in the crowning on the tooth surface. x u is not shown in Fig. 2 because it is very small. n can be considered that there is no effect of x u . When the crowning is not given, the tooth surface of the quasicomplementary crown gear is a plane, namely, an octoid tooth profile. The quasi-complementary crown gear is rotated about the z axis by angle and generates the tooth surface of the straight bevel gear, which is rotated about its axis by cos f /sin 0. We call this rotation angle of the quasi-complementary crown gear the generating angle. When the generating angle is , the coordinate system O-xyz is rotated about the z axis by in the coordinate system O-XYZ", " In addition, a complicated grid model is not needed beause the real values related to gear accuracy can be detected sing the coordinates of arbitrary points. 3.1 Outline of Measurement Method. The aim of the meaurement of the straight bevel gear accuracy is fundamentally to ompare the measured real tooth surface with the theoretical one. he coordinate measurement of the real gear tooth surface proides the information about some factors related to the gear accuacy mentioned in the previous section. We consider the pressure ngle , tooth angle error as shown in Fig. 2, workpiece setting ngle r as shown in Fig. 1, and apex to back l as shown in Fig. 3 s the factors related to the gear accuracy in this paper. These actors are invariable because these factors do not change during he end milling process. For measurement, the straight bevel gear is set arbitrary on a MM. The positions of the gear axis and the datum plane must be etermined by measurement independent of the tooth surface easurement because the gear is set arbitrarily. We can make the rigin O and gear axis zg in the coordinate system O-xgygzg atached to the gear coincide with the origin Om and axis zm in the oordinate system Om-xmymzm of the CMM, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003643_aim.2011.6027000-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003643_aim.2011.6027000-Figure3-1.png", "caption": "Fig. 3. Appearance of pressurization", "texts": [ " Overview of the Mechanism The mechanism consists of many units; each unit imitates an intestinal circular muscle. Fig. 2 shows a cross section of a unit. Each unit consists of a cylindrical tube, a straight-fiber-type artificial muscle, and flanges. The cylindrical tube is arranged inside the artificial muscle, and the two ends are joined by the flanges. The space enclosed by the artificial muscle, the cylindrical tube, and the flanges forms a chamber. This chamber can be pressurized through an air vent. Fig. 3 shows how the appearance of the unit changes when the unit is pressurized. Both the artificial muscle and the cylindrical tube expand during pressurization; the artificial muscle expands only in the radial direction, and the unit contracts in the axial direction. The cylindrical tube includes four fibers that become constrained when pressurizing the unit. Therefore, fold lines occur on the cylindrical tube in four directions, and the expansion is divided into four parts. As seen from the axial direction, upon expansion, the cylindrical tube takes the shape of four quarter circles that push out toward the center" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000566_ac00245a028-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000566_ac00245a028-Figure1-1.png", "caption": "Figure 1. A schematic representation of the cell. See text scription.", "texts": [ " Even though small substrates could conceivably be introduced into this thin layer through the ion exchange membrane, there are no obvious means to rapidly determine the activity of an entrapped enzyme in this system. The method for control of solution potential described here also utilizes a permeable, thin membrane to separate the thin-layer control electrode from the external bulk solution in the cell. The thin layer itself is very thin (0.1-3.0 Km) and contains an entrapped oxidase enzyme, free or immobilized, and the mediator titrant. Figure 1 is a schematic representation of the cell and the reactions which take place in the case of galactose oxidase. The two working electrodes used in the cell are shown; the platinum working electrode (Pt) is used for the amperometric determination of HzOz and the gold control electrode for the control of solution potential within the enzyme layer. Control of the enzyme oxidation state (Cu) is therefore mediated by the ferricyanide-ferrocyanide couple. Thus, by measuring the steady-state hydrogen peroxide production with the platinum electrode behind this membrane assembly, the relative steady-state activity of an oxidase enzyme in the thin layer can be determined in the presence of a substrate which is introduced into the bulk solution and allowed to equilibrate across the outer separator membrane" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002410_s00248-008-9468-6-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002410_s00248-008-9468-6-Figure4-1.png", "caption": "Figure 4 Example simulation trajectory of the model bacterial cell, projected onto a plane. The algal cell swims along the x-axis", "texts": [ " We then calculate the \u2018tracking fraction\u2019 S(t\u2217), defined as the fraction of simulations in which the bacterial cell is still tracking the algal cell at dimensionless time t\u2217. (In statistical jargon, S(t\u2217) would be called the \u2018empirical survival function\u2019, hence the symbol S.) The standard error SE(t\u2217) of S(t\u2217) is estimated by SE(t\u2217) = \u221a S(t\u2217)[1 \u2212 S(t\u2217)]/N, (8) where N is the number of simulations [26]. We shall first present data for the default set of parameter values listed at the end of the previous section and later look at the effect of altering the parameter values. Figure 4 shows an example trajectory of the model bacterial cell, projected onto a plane, for a simulation run with the default parameter values. The trajectory bears similarities to the experimental results shown in Fig. 1, especially in that the primary motion of the bacterial cell is almost at right angles to the motion of the algal cell, so that the bacterial cell\u2019s motion is a zigzag. This zigzag motion is not a once-off but occurs in most of the simulations, as revealed by the probability density function for the angle \u03c6 between the bacterial and algal swimming directions (panel A of Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003013_jeb.033639-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003013_jeb.033639-Figure1-1.png", "caption": "Fig. 1. General anatomy of the long (TrLONG) and lateral (TrLAT) heads of the triceps brachii of the forelimb of goats. Also shown are the approximate locations of sonomicrometric crystals and EMG electrodes used to measure muscle function. See the text for a more detailed description of the anatomy and methods.", "texts": [ " A relatively large, long-fibered condition in a multiarticular muscle may reflect a functional requirement for fascicle shortening and work production, rather than strut-like isometric force production (Payne et al., 2005a; Payne et al., 2005b). Additionally, if a multiarticular muscle is more massive than those proximal to it, it would not make sense for it to transfer energy from smaller muscles without performing work itself. The long head of the triceps brachii (TrLONG) is a flexor at the shoulder joint (i.e. it rotates the humerus caudally in the sagittal plane; Fig.1). Consequently, it is able to transfer energy from shoulder extension to elbow extension in the manner consistent with the strut-like function described above. However, morphological properties of the TrLONG suggest a role in work production less commonly ascribed to biarticular muscles. First, it is relatively long fibered (Payne et al., 2005b), an architecture associated with work or displacement producing muscles (Lieber, 2002). Second, it is massive relative to other intrinsic and extrinsic forelimb musculature (Payne et al", " By contrast, Witte et al. (Witte et al., 2002) (and references therein) assume an isometric role for the TrLONG in the early phase of stance in small mammals. We measured aspects of in vivo muscle function including activation, strain, and strain rate in the TrLONG of goats (Capra hircus Linnaeus 1758) to broaden understanding of the functional repertoires of biarticular muscles. The functional role of the TrLONG was also compared with its monoarticular agonist, the lateral head of the triceps (TrLAT; Fig.1). Based on a previous comparison of both muscles during jumping (Carroll et al., 2008), we predicted that the TrLAT would exhibit a stretch shortening pattern in parallel with the flexion and extension of the elbow, whereas the TrLONG would shorten throughout stance in a pattern similar to that predicted by Goslow et al. (Goslow et al., 1981) for running dogs. Changes in muscle force, work, and mechanical power associated with changes in speed that are produced by the musculoskeletal system are linked to the energy cost of locomotion (or rate of oxygen uptake; VO2) (e", "5m long and 0.75m wide). During data collection an effort was made to record similar speeds (usually at 0.5ms\u20131 increments) across animals; however, we also sought to elicit the fastest or slowest speed at which an animal would perform a certain gait. Thus, minimum and maximum speeds within a gait differed among some individuals. Values are reported as means \u00b1 s.e.m., unless otherwise noted. The long head (TrLONG) and the lateral head (TrLAT) of the triceps both insert on the olecranon process of the ulna (Fig.1). A much smaller (<5% mass of combined TrLONG and TrLAT) monoarticular medial head is also present but was not measured. The lateral head originates from the humerus, whereas the long head originates from the distal-most third of the ventro-caudal margin of the scapula. Consequently, the TrLAT is monoarticular and the TrLONG is biarticular. Architecturally, the TrLAT contains parallel fascicles that run the extent of its length (ranging from an average of 6.0\u00b12cm in a 25kg goat to 8.0\u00b13cm, in a 45kg goat)", " Digitized data were filtered at 30Hz with a fourth-order recursive (zero-lag) Butterworth filter, and used to determine hoof position, limb segment position, and elbow and shoulder angles for each sequence. The camera was controlled by an analog trigger, and the voltage signal from this trigger was used to synchronize camera recordings with EMG and sonomicrometry recordings. Shoulder flexion was defined with respect to the caudal (posterior) angle formed between the humerus and scapula in the sagittal plane (Fig.1). X-ray video kinematics were captured from a lateral X-ray video, recorded at 125Hz, of a 15kg goat walking and trotting on a motorized treadmill (0.7 m 1.20 m) at 0.9 and 2.5 m s\u20131, respectively. A smaller goat was required for X-ray recordings, in order to fit the scapula, humerus and olecranon on the 30.5cm diameter fluoroscopic screen for a full stride. Videos were recorded at 125Hz with a high-speed Photron Fastcam camera (Photron, San Diego, CA, USA) mounted on the fluoroscope of the X-ray C-arm (Model 9400 C-Arms International, Hamburg, PA, USA)", " In contrast to the shortening observed in the biarticular goat TrLONG, hind limb biarticular muscles in running turkeys (Roberts et al., 1997) and guinea fowl (Daley and Biewener, 2003), hopping wallabies (Biewener et al., 1998) and walking and running humans (Ishikawa et al., 2005a; Lichtwark et al., 2007) display nearly isometric muscle function. Consequently, differences in their contractile behavior probably reflect other anatomical differences. First, whereas the TrLONG has a long moment arm at the shoulder relative to its moment arm at the elbow (Fig.1), the proximal moment arms of the hind limb lateral and medial gastrocnemius at the knee are small relative to the muscles\u2019 moment arm at the ankle, reducing the influence of the knee joint on LG and MG length change behavior (Kaya et al., 2005). Second, fascicle length changes of the LG and MG are buffered from ankle joint kinematics by their comparatively long compliant tendon (Lichtwark and Wilson, 2006). By contrast, the tendon attachment of TrLONG at the elbow is extremely short, so that TrLONG fascicle length changes are more closely coupled to joint kinematics" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002090_9780470264003-Figure10.6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002090_9780470264003-Figure10.6-1.png", "caption": "Figure 10.6 Frying pan TRIZ separation example. (From Shpakovsky and Novitskaya, 2002b.)", "texts": [ " Over the years, one improvement proposed replaced the conventional conical shape of the board with a toric. A chop is pushed up on the board surface and then turns over following the geometry. However, there is a contradiction. It is very difficult to extract the chop from the frying pan because of the board\u2019s shape, but it is easy to turn the chop on such a frying pan. This contradiction can be resolved by space separation principles. For this purpose the board is made traditional (i.e., conical) on one side of the frying pan and toric on the other side, as in Figure 10.6. TRIZ FUNDAMENTALS 207 208 DFSS INNOVATION FOR MEDICAL DEVICES 76 Standard Solutions These are generic system modifications for the model developed using substance-field analysis. These solutions can be grouped into five major categories: 1. Improving the systems with no or little change: 13 standard solutions 2. Improving the system by changing the system: 23 standard solutions 3. System transitions: 6 standard solutions 4. Detection and measurement: 17 standard solutions 5. Strategies for simplification and improvements: 17 standard solutions For example, in recycling household wastes, we used to place all waste into a single trash container; then we had paper, metal/plastic, and other waste; today, many municipalities require separation of wastes into specialized categories" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003147_ijmee.38.2.5-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003147_ijmee.38.2.5-Figure1-1.png", "caption": "Fig. 1 Sample fast-return actuator \u2013 the students\u2019 fi rst woodworking project. Pulley at rear is for optional drive system by DC motor and pulley.", "texts": [ " In a later course (Manufacturing Processes), students develop computer numerical control (CNC) and welding skills. Students worked in teams of two for almost all projects. In the shop, this buddy-system arrangement helped ensure students were attentive to each other\u2019s safety; no signifi cant injuries occurred throughout the course. Woodworking projects: fast-return actuator and acoustic guitar As their introductory project to woodworking equipment, students constructed a simple fast-return actuator mechanism (Fig. 1). This actuator (an inversion of the slider-crank mechanism) was also analyzed in the concurrent Dynamics class. This project taught skills on the miter saw, table saw, drill press, sander and band-saw. The basic design was adapted and modifi ed from Levy [9]. Mechanical engineering students sometimes have pre-existing skills in woodworking; this project was designed to allow both basic and advanced versions, to provide challenge to all levels. This project typically took one lab period. For their second woodworking project, students designed and built simple acoustic guitars (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003731_tmag.2011.2151845-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003731_tmag.2011.2151845-Figure3-1.png", "caption": "Fig. 3. Comparison between two FSPMLMs. (a) Proposed FSPMLM. (b) Existing FSPMLM.", "texts": [ " The finite element analysis (FEA) results under normal and fault conditions are compared and validated by the ones from the experiment. Theconfigurationof theproposedFSPMLMisshowninFig.1. The existing FSPMLM presented in [4] is shown in Fig. 2. Manuscript received February 21, 2011; accepted April 29, 2011. Date of current version September 23, 2011. Corresponding author: H. Yu (e-mail: htyu@seu.edu.cn). Color versions of one or more of the figures in this paper are available online at http://ieeexplore.ieee.org. Digital Object Identifier 10.1109/TMAG.2011.2151845 Fig. 3 shows the differences between two FSPMLMs. The primary teeth of existing FSPMLM are uniformly arranged, which means that the values of tooth pitches of the primary are identical, . As shown in Fig. 3(b), the two alternate teeth of primary, for example primary teeth l and 3, do not exactly face the corresponding secondary teeth. This will result in a lower flux. Therefore, different values of tooth pitches are used in proposed FSPMLM, , to increase the flux value of coils. In addition, each E-shaped iron core is used to replace the two U-shaped iron core and one flux-barrier, thus reducing the manufacturing cost. This coupling E-shaped configuration can also minimize the cogging force. The armature flux linkage changes with magnetic reluctance at different positions" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000319_095440605x8478-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000319_095440605x8478-Figure6-1.png", "caption": "Fig. 6 Coordinate systems of the example model", "texts": [ " When the dimensions of the prismatic beam are a, b and c with respect to the x, y and z axes, the compliance matrix Ch is [3] Ch \u00bc a Ebc 0 0 0 0 4a3 Eb3c \u00fe a Gbc 0 0 0 0 4a3 Ebc3 \u00fe a Gbc 0 0 0 0 Caa 0 0 6a2 Ebc3 0 0 6a2 Eb3c 0 0 2 66666666666666664 0 0 0 6a2 Eb3c 6a2 Ebc3 0 0 0 12a Ebc3 0 0 12a Eb3c 3 77777777777775 (46) where Caa is given by Caa \u00bc a Gk2bc3 for b . c a Gk2b3c for c , b 8>< >: (47) and k2 is a geometrical constant determined by b and c. Themasses and the inertia moments for the equivalent matrix are listed in Table 1. The compliance matrix of a rectangular shell is obtained by equation (46). The stiffness matrix with respect to the flexure hinge coordinate is calculated from the inverse of the compliance matrix of the rectangular shell. The coordinate systems of the compliant mechanism are shown in Fig. 6. The Euler angles for the rotational transformation matrices are listed in Table 2. In addition, the vectors d and l are listed in Table 3. Let model A be the equation of motion proposed by this study, and model B be the conventional equation of motion presented in references [12] to [14]. Then, the elements of the equivalent stiffness are given by equations (40) and (41) for model A, and by equations (40) and (43) for model B. Because the modal frequencies calculated using equation (42) resulted in a discrepancy with some imaginary values, the diagonal equivalent stiffness of model B was used in equation (40) instead of equation (42)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure6-1.png", "caption": "Fig. 6. Datum surface is pitch cylinder.", "texts": [ " 5(a), a3 is parallel to unit vector a, so there exists q2 \u00bc AIa \u00f0N\u00f01\u00de v\u00f021\u00de p \u00de \u00bc 0, noninterference condition is satisfied naturally; for intersected-axis drive (A = 0), then there is q = IsinR(a \u00b7 P(0)). In Fig. 5(b), a3 \u00c6 a = a3 \u00c6 P(0) = 0, so a3 is parallel to a \u00b7 P(0), there exists q2 \u00bc I sin R\u00f0a P\u00f00\u00de\u00de \u00f0N\u00f01\u00de v\u00f021\u00de p \u00de \u00bc 0, noninterference condition is also satisfied naturally. Paralleled-axis drive is applied widely in the engineering, and we can acquire basic properties of this type of drive from general principles of normal circular-arc gear drive concluded above. In Fig. 6, R\u00f01\u00dep and R\u00f02\u00dep are a pair of pitch cylinders, IA is instantaneous axis (line of action), R1 and R2 are the radii of two pitch cylinders. Additionally, u1, u2 are the rotational angles of two gears. Under the condition of constant speed ratio, the following relationship can be obtained: A \u00bc R1 \u00fe R2 I \u00bc u2 u1 \u00bc R1 R2 \u00f024\u00de If let conjugate point P move along IA with some motion rule and R\u00f01\u00dep ;R\u00f02\u00dep rotate about z1-axis and z2-axis at the speed ratio I, respectively, and then the locus, formed by the point P on two pitch cylinders, are a pair of conjugate directrixes C\u00f01\u00dep ;C\u00f02\u00dep " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001586_s0022112078001755-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001586_s0022112078001755-Figure1-1.png", "caption": "FIGURE 1. Definition sketch of the co-ordinate system.", "texts": [ " It is supposed that t,he axis of the body always lies in the plane 2 = 0. This is made possible by assuming that the 2 components of the force acting on the fish balance at each cross-section, assumed symmetric with respect to the plane 2 = 0, and therefore the resultant of the forces has only X and Y components. Thus the Z components of vectors are ignored in the following analysis. The depth s of a cross-section is used as the size in the 2 direction. The angle between the S axis and the body axis is denoted by 8 (figure 1). It is convenient to define a tangential unit vector t , a normal unit vector n and a unit vector ex in the -3 direction by t = (cos8,sin8), n = (-sinO,cos8), e , = ( - l , O ) , (2. la, b , c ) where t points from the nose to the tail and turning t anticlockwise through 90\" yields n. Thus in terms of 5 (a Lagrangian co-ordinate along the body axis) the position X of a cross-section is written as X = ( X , Y ) = X , + < t ( - a < f [ < a ) , ( 2 . 2 ) the nose lying a t [ = - a and the posterior end at f [ = a", " A rough reading of Bainbridge's figures ,which give the distribution of the mean speed of lateral movement along the body length, shows that the fractional distances x, of the point of minimum amplitude from the snout relative to the length from the snout to the anterior edge of the caudal fin are 0.36 (dace), 0.31 (bream) and 0.29 (goldfish). In the present analysis the corresponding fractional length would be defined by Rep,, which is 0.42 for 6 = 1 or 0 with y* = 0.05, CT = 0.6 and v = 10. According to Fierstine & Walters (1968), for the wavyback skipjack (a scombroid fish) this point of minimum amplitude is somewhere near the base of the pectoral fin. A rough estimate of this position from their figure 1 shows that the value of x, is around 0-31. To minimize the yawing movements at the snout, it may be preferable to have a smaller value of xm as well as for the yawing amplitude itself to be minimized. The observed values of x,, which are smaller than that in the 550 T. Kambe present analysis in spite of the similar lateral body shapes, may be partly due to the rather asymmetric mass distribution (more mass in the anterior half of the body of scombrids; see Fierstine & Walters) and fin arrangement" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003648_icems.2011.6073664-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003648_icems.2011.6073664-Figure1-1.png", "caption": "Fig. 1 Cross section of the Dual-Star FSPM motors and their flux distribution with one phase excited with constant current.", "texts": [ ". INTRODUCT\u0130ON HANKS to their high torque density and their simple rotor structure, which is similar to that of the Switched Reluctance Motors (SRMs), the Flux-Switching Permanent Magnet (FSPM) motors attract more and more attentions in the critical application as the Hybrid Vehicles and aerospace [1][2], in which the reliability of system is essential. In order to enhance the fault tolerant capability of electrical machines, the redundant structure can be applied [3] [4] [5]. As shown in the Fig. 1 (I), both the machines have one primary star and one redundant star, which are excited independently as shown in the Fig. 4. In the double layer FSPM motor, there are two layers of adjacent phases in the same stator slot. The primary star constituted by the phases A1, B1 and C1 and the redundant star constituted by the phases A2, B2 and C2. This is the same for the single layer FSPM motor, while the only difference is that there is only one layer of the same phase in each stator slot. Under normal mode, only the primary star is excited for the double layer FSPM motor as well as for the single layer FSPM motor", " ELECTROMAGNET\u0130C CHARACTER\u0130ST\u0130CS OF D\u0130FFERENT FSPM MOTORS Since the two FSPM motors have the similar stator as well as rotor and, the only difference is the winding geometry, the flux path of one phase excited of these two FSPM motors can then be establised by using one common cross section shown in the Fig. 2 (a). Based on the flux path, the equivalent lumped magnetic circuit can be obtained as shown in the Fig. 2 (b). \u0130t should be noted that for the double layer FSPM motor, the two stator teeth of the same phase are symetrical (see the Fig. 1 (II) (a)) and the turn number of the equivalent lumped magnetic circuit should be one half of the total phase. Furthermore, the obained inductance by the Fig. 2 (b) should be multipled by two times to obtain the phase inductances of the double layer FSPM motor. T The general expression decribing the magnetic reluctance in the Fig. 2 (b) could be established as: 1 (1) where and are respectively the relative permeability and the permeability of free space, l and s are respectively the length and the area of the cross section of the flux path", " The relationship between the reluctance (R), the inductance (L) and the phase turn number (N) (for the two machines, the phase turn numbers are the same) could be established as (3) and the expression of the phase inductance of these two FSPM motors could be finally obtained as follows: (3) 2 2\u2044 2 (4) where and are respectively the phase self inductances of the double layer and the single layer FSPM motors. It is found that the self inductance of the double layer machine ( ) is practically twice as low as that of the machine with simple layer ( ). On the contrary, for the mutual inductance between phases of both FSPM motors, it can be noted that FSPM motor with single layer has lower mutual inductances than those of FSPM motor with double layer because as in Fig. 1, there are obviously more mutual fluxes between phases in FSPM motor with double layer than those in FSPM motor with single layer [6] [7] [8]. The self and mutual inductances are also calculated by Finite Element Method (FEM) 2D, the results of self and mutual inductances against rotor positions are illustrated in Fig. 3. Only self and mutual inductances of the phase A1 are chosen, the other phases (phase B ant phase C) are similar to the phase A while with a phase separation of one-third electrical cycle" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002090_9780470264003-Figure12.3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002090_9780470264003-Figure12.3-1.png", "caption": "Figure 12.3 Boothroyd\u2013Dewhurst Inc. (BDI) software.", "texts": [ " The DFMA team will then continue to evaluate design changes as the final device concept moves into the optimize phase. The team will also issue an updated report prior to the characterize tollgate review. A DFMA specialist will be a valuable addition in the characterize phase. The specialist will help the team compare the metrics generated during DFMA analysis to the targets and provide feedback to the team. The specialist will also begin to track the concept\u2019s device cost using a commercially available software tool such as the tool depicted in Figure 12.3. Such concurrent costing software can estimate the cost of individual components with 23 different shape-forming processes. Before starting DFMA analysis on the device design concept under consideration in the characterize phase, the DFSS team should consider a list of items as potential inputs to the analysis, such as manufacturing quality data on an existing baseline device, current manufacturability issues, field reliability issues (field failures), serviceability issues, customer requirements in the QFD, product requirements definition, competitive benchmark analysis, supplier management strategy, lifetime and yearly volume estimates, device cost targets, and previous prototypes results, if applicable" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001100_bf01228535-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001100_bf01228535-Figure5-1.png", "caption": "Fig. 5. The curvilinear coordinates (~, r/). ~ = arc length around a unit circle with centre at the force centre (0, - 1). r/= radial distance ratio f rom this unit circle.", "texts": [ " The trajectory is formed by the repetition of a trajectory section at regular angular intervals (1-N,/Np)2~, which is exact; the accuracy of the approximation is the same for each section, there being no secular accumulation of error, and this solution is neither restricted to small relative displacements nor to a few revolutions of the station. 6. The Trajectories of Low Ejection Speed as Prolate Cycloids The geometric nature of this new first-order solution becomes apparent under a transformation to the curvilinear coordinates (~, r/) shown in Figure 5. The coordinate ~ is a measure of arc length around the circle rp-- 1, and the coordinate r/measures distance ratio from that circle, in the radial direction. The origin of this coordinate system is at the station, and the relationship to the polar coordinates (rp, 7) is 7/7 6 \" = - ~ -t ; r / = r p - 1. (36) 2 A NEW FIRST-ORDER SOLUTION FOR THE RELATIVE TRAJECTORIES OF A PROBE 85 The first-order solution in this curvilinear coordinate system is then = (N~/Np - 1)Ep - (Ns /Np + 1)ep sin Ep + ~ - ~*, (37) r / = (ap - 1) - apep cos Ep" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000669_s095679250400573x-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000669_s095679250400573x-Figure1-1.png", "caption": "Figure 1. Solutions a are given by the intersection of the solid curves y = sin(2ah) and y = 2(ah)/(A+h), and these are stable if the corresponding value of cos(2ah) (the dotted curve) is negative. Thus, the first (trivial) root is unstable, the second stable (provided it lies in ah > \u03c0/4), the third unstable, and so on, nontrivial solutions alternating between stable and unstable as a increases for as long as roots exist.", "texts": [ " Since we impose strong anchoring \u03b8 = 0 at z = 0, to leading order the solution for the director angle \u03b8 is \u03b8 = a(x, y, t)z. (3.12) Thus, ah represents the total angle turned through by the director across the sample. To satisfy the free surface condition (3.9) a(x, y, t) satisfies a(x, y, t) = A+ 2 sin [ 2a(x, y, t)h(x, y, t) ] , (3.13) where the dimensionless anchoring strength A+ = \u03b4LA+/K . Nontrivial solutions of (3.13) exist only if A+h > 1 everywhere; if this is the case then there may be multiple solutions of (3.13) (see Figure 1). Solutions a may be found in terms of h by considering the intersections of the curves y = 2ah/(A+h) and y = sin 2ah. In the case that multiple solutions exist, not all will be stable. By considering the sign of the second variation in (3.6), given by (\u2206J)2 = \u03b52 2 \u222b\u222b\u222b \u2126 \u03b72 z dV \u2212 \u03b52A+ 2 \u222b\u222b \u2202\u2126+ \u03b72 cos 2ah dS, (3.14) it may be deduced that roots for which cos(2ah) < 0 are stable (small variations in the solution only increase the energy), while those for which cos(2ah) > 0 are unstable. Hence, considering Figure 1, only if A+h > \u03c0/2 everywhere will stable solutions for a exist. The stable configuration observed in experiments will correspond to the smallest (positive, without loss of generality) nontrivial root of (3.13), as this is the lowest energy state (recall that ah represents the angle the director turns through across the film). Thus, when nontrivial solutions exist, the observed solution will be such that \u03c0 4h < a < \u03c0 2h . (3.15) The free elastic energy W is then given by W = \u03b82 z 2 = a(x, y, t)2 2 ", " (Table I) [6] yield \u03b11 = \u22120.2191, \u03b12 = \u22121.3365, \u03b13 = \u22120.0133, \u03b15 = 0.9431 \u03b16 = \u22120.4068 for the normalised viscosities. In this case the functions Fi(\u03bb) are all positive. They are sketched in figure 2, for \u03c0/2 < \u03bb < \u03c0, which, by (3.15), is the range of interest. Thus the Ii are all positive for MBBA. Note that the film thickness h cannot go to zero within the framework of this model, since for sufficiently small thicknesses h no stable energy minima will exist (the stable roots a all disappear; see Figure 1). Hence we are restricted to considering films sufficiently thick that stable energy minima exist. 3.2.2 Gravity parallel to the film When gravity in the (\u2212x)-direction is assumed to be the driving mechanism, which is equivalent to a flow driven by a constant influx supplied at x = +\u221e, (3.19) is replaced by pz = 0, and (3.17) is modified with an extra gravity term (\u2212B1) on the right-hand side, where B1 = B/\u03b4. An analogous procedure then leads to \u2202h \u2202t + \u2202 \u2202x [( C ( \u22072h ) x \u2212 B1 \u2212 Naax ) I1 \u2212 NhaaxI2 ] + \u2202 \u2202y [( C ( \u22072h ) y \u2212 Naay ) I3 \u2212 NhaayI4 ] = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003132_jmes_jour_1969_011_008_02-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003132_jmes_jour_1969_011_008_02-Figure9-1.png", "caption": "Fig. 9", "texts": [ " 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 bands may coalesce, the asymptotic solutions, equations (13), (14) and (15), being applicable. The stability map is shown in Fig. 8, where the form of the mode coupled instabilities is also indicated. It is seen that the instability bands accociated with the higher modes are negligibly thin. This is to be expected as these modes (+uz and +2z in Fig. 7) consist principally of base vibrations and therefore the asymmetry of the shaft has little effect. UniformIy asymmetric shaft, with asymmetric 'perfect' boundary conditions The shaft of Fig. 9 has a generating system consisting of pin-pin shaft modes in the Oxy plane, and free-pin shaft at DEAKIN UNIV LIBRARY on August 12, 2015jms.sagepub.comDownloaded from PARAMETRICALLY EXCITED LATERAL VIBRATIONS OF AN ASYMMETRIC SLENDER SHAFT 65 modes in the 0x2 plane. In evaluating the coefficients v, p the notation and numerical data from (12) can be utilized directly. Equations (5) then yield vyil = C I W ~ , ~ ~ with vy,, = 0 (i # j ) vzIl = c(wzizy with vat, = 0 (i # j ) It is convenient in this case to recast the equations (4) in terms of the dimensionless time variable 7 = SZt where EZ 522 = - pAL4 the shaft speed is now denoted by the ratio w n = n With the modal shapes and inertia coefficients given in (12) it is then found that 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 VoZII No I I969 at DEAKIN UNIV LIBRARY on August 12, 2015jms" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001171_j.commatsci.2006.03.017-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001171_j.commatsci.2006.03.017-Figure3-1.png", "caption": "Fig. 3. Sketch for three-dimensional conical heat source models.", "texts": [ " Thus, surface distribution energy input models of Gauss and double ellipse, commonly employed in GTAW and other welding processes, are unsuitable for VPPAW process; the three-dimensional high-density energy input model, however, is preferred [18,45]. Some researches regarding the 3-D high-density heat input model have been conducted [18,45\u201348] and it is shown that 3-D energy distribution models of cone and double ellipsoid can accurately compute temperature fields of high-density energy input welding processes. Although gouging and puddling of plasma arc are coupled in both 3-D double ellipsoidal heat source model and 3-D conical heat source model (see Fig. 3 [18]), the latter one is preferred to simulate the keyhole-mode plasma arc welding processes (e.g. DPAW or VPPAW). Wang and Wu et al. modify the 3-D conical energy distribution model (see Fig. 3) to properly describe the keyhole-mode PAW processes, and the modified conical heat source model is given as [18] qarc\u00f0x; y; z; t\u00de \u00bc gIU 0 2pr2 0\u00f0z\u00de exp _r2 2r2 0\u00f0z\u00de \u00f013\u00de where _r \u00bc \u00f0x2 \u00fe y2\u00de1=2 \u00bc \u00bd\u00f0x x0\u00de2 \u00fe \u00f0y y0 vt\u00de2 1=2 r0\u00f0z\u00de \u00bc re\u00bdln\u00f0z\u00de ln\u00f0zi\u00de \u00fe ri\u00bdln\u00f0ze\u00de ln\u00f0z\u00de ln\u00f0ze\u00de ln\u00f0zi\u00de where x0 and y0 are the location parameters for energy distribution (the values of x0 and y0 used in the study are 0 and 10, respectively), ze and zi the z-component coordinates for the top and bottom workpiece surfaces, respectively, re and ri the radii of Gaussian energy distribution for the top and bottom workpiece surfaces, respectively (these parameters can be obtained through experiments and researchers\u2019 experience)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003140_iros.2009.5354270-Figure12-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003140_iros.2009.5354270-Figure12-1.png", "caption": "Fig. 12. Robotic action patterns.", "texts": [ " Experimental Condition and Method We attempted to avoid the stagnant state at the bottom of the horizontal bar by applying various types of rewards. In this experiment, the rewards based on the robotic physical quantities\u2014decrease in swing height, swing angle, tip angle, and mechanical energy\u2014as listed in Table II were given to the robotic actions in the Q-Learning process. The tip angle t\u03b8 was calculated by means of the kinematic information of the robot, as shown in Fig. 11. In addition to this condition, the movable ranges of enabled motors were constrained as shown in Fig. 12 in order to imitate those of human beings; we expected that the robot may acquire a human-like motion. Similar to the experiment in chapter IV, the Q-Learning with \u03b5-greedy method was applied. Then, \u03b5 was reduced with the time transition at the rate of 2.0 \u00d7 10\u22126 per learning step. First, we executed the Q-Learning for each reward by using the ODE-based dynamic simulator of the giant-swing robot in order to reduce the learning time and to avoid the fatigue breakdown. Subsequently, we attempted to pick up the effective learning results and actually implemented them in the real robot so as to examine the performance and applicability of the learning results obtained in the simulation" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003361_j.robot.2010.10.007-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003361_j.robot.2010.10.007-Figure2-1.png", "caption": "Fig. 2. Footstep position in the double support phase (a). Positions reachable from the current position of the robot by a single step according to the constraints (3.13), (3.14) and (3.16) (b).", "texts": [ " A step is feasible if there are no collisions between the robot and its environment during the execution of the step, and if in the double support phase at the beginning and the end of the step both feet of the robot rest on a flat horizontal surface. These two conditions are evaluated by approximating the robot by a hull containing all points of the robot during the execution of a particular step. The shape of this hull is such that it allows efficient collision checking as well as efficient checking whether the entire surface of the sole of each robot\u2019s foot is in contact with a flat horizontal surface in the robot\u2019s environment. Fig. 2(a) shows the robot footstep position during a double support phase. The parameterization of the footstep position used in this article is based on the straight line segment with endpoints R and R\u2032\u2032, referred to in Fig. 2(a) as a path segment. The SSP approach proposed in Section 5 results in a path representing a chain of such segments. The robot reference frame SR is centered at the point R with z-axis antiparallel to the gravity axis and x-axis parallel to the path segment. The standing foot is assigned the reference frame SF and the swinging foot the reference frame SF \u2032\u2032 . The x-axis of each of these reference frames defines the orientation of the corresponding foot. The robot footstep position is defined by vector c = pT , z, \u03b1, \u03be, 1l, 1z, 1\u03b1st , 1\u03b1sw T (3", " This can be formulated by introducing a function fnext as follows c\u2032 = fnext(w,w\u2032, \u03be , 1\u03b1\u2032st). The definition of the function fnext is given in Appendix C. Because of the constraint (3.3), themaximumchange ofwalking direction is 1\u03b1max, which can be formally written as\u03b1\u2032 \u2212 \u03b1 \u2264 1\u03b1max. (3.12) From (3.5) and (3.9) it follows thatz \u2032 \u2212 z \u2264 1zmax, (3.13) which defines the range of z \u2032-component of the posew\u2032. From (3.2) and (3.8) it follows thatp\u2032 \u2212 p \u2264 1lmax. (3.14) Since the x-axis of SR is parallel to the current path segment, cf. Fig. 2(a), the following holdsp\u2032 \u2212 p > 0\u21d2 \u03b1\u2032 = \u0338 (p\u2032 \u2212 p), (3.15) where symbol \u0338 u denotes the angle of a vector u = [xu, yu]T , i.e. \u0338 u = a tan 2(yu, xu). From (3.12) and (3.15) it follows thatp\u2032 \u2212 p > 0\u21d2 \u0338 p\u2032 \u2212 p \u2212 \u03b1 \u2264 1\u03b1max. (3.16) Constraints (3.13), (3.14) and (3.16) are illustrated in Fig. 2(b). These constraints, together with the constraints (3.4), (3.6) and (3.12), define the set of all posesw\u2032 that could possibly be reached from c by a single step, assuming an appropriate adjustment of the standing foot orientation. This set is denoted in the following by Wnext(c). An exact definition of setWnext(c) is given in Appendix C. Using set Wnext(c) and function fnext, set Cnext(c) \u2282 C can be formally defined as the set of all configurations c\u2032 = fnext(w,w\u2032, \u03be , 1\u03b1\u2032st), where w\u2032 \u2208 Wnext(c) and 1\u03b1\u2032st satisfies constraints (3", "28 m Foot width wF 0.12 m Distance between the centers of the robot\u2019s feet when the robot stands with its feet closed lfd 0.26 m Maximum change in walking direction in one step 1\u03b1max 45\u00b0 Height of the component V (1) of the hull shown in Fig. 4 h(1) 0.35 m Maximum step clearance h(2) 0.175 m Radius of the smallest cylinder inside which the robot can turn in place rrot 0.31 m Radius of the circle which completely contains the sole of a robot\u2019s foot rf 0.18 m Distance from the foot to the safety margin L (cf. Fig. 2) \u03b5 0.01 m complexity of the scenario as well as on the robot\u2019s initial and goal position in a given scenario. The scenario #3 proved to be themost difficult for the proposed SSP method. For this scenario, a feasible step sequence was found in less than 0.3 s on average. However, the maximum execution time exceeded 1.2 s in one particular situation. After one of the obstacles was moved by only 0.01 m, an appropriate step sequencewas found in 770ms for the same initial position. Results of the slightly modified scenario are presented in Table 4 as scenario #3\u2217", " Although the assumptions about the locomotion abilities of the robot used as the basis for this work are not too restrictive, their validity should be evaluated by experiments with a real biped robot. The authors would like to thank the anonymous reviewers of this paper for their invaluable remarks and suggestions. In this appendix, the derivation of the foot collision constraint (3.6) is given. Assuming that (3.3) and (3.4) are satisfied where 1\u03b1max \u2264 \u03c0/4, the feet are not in collision with one another if the point V of the standing foot shown in Fig. 2(a) and the swinging foot are at the opposite sides of the straight line L. The distance \u03b5 of line L from the swinging foot defines a safety margin, i.e. the minimal allowed distance between the feet. Let FpV be the position of point V relative to the standing foot reference frame SF . It is defined by FpV = \u2212 lF 2 sgn(\u03be \u2212 0.5) wF 2 . (A.1) Position RpV of point V relative to the robot coordinate system SR is defined by RpV = R(1\u03b1st) FpV \u2212 0 sgn(\u03be \u2212 0.5) lfd 2 , (A.2) where R is a 2D rotation matrix, defined by R(\u03d5) = [ cos\u03d5 \u2212 sin\u03d5 sin\u03d5 cos\u03d5 ] " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000513_2004-01-1058-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000513_2004-01-1058-Figure4-1.png", "caption": "Figure 4: Electromagnetically Controlled Pilot Clutch Type TMD", "texts": [ " Torque Management Devices can be manually operated (dog clutch) or passively or actively controlled. Passively controlled Torque Management Systems are either torque sensing (helical gear type, multi-plate type etc) or speed sensing like the commonly known viscous coupling [1]. This paper mainly concentrates on actively controlled Torque Management Devices. Therefore these devices will be explained in more detail. Several systems available are electromagnetically controlled by a pilot clutch. One example is shown in the Figure 4. In this system a small clutch pack - the pilot clutch - can be activated by an electromagnet. The magnetic force pushes the clutch plates in the pilot clutch pack together, which then activates a ball and ramp mechanism once a speed difference occurs between input and output. This ball and ramp mechanism acts on the main clutch pack thus providing the required locking torque. Other systems in the market are activated hydraulically. Typically these systems contain a pump, which creates hydraulic pressure, under speed difference, that acts on a conventional wet clutch pack" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003011_gt2010-22632-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003011_gt2010-22632-Figure4-1.png", "caption": "Figure 4. DOMAIN AND POST\u2013PROCESSING SURFACES", "texts": [ " This boundary condition fixes the mass flux across the boundary and allows the total pressure to change. Due to computational time constraints, this was not possible to do for all gears. The outlet boundary for all simulations is a pressure outlet condi- tion, in which static pressure is prescribed as zero Pascals, once again replicating the conditions in which the experiment was run. Four surfaces are used for post-processing of static pressure and mass-flow-rates for analysis, and these are all shown in Figure 4 . The perspective for this view is marked on Figure 3, and shows a section of the domain through a periodic boundary. Nose Restriction is axially located half way along the restriction at the \u201cnose\u201d to the shroud. Shroud Outlet is at the exit to the shroud, where the flow passing between gear and shroud passes into the back chamber. These surfaces are annular in shape for a whole gear. Gear Inlet is effectively a conical surface which intersects the domain at a point just after the entrance to the gear teeth and is inclined at the back cone angle (90\u25e6 minus the pitch cone angle)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000402_icpp.2005.9-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000402_icpp.2005.9-Figure8-1.png", "caption": "Figure 8. Samples of routing.", "texts": [ " If the source s can know all the surfaces [xd : xd, 0 : yd, 0 : zd], [0 : xd, yd : yd, 0 : zd], and [0 : xd, 0 : yd, zd : zd] can be reached, based on the sufficient and necessary condition for the existence of minimal path in Theorem 2, a minimal routing is feasible from d to s (i.e., a minimal routing from s to d). Figure 7 shows some samples of this feasibility check. After the feasibility check, the routing process considering three preferred directions: the +X , +Y , and +Z is similar to that in 2-D meshes which only considers +X and +Y directions. Figure 8 shows some samples of routing under our MCC model in 3-D meshes. In summary, the contributions of this paper are listed as the following: (a) We have rewritten the MCC model in 2-D meshes without using global information so that the shape information at our boundaries can be used not only to ensure the existence of a minimal path, but also to form a minimal routing by routing decisions at intermediate nodes along the path. This routing will find a minimal path from source s to destination d whenever it exists" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.82-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.82-1.png", "caption": "Fig. 14.82 Single-person hanging scaffold", "texts": [ " Hanging scaffolds can be divided into the following kinds: \u2022 Stationary, hand-driven, single-person (cradle) scaffolds\u2022 Stationary, hand- or electrically driven scaffolds\u2022 Mobile, hand- or electrically driven scaffolds\u2022 Stationary, hand- or electrically driven sectional scaffolds Stationary, single-person hanging scaffolds are cradles with a single-person workstation. The cradle is suspended by a hoisting cable from a boom placed on the roof of a building. The vertical motion of the scaffold is effected by means of a two-crank hand winch operated by the person in the cradle. There are containers for materials and tools on both sides of the cradle. A single-person hanging scaffold is shown in Fig. 14.82. Specifications of single-person hanging scaffolds are listed in Table 14.12. Stationary Hanging Scaffolds These are scaffolds in which the work station is a platform with winches, suspended on cables. Hand-driven and electrically driven hanging scaffolds are used. A stationary hanging scaffold with a hand drive is shown in Fig. 14.83. The hanging scaffold shown in Fig. 14.83 consists of three main units: a platform, two winches with a hoisting cable, and a boom. The platform is a steel frame (lined with boards) with a take-down guardrail" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003961_detc2013-12361-Figure11-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003961_detc2013-12361-Figure11-1.png", "caption": "FIGURE 11. Sun Gear: Multiple Tooth Spall Case", "texts": [ " For all test cases and test conditions, data were acquired at 50 kHz sampling rate with a 25 kHz aliasing filter cut-off frequency. The data were acquired for 40 sec. per set. In total, there are two sets of baseline cases, two sets of single spall cases, one test for the multiple spall, and one test for the cracked sun gear. The single spall case is shown in Figure 10. This spall covers around 75% of the toothface. No other ap- preciable damage was noticed via visual inspection on any of the other teeth. Another component tested was the sun gear shown in Figure 11. This gear had four teeth with severe damage. On ST10, a spall about one-third the facewidth exists. On tooth STID12, there is a chip at the tip which extends about 20% of the facewidth. ST14 has spall covering about 80% of the facewidth and, on ST15, almost the full facewidth is spalled. A simulated crack was machined using Electrical Discharge Machining (EDM) and is shown in Figure 12. The notch was placed in the tooth fillet region along the complete width of the tooth. It had a crack depth of about 25 percent of the total tooth cross section length and a circular path similar to that which would naturally occur [39]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.112-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.112-1.png", "caption": "Fig. 14.112 Plastering unit working in tandem with equipment for transporting dry mortar", "texts": [ " Liquefied mortar is pumped through a hose from the vessel to the plastering Vibrator for aiding flow of cement in silos Valve closing flow of plaster from silo to vessel Pressure vessel Additional jet of air aiding high bulk density mortar feeding Aeration fabric for liquefying mortar in vessel and transporting it Control box Compressor Compressor drive\u2019s electric motor Fig. 14.111 Equipment for transporting dry mortar from silo to plastering unit unit. If mortar with a high bulk density is to be pumped, the valve should be opened in order to aid the flow of mortar with an additional jet of air. Portable equipment for transporting dry mortar, working in tandem with a plastering unit, is shown in Fig. 14.112. Besides a vibrator, an aeration unit is introduced in order to ensure proper flow of cement from the silo to the reservoir. The cement in the reservoir is liquefied by means of blow-in nozzles. The plastering unit\u2019s mortar reservoir is equipped with a lid with a fill-up signalling gauge and an air filter. Dry mortar feeding systems can be made as mobile (equipped with a driving axle) or portable. Mixers for Mortars and Plaster Mixes Mixers for mortars and plaster mixes form a highly diverse class of plastering machines in terms of their size and principle of operation", " The following kinds of machines are distinguished: \u2022 Electrically driven, hand-operated mixers\u2022 Continuous-type mixers\u2022 Batch-type mixers The machines are used for preparing masonry mortars, plasters, self-leveling mixtures, and so on. Electrically driven, hand-operated mixers work by one or two electrically driven agitators that are manually introduced into a container to mix the contents. Continuous-type mixers can be equipped with an open hopper (Fig. 14.113) or a dry mortar bin working with a dry mortar transport system (Fig. 14.112). A diagram of a continuous-type mixer is shown in Fig. 14.113. In this mixer dry components are fed into the hopper and transferred by raking-out paddles and the feeding screw into the mixing unit, where water is added. The rate of flow of water into the mixer is controlled by a valve and a flowmeter. Batch-type mixers perform the same function as continuous-type mixers, except that their operation is periodical and consists of the charging of components, their mixing, and discharging in succession" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002349_s12239-009-0039-8-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002349_s12239-009-0039-8-Figure3-1.png", "caption": "Figure 3. Speed ratio changing.", "texts": [ " The circular contact area improves the torque capacity and reduces the transmission size. It is also more effective in supporting the extremely high stresses on the contact point, which is an inescapable problem of the traction drive CVT. Further, the spin loss is minimal in the circular contact area at every speed ratio, while many of other traction drive CVTs have an elliptical contact area, which causes large spin. The counter rotor can be tilted to change the radii of rotation of each of the rotors, which determine the speed ratio, as illustrated in Figure 3. The cam and follower mechanism, known as the ratio changer, controls the tilt angle. There is a tilt handle under each of the counter rotors, which is guided by the spiral groove on the drumshaped cam. As this drum-cam rotates, the counter rotors are synchronously inclined. The pressure to produce the traction force on the contact point is generated by the coil spring and the loading cam. The coil spring is installed in the counter rotor assembly with the designed preloading, which improves the efficiency of the initial operation" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002034_pesc.2007.4342050-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002034_pesc.2007.4342050-Figure2-1.png", "caption": "Figure 2. Estimated and ideal rotor flux reference frames According to (3), the injected signal creates an electromagnetic torque oscillation", "texts": [ " Low-frequency Signal Injection Method The LF signal injection method presented in [8] is implemented on SMPM. In this paper, it\u2019s applied for IPMSM. Inserting (2) into (1), the steady-state equation can be written as sq s sq r d sd qu R i L i e (7) where q r me is the back EMF on q axis. In real system, if there is an rotor position error , the LF current signal ( ) ( )2 cos c c c i t I t , which is superimposed on the estimated d axis, could create harmonic components cdi and cqi on the ideal d- and q-axis respectively, as is showed in Fig. 2. ( ) 1.5 ec m cq T t P i (8) From (4) and (6), it can be seen that, this torque oscillation can cause a rotor speed response and finally a q-axis back EMF response. After some mathematical deductions as is done in [8], the resulting response of back EMF on the estimated q-axis can be expressed by (8). 2 23( ) sin( ) 2 m c cq c c P I e t t J (9) The exact rotor position can be obtained by adjusting the error angle to zero, which also leads to 0cqe . However, can\u2019t be obtained easily. As a result, an error signal F is constructed" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003361_j.robot.2010.10.007-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003361_j.robot.2010.10.007-Figure4-1.png", "caption": "Fig. 4. Hull V (c) and cylindrical solids V (i).", "texts": [ " hulls representing the robot walking straight ahead, 2. hulls representing the robot turning in place and 3. hulls representing the robot stepping over/onto an obstacle. In this paper, we propose a modification of the robot model presented in [7,8]. Instead of modeling an action such as straight ahead walking, turning in place and stepping over an obstacle, the approach considered herein is to assign a hull V (c) to each configuration c. The top and a side view of a hull V (c), where c = pT , z, \u03b1, \u03be, 1l, 1z, 1\u03b1st , 1\u03b1sw T , are shown in Fig. 4 for \u03be = 0. The hull is uniquely defined by the components of configuration c. It consists of 4 cylindrical solids, V (1) approximating the shape of the robot\u2019s upper body and upper parts of its legs, V (2) approximating the shape of lower parts of the robot\u2019s legs and V (3) and V (4) which represent the lowest parts of the robot\u2019s legs including its feet. These cylindrical solids are defined by r (1) = rrot, z(1) = z + 1z(1), B(1) = p1p2, r (2) = rf , z(2) = z + 1z(2), B(2) = p3p4 \u222a p5p6 \u222a \u222a \u222a , r (3) = rf , z(3) = z, B(3) = p8, r (4) = rf , z(4) = z + 1z, B(4) = p7, where pipj denotes the straight line segment connecting two points pi and pj and denotes the arc of a circle with the center in pi connecting two points pj and pk lying on the circumference of the circle", " The base of the hull V (c) is the union of the bases of its components V (3) and V (4). For the SSP approach proposed in this paper to be applicable to a particular robot, the robot must have the following property. Property 1. The mechanical abilities of the robot allow it to achieve every configuration c \u2208 C by standing in a stable posture such that the robot is completely contained inside hull V (c), while the soles of the robot\u2019s feet are completely contained inside the base of this hull. The obstacle shown in Fig. 4 demonstrates that stepping over an object and changing the walking level are treated in a unified manner. The gaps between two walking surfaces are handled analogously. A limitation of the robot model presented in this section is that themaximumheight andwidth of an obstaclewhich the robot can step over represent constant parameters, while, in reality, the maximum surmountable height of an object may depend on itswidth. Another limitation is that the proposedmodel is designed for motion planning using a 2", " Variations of the free space detection time tFS and the time tNNF needed for computation of NNF for different initial and goal positions are rather small. On the other hand, SSP time tSSP significantly depends on the Table 2 Robot model parameters. Parameter description Symbol Value Maximum step length 1lmax 0.50 m Maximum step height 1zmax 0.175 m Foot length lF 0.28 m Foot width wF 0.12 m Distance between the centers of the robot\u2019s feet when the robot stands with its feet closed lfd 0.26 m Maximum change in walking direction in one step 1\u03b1max 45\u00b0 Height of the component V (1) of the hull shown in Fig. 4 h(1) 0.35 m Maximum step clearance h(2) 0.175 m Radius of the smallest cylinder inside which the robot can turn in place rrot 0.31 m Radius of the circle which completely contains the sole of a robot\u2019s foot rf 0.18 m Distance from the foot to the safety margin L (cf. Fig. 2) \u03b5 0.01 m complexity of the scenario as well as on the robot\u2019s initial and goal position in a given scenario. The scenario #3 proved to be themost difficult for the proposed SSP method. For this scenario, a feasible step sequence was found in less than 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002224_0020-7403(72)90101-4-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002224_0020-7403(72)90101-4-Figure2-1.png", "caption": "FIG. 2. Schematic drawing of deformed envelope.", "texts": [ " The envelope is assumed to be of in6nite extent in the x s direction; thus the deformation of one sheet is described by y~ = x~+u~(x 1, t), (2.1) where u~ are the Cartesian components of the displacement vector. A similar form holds for the other sheet. The deformed state of such an envelope is schematically illustrated in Fig. 3. Here it is noted that since the envelope, which cormi~ts of two sheets, is resting on a base which is aesmned rigid that, in general, there will be a fiat portion whose length is denoted by 2L in Fig. 2. The long fluid storage bag: A contact problem for a closed membrane 433 The components of the deformation tensor in the co-ordinate frame x~ are given by (cf. Green and Zerna a) . , , du l . ldux\\ s . ldus\\ s ldus\\ s o , 1 = (2.2) du s O x s = ~ Css-- 1, Cxs= dus ' ~xx' Css=O, Css= 1. Since the two sets of cords make equal angles ( T- ~) with the z s axis, the inextensibility condition may be stated as follows Clxsinsa+OsscosSc~+Oxssin2c~ = 1, t (2.3) Cn sin s a + C88 cos s ~ - Cxs sin 2~ 1. ! Using equation (2" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001920_1.2988480-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001920_1.2988480-Figure5-1.png", "caption": "Fig. 5 Measurement of the tooth surface", "texts": [ " We can make the rigin O and gear axis zg in the coordinate system O-xgygzg atached to the gear coincide with the origin Om and axis zm in the oordinate system Om-xmymzm of the CMM, respectively. Howver, the angle by which the gear is rotated about its axis is un- ertain. Therefore, we must define an unknown angle between ournal of Mechanical Design om: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/17/201 the xg and xm axes as shown in Fig. 4. Therefore, the tooth surface of the straight bevel gear and its unit normal are rewritten as xm and nm in Om-xmymzm as follows: xm u, ; = C xg u, + 0,0, T nm ; = C ng 6 where is the error of the apex to back l. Figure 5 shows that the gear tooth surface xm and a spherical probe of radius r0 of the CMM are in contact with each other at point Q. The coordinates of the probe center P are theoretically expressed as a position vector P Px , Py , Pz in Om-xmymzm as follows: P u, ; = xm u, ; + r0nm ; 7 On the other hand, the coordinates of the probe center P are measured by the CMM. The measured coordinates are expressed as a position vector M Mx ,My ,Mz . If is known and the real tooth surface is manufactured without errors, P should be equal to M as follows: P \u2212 M = 0 8 In practice, however, assuming that the coordinates of n points on the real tooth surface are measured, each residual Ei i =1,2 , " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002858_s12239-009-0081-6-Figure14-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002858_s12239-009-0081-6-Figure14-1.png", "caption": "Figure 14. Normalized percentages of the parameters effects on the transmitted vibration for (a) vertical harmonic force and (b) harmonic torque, at the mid-point of the driver\u2019s seat rail.", "texts": [ " Generally, interaction of A, B,\u2026, and K parameters may be obtained by employing the following equation (Montgomery, 2005): (1) If the effect includes the considered factor, the corresponding parenthesis appears with a minus sign (Table 2), and if it does not include the considered factor, it appears with a plus sign (Montgomey, 2005). AB...K( )Interaction= 1 2 k 1\u2013 -------- a 1\u00b1( ) b 1\u00b1( )... k 1\u00b1( )[ ] Once the interaction of the effects of the parameters has been computed, the sum of squares of interaction effects may be calculated as follows: (2) Using Equations (1) and (2), 15 estimated effects and sums of squares may be calculated. The results are shown in Tables 3 and 4. The normalized percentages of the parameters effects on the vibration transmission are illustrated in Figure 14 for both vertical force and moment inputs. As it may be seen from Figure 14, factor D, which denotes the stiffness of connection of the engine supporting member to the body, had a more pronounced effect (over 60% contribution) in the vibration transmission in the first input, whereas factor A had the dominant effect in the second input. However, the remarkable difference between the effects of parameters D and A in the first input and the small difference between these effects in the second input justifies the conclusion that a higher value of the parameter D may be chosen" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.130-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.130-1.png", "caption": "Fig. 14.130 Excavator with associated equipment for making deep point-excavations", "texts": [ " An example of a remote-controlled machine is the pull shovel shown in Fig. 14.129. It is radio-controlled and has an operating range (distance from the control panel) of 1500 m. The working system\u2019s cylinders and travel drive are controlled by levers on the control console. Two video cameras, mounted outside the excavator, transmit a picture of the working area. A third camera inside the operator\u2019s cabin shows the indications of the gauges. Another example of a remote-controlled excavator is the excavator shown in Fig. 14.130 [14.53]. The machine is intended for making deep excavations, vertical transport of output, and loading it into transport means. A slewing body with excavating tools and containers or negative-pressure conduits for transporting output are mounted on the crawler chassis. The excavation\u2019s minimum diameter is 3 m. The excavation rate and the output loading rate can be adjusted to the ground\u2019s properties. If soft rock is encountered, a milling drum can be mounted. The excavator is used in mountainous terrain not easily accessible to pile-drivers to avoid the hazards (landslides, subsoil waters, toxic gases, and falling objects) to which the operator would otherwise be exposed" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000958_j.aca.2004.10.063-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000958_j.aca.2004.10.063-Figure1-1.png", "caption": "Fig. 1. Schematic diagram of reflective optical microscope for the measurement of membrane thickness. The distance of focus difference between c and c\u2032 was equal to the thickness (\u03b4m) of immobilized membrane. The gray area indicates the membrane.", "texts": [ " Thickness estimation of hydrated membrane The hydrated protein membrane was like a soft gel so that the measurement of real thickness was hardly measured by contact tip mode (e.g., -step method). An indium-doped tin oxide (ITO) glass was chosen as the basal plane and the protein-immobilized substrate. The area of Dp\u2013GA or BSA\u2013GA membranes formed on ITO glass was restricted within a 3 mm diameter circle the same as the area of disk electrode. Principle of the measurement was based on the difference of the focal distance between the top of membrane and the basal plane, referred to the membrane thickness (\u03b4m). Fig. 1 elucidates the method to measure the membrane thickness. Points c and c\u2032 were respectively focused on ITO basal plane and the top of membrane at 400\u00d7 magnification, and then the focus difference between the two points was equal to \u03b4m. For more accurate estimation, the measuring pathways were selected on two orthogonal straight lines (baselines) across the center of the membrane. Thirty points chosen with a 0.2 mm interval between two points along the orthogonal lines were separately measured. Moreover, a and b points as shown in Fig. 1 locating outside the membrane were assigned to be the baseline of ITO plane. 2.6. Procedure i n o m t d i s F 2 a .3. Determination of actual surface area The electrode were first polished with sand papers of 00 mesh and then Al2O3 powders to obtain a clean and mooth surface. At first 1N H2SO4 solution was purged with igh-purity nitrogen for 30 min to remove oxygen prior to ach measurement according to the established procedure as reviously described elsewhere [15]. An Au electrode in N2aturated 1N H2SO4 was successively measured by the CV ith the scanning rate of 100 mV s\u22121 from \u22120", "6 V that exceeded the need of oxidation of ferrocyanide and the current was monitored as a function of time. Subsequently, the RDE method was employed with the rotation rate ranging from 400 to 2500 rpm to determine the steady-state current. The diffusion and partition coefficients were calculated from the current values measured by PS and RDE methods. 3. Results and discussion 3.1. Thickness of immobilized membranes The thickness of hydrated and dry immobilized membranes was measured by means of the focus-difference method of a reflection microscope. The variation in a\u2013b altitude (Fig. 1) was smaller than 2 m, which was much lower than the membrane thickness (described later). Furthermore, the influence of ITO surface roughness on the thickness calculation was negligible. Therefore, the a\u2013b baseline was standardized to calibrate the membrane thickness. Table 1 shows t m v f B g s o t c 1 g c f r t c e T s 3% Dp and 1.5% BSA solution. The amount of BSA or Dp molecules cross-linked should be the same when the added GA concentration was greater than 1%. It was the reason for dry Dp\u2013GA or BSA\u2013GA membranes with almost the same thickness whatever GA concentration altered" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003271_s00170-011-3475-3-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003271_s00170-011-3475-3-Figure4-1.png", "caption": "Fig. 4 Circular pitch and angles", "texts": [ "3 The tool path equations for the left tooth flank Similarly, the location on the left involute curve of the end mill was taken into consideration while parametric tool path expressions for the left tooth flank of the gear were derived (see Fig. 3) [22]. The parametric expressions of the tooth path for the left tooth flank of the gear from Fig. 3 were derived as follows: Xp0 l \u00bc rn:Cos 2invamax \u00fe q invan\u00f0 \u00de \u00fe Re:Sin an 8\u00f0 \u00de \u00f06\u00de Yp0 l \u00bc rn:Sin 2invamax \u00fe q invan\u00f0 \u00de \u00fe Re:Cos an 8\u00f0 \u00de \u00f07\u00de Zp0 l \u00bc b Constant\u00f0 \u00de \u00f08\u00de where, to calculate 8 in Eqs. 6 and 7, angles corresponding to a tooth thickness and width of space on the spur gear were taken into consideration (see Fig. 4). The angles 8 and \u03b8 were derived as 8 =2inv\u03b1max\u2212 invan and \u03b8=\u03b8p\u22122(invamax\u2212inv\u03b8p) from Figs. 3 and 4. In a related paper, rotation procedures to Eqs. 1, 2, 3, 6, 7 and 8 were applied to cut the whole of teeth. The following expressions were used for this operation. Xrmk Yrmk Zrmk 2 4 3 5 \u00bc Cosbk Sinbk 0 Sinbk Cosbk 0 0 0 1 2 4 3 5: Xpr Ypr Zpr 2 4 3 5 \u00f09\u00de Xlmk Ylmk Zlmk 2 4 3 5 \u00bc Cosbk Sinbk 0 Sinbk Cosbk 0 0 0 1 2 4 3 5: Xp0 l Yp0 l Zp0 l 2 4 3 5 \u00f010\u00de and k=1,......... ....,N. where, N and bk were described as the number of teeth of the gear and rotation angle, respectively", "1 Variations of the cutting errors In this paper, a programme in MATLAB is prepared by using mathematical expressions derived above to investigate the cutting error of the tooth profile curve according to design parameters of the gear (see Fig. 11). Parameters such as am, N, f and m are changed in this programme, and the obtained results are illustrated graphically (see Figs. 12, 13 and 14). In the paper, the graphics are drawn for variations of deviation quantities on the right involute curve, but it is not drawn for the left involute curve because the right and left involute curve on the tooth profile are symmetrical to each other (see Fig. 4), where interval 0\u2264an\u2264amax is described for angle an corresponding to each point on the involute curve. \u201cAlpham\u201d and \u201calphan\u201d in the figures are used instead of angle am and an, respectively (i.e. alpham=am and alphan=an). It is seen that the cutting errors on the tooth profile curve increase depending on increments of am in the figures above. Quantities of these errors further increase when an is equal to \u03b1max (see Figs. 12, 13 and 14). Because distance GH increases depending on the increment in the radius rG, ra and rH, quantity of the cutting errors increase (see Fig. 10). Besides, it is seen that the cutting errors also increase depending on increments in values N, f and m of the gear (see Figs. 13 and 14). For the increment in these values raises the dimensions such as rr, rb, rp, ro of the gear (see Fig. 4). These rise in the dimensions cause to increase interval GH, and this case increases the cutting errors of the tooth profile. As a result, it is seen that the increment or decrement in the length GH according to parameters such asN, f, m of the gear affects the cutting errors of the tooth profile. In the spur gears manufactured in a three-axis CNC milling machine according to the radial cutting method, it is seen that the cutting error in the tooth profile increases depending on increment of the parameters such as N, am and m of the gear" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001044_ac60366a003-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001044_ac60366a003-Figure1-1.png", "caption": "Figure 1. Schematic of flash photolysis instrument", "texts": [ " However, several major modifications ' Present address, Physics and Analysis Branch, Research and Development Department, Phillips Petroleum Co., Bartlesville, Okla. 74004. have been made to facilitate these measurements, and these are described below. Flash Radiation. A new optical arrangement has been adopted for this work. I t differs from that used previously ( 4 ) in that a hemispherical mirror has replaced the parabolic mirror; the flash lamp is inserted through the center of the mirror and mounted along the focal axis as shown in Figure 1. The mirror itself is 3 inches in diameter, is made of glass, and the inner reflecting surface is covered with a high-ultraviolet-reflectance coating (PhotoTechnical Research, Inc., Cambridge, Mass.). The double convex quartz lens (3-in. diameter, F = 3.5 in.) projects a diffuse spot of light, reflected into the photolysis cell by a flat mirror having the same high-UV-reflectance coating as the spherical mirror. An estimated 50-7506 increase in photon flux in the vicinity of the monitoring electrode has been obtained with the spherical optics relative to the previous linear parabolic mirror arrangement ( 4 ) " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002858_s12239-009-0081-6-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002858_s12239-009-0081-6-Figure2-1.png", "caption": "Figure 2. Engine supporting system (EM1 to EM3 are the engine mounts).", "texts": [ " Since natural frequencies of the engine block are much higher than those of the mounting system, it is reasonable to model the engine block with a lumped mass having the relevant inertia. This lumped mass element was connected to the engine-mount system by rigid elements (REB2 elements). To avoid artificial stress concentration around the RBE2 element, many related nodes were chosen to construct the connection. Engine mounts were modeled using solid elements. The FEA model of the engine-supporting system is shown in Figure 2. In Figure 2, the connection points of the transverse engine supporting system to the BIW (body in-white) are referred to as EM1, EM2, and EM3, respectively. To evaluate the transmitted accelerations, responses measured at the mid-point of the driver\u2019s seat rail were used in a harshness optimization objective function. Modal analysis was performed to determine the natural frequencies, mode shapes, and frequency constraints. It is evident that the peak values of the transmissibility curves must occur at frequencies that are far enough from the resonance frequencies of the vehicle subsystems (including the engine block, body, and suspension system components)", "5 g lateral and \u22121 g vertical accele- rations), and (9) 1st gear WOT (1 g longitudinal and \u22121 g vertical accelerations). Some design criteria were also used. For example, the ratio of the rubber mount stiffness to the metallic mount bracket must be less than 0.1 (Shariyat, 2006). In the sensitivity analysis, design parameters were denoted by specific symbols. These symbols as well as their upper and lower limits are listed below: \u2022 Factor A (supporting plates/under frame thickness): It may vary from 0.5 to 1.2 (mm). \u2022 Factor B (Elastic stiffness of the rubber engine mounts shown at the bottom of Figure 2): This factor may vary within \u00b150% of the initial value. \u2022 Factor C (Elastic stiffness of the rubber engine mount shown at the top of Figure 2): Variations within \u00b150% of the initial value are permitted. \u2022 Factor D (Stiffness of the connections of the engine sup- porting member to the BIW): Variations are considered to be within \u00b150% of the initial values. Factor D change may be achieved by altering the connection type or by using thin vibration isolators between the engine supporting structure and the body. The effect of each parameter in reducing the vibration transmissibility and amplitudes was determined, and a sensitivity analysis was performed to determine the influence of each parameter on the vibration suppression" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003722_lindi.2011.6031137-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003722_lindi.2011.6031137-Figure2-1.png", "caption": "Figure 2. The flux differential detector", "texts": [ " To avoid the harmful effects of faults on the systems in which the SRMs are used it is necessary to detect the faults already in their incipient phase [8]. \u2013 143 \u2013978-1-4577-1841-0/11/$26.00 \u00a92011 IEEE III. FAULT DETECTION The winding faults of the SRM can be sensed by several failure detectors [9]. In a first approach a simple fault detection device can be applied: the over-current detector given in Fig. 1 [10]. Its efficiency is limited due to insufficiently fast response time and the inability to detect all types of faults. Another simple detector, the flux differential one is shown in Fig. 2. It requires additional search coils wrapped around the stator poles. The search coils of each phase are connected in series opposing. Hence during normal operation the induced voltages of the search coils are equal and opposite, leaving a zero voltage at the terminals of the series pair. When a fault occurs the imbalance in the pole fluxes induces a voltage in one search coil that is greater than the voltage in the other coil, producing a voltage that can be detected with a bidirectional comparator", " The results were plotted versus the rotor's displacement (see Fig. 6). In position 0 pole A is perfectly aligned with a rotor pole. The decreased magnetic flux due to the shorted turns in coil A can be clearly observed. \u2013 145 \u2013 By having computed also the variation of the magnetic flux function of time, the emf induced in the search coils having 200 turns round pole A can be plotted (see Fig. 7). By computing the emf in both search coils placed on opposite poles (A and A') the voltage differences to be sensed by the detector given in Fig. 2 can be plotted for both machine conditions taken into study (see Fig. 8). As it can be seen from figure 8, the voltage difference in case of shorted winding turns conditions is enough significant to be sensed by a precise differential detector. V. DYNAMIC SIMULATION OF SRMS Dynamic simulations are one of the best choices to emphasize the effects of different winding faults on the SRM's behavior [14]. For this purpose the same SRM as in the previous case was simulated. The machine and its power converter must be simulated by means of coupling two software packages [15]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000585_j.intermet.2006.03.012-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000585_j.intermet.2006.03.012-Figure1-1.png", "caption": "Fig. 1. A schematic illustration of 3DMW freeforming process.", "texts": [ " In the present study, the influence of arc current on diameter, height, and contact angle with the substrate of intermetallic alloy beads in the TieNi and TieFe systems was investigated. Their microstructural changes associated with arc current were observed and discussed on the formation of their simple 3D objects. The idea of 3DMW consists of a combination of two different techniques; freeform fabrication, also called rapid prototyping, and TIG (Tungsten Inert Gas) welding. The principle of 3DMW is shown in Fig. 1. A metal substrate is placed on an xey stage under the tungsten electrode for arc discharge. When pulsed micro-arcs are emitted, the tip of a thin Ti wire fed by the rotation of two opposed capstans is fused and a micro-metal bead is formed instantaneously. A fused bead is welded to a metal substrate or previously formed beads. Then a Ni (or Fe) wire fed alternately from the opposed capstans is fused and a Ni (or Fe) bead is formed on this Ti bead resulting in the formation of a TieNi (or Fe) bead" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003123_robio.2009.4913293-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003123_robio.2009.4913293-Figure1-1.png", "caption": "Fig. 1. Reference Frame Definition Placing Cluster Center at Triangle Centroid", "texts": [ " This framework was successfully demonstrated for both holonomic and non-holonomic two-robot systems, including several clusterspace-based versions of regulated motion [8], automated trajectory control [9], [10], human-in-the-loop piloting [11], [12], and potential field-based obstacle avoidance [13]-[15]. The method was also implemented for three-robot [16] and four-robot [17] non-holonomic systems. 978-1-4244-2679-9/08/$25.00 \u00a92008 IEEE 1911 To further develop the application of the cluster space framework, we have applied it to the specification and control of three differential drive robots operating in a plane [16]. This section reviews the selection of cluster space variables, the derivation of the relevant kinematic transforms, and the formulation of an appropriate control architecture. Figure 1 depicts the relevant reference frames for the planar 3-robot problem. We have chosen to locate the cluster frame {C} at the cluster\u2019s centroid, oriented with Yc pointing toward Robot 1. Based on this, the nine robot space state variables (three robots with three DOF per robot) are mapped into nine cluster space variables for a nine DOF cluster. Given the parameters defined by Figure 1, the robot space pose vector is defined as: \u2212\u2192 R = (x1, y1, \u03b81, x2, y2, \u03b82, x3, y3, \u03b83) T (1) where (xi, yi, \u03b8i) T defines the position and orientation of robot i. The cluster space pose vector definition is given by: \u2212\u2192 C = (xc, yc, \u03b8c, \u03c61, \u03c62, \u03c63, p, q, \u03b2) T (2) where (xc, yc, \u03b8c) T is the cluster position and orientation, \u03c6i is the yaw orientation of rover i relative to the cluster, p and q are the distances from rover 1 to rover 2 and 3, respectively, and \u03b2 is the skew angle with vertex on rover 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.126-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.126-1.png", "caption": "Fig. 14.126 Disk sander for parquet floors", "texts": [ " The sanding drum\u2019s working width is up to 250 mm and is limited by the necessity for the drum to exert appropriate linear pressures (about 20 N/cm) on the surface. The design of a single-drum sander is shown in Fig. 14.125. A drum sander may weigh as much as 90 kg and the power of the driving motor can reach 3.5 kW at a drum rotational speed of about 2300 rpm. Electric motors are exclusively used to drive sanders because the latter are intended mainly for work indoors and the possibility of ignition of the wood dust collected in dust bags has to be eliminated. The described drum sander cannot sand the floor under heaters. For this purpose disk sanders (Fig. 14.126) are used. The working tool in this machine is an abrasive disk mounted at an angle of about 3\u25e6 relative to the base. Sandpaper is Velcro-fastened to the disk or clamped with a nut. Because of the disk\u2019s inclination, its working base\u2013contact surface amounts to about a third of the disk\u2019s surface area. Power is transmitted to the disk by a V-belt or a cogbelt. The driving motor\u2019s power does not exceed Part B 1 4 .7 2.0 kW at a rotational speed of about 4300 rpm. In currently manufactured disk sanders, disks about 180 mm in diameter are used" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000421_tmag.1983.1062848-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000421_tmag.1983.1062848-Figure1-1.png", "caption": "Fig. 1 :f ixed- 'and'moving parts of the t!leC%rOmagnetic", "texts": [ " The ana lys i s i s based on a non l i n e a r f i n i t e e l e - ment computation of the magnetic field combined with an a n a l y t i c a l or numerical evaluation of the eddy d i s t r i - bu t ion in the conduct ing s lab . The paper descr ibes the way i n which t h e eddy c u r r e n t d i s t r i b u t i o n l e a d s t o t h e de f in i t i on o f a power loss f a c t o r , which al lows the determination of the electromagnetic torque. P a r t i c u l a r a t t e n t i o n i s brought t o t h e numeric a l problem caused by the ve loc i ty vec tor which lead us to use a special a lgori thm cal led the \"upwind\" f i n i t e element method. DESCRIPTION OF THE SYSTEM (Fig.1) A variable number,of magnetic poles energized by brake d i r e c t windings a re f ixed in f ron t o f t h e wheel. The r e l a t i v e r o t a t i o n o f the pole and the wheel generates an eddy cu r ren t s d i s t r ibu t ion the va r i a t ion o f which produce a braking torque. The a n a i y t i c a l method i s a c lass ica l Four ie r ana lys i s o f t he cu r ren t d i s t r ibu t ion . The f i n i t e element method i s based on a magnetodynamic model including a ve loc i ty vec tor ", " To compute t h e r e s i s t i v i t y f a c t o r and only for this ,we assume : -the permeabili ty constant - the f lux densi ty B with periodic X-variation -vector B perpendicular to the ac t ive sur face of t h e r o t o r so (see Fig. 2 l e f t ) Bx(x ,y)=G B Z ( X , Y ) = ~ ( ~ ) if y -UO,+UO BY(X,Y)=O =O if Y +UO o r y -UO -the eddy-currents developed themselTres on t h e exte rna l sur face( th in depth pene t ra t ion) . We consider two mode l s . i n the f i r s t model t h e eddy-curr e n t s have a phys ica l d i s t r ibu t ion on t h e e n t i r e developed surface o f t h e r o t o r ( s u r f a c e AECDEF on Fig 1 ) In th i s ca se the lo s ses a re ca l cu la t ed wh i th c l a s s i ca l Four i e r ' s ana lys i s . In the second model eddy-currents a- , r e r e c t i l i - n e a r and f o r c a l c u l a t e t h e l o s s e s we consider t h e t o p p a r t of the sur face CD of t h e r o t o r ( F i g 2 r i g h t ) using also Fourier ' s analysis .We then obtain the re - s i s t i v i t y f a c t o r as t h e r a t i o of these two previous results and t ak ing in to accoun t t he r e s i s t ance of t h e two c i r c u i t s " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000661_s0301-679x(03)00033-1-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000661_s0301-679x(03)00033-1-Figure1-1.png", "caption": "Fig. 1. The structure of the overrunning clutch for the pulse-CVT.", "texts": [ " A non-Newton rheogical model was utilized to calculate the frictional force and traction coefficient. An inverse solution is pursued to solve the simultaneous equations. An example problem of an overrunning clutch with Pulse-CVT with an input power of 1500 W is analyzed. The scientific and technological persons have tested the dynamic loading and friction behaviors of overrunning clutch and providing a better testimony and identifying the correctness of various models to the research in this paper [4]. Fig. 1 schematically shows the structure of the overrunning clutch for the pulse-CVT. The clutch mainly contains of an outer rim, a few rollers, a planet wheel (inter rim), and springs. The outer rim is rigidly connected to the output rocker attached to the pulse-CVT, and the planet wheel is rigidly connected to the output shaft of the clutch. The spring retains the rollers in the wedged space between the planet wheel and the outer rim. When the rotor rotates in a given direction, it will convert the reciprocating motion of the rocker of the Nomenclature b Hertzian-half length, mm E1;E2 Young\u2019s modulus E0 equivalent Young\u2019 modulus, 2 E0 \u00bc 1 m21 E1 \u00fe 1 m22 E2f traction coefficient F total frictional force, N G dimensionless material constant, G \u00bc aE0 H, H0 dimensionless film thickness, dimensionless initial-minimum film thickness h, h0 minimum film thickness, lm initial-minimum film thickness, lm h1 film thickness, lm J mass moment of inertia of the output shaft (including out rim, rollers, planet, etc" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000235_bf02905937-Figure19-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000235_bf02905937-Figure19-1.png", "caption": "Figure 19. Membrane element", "texts": [ " At each increment of the calculation process, the equilibrium equations are concluded from the principles of virtual work: \u03b4wint = \u03b4wext (31) Whereby fj represents the component j of specific volumic load density and pj the component j of surface density load. So represents the surface where pj is applied. The principles of virtual work are then formulated as:\u222b Go ( Cijkl \u03b5kl + \u03c3oij ) \u03b4\u03b5ij dG o = \u222b Go fj\u03b4ujdG o + \u222b So pj\u03b4ujdS o (32) The above equation is nonlinear, its solution is carried out via an iterative algorithm within an incremental loading process. The physical movement of Go to G is generated by a series of small perturbations, see Bathe [7]. A triangular finite element was used for the fabric modeling, see Figure 19. The surface meshing is topologically similar to that used for the initial Go form-finding. In this case, the Green-Lagrange strain tensor is written as:\u23a7\u23aa\u23aa\u23a8\u23aa\u23aa\u23aa\u23a9 \u03b5xx = \u2202u \u2202x + 1 2 [( \u2202u \u2202x )2 + ( \u2202v \u2202x )2 + ( \u2202w \u2202x )2 ] \u03b5yy = \u2202v \u2202y + 1 2 [( \u2202u \u2202y )2 + ( \u2202v \u2202y )2 + ( \u2202w \u2202y )2 ] \u03b3xy = \u2202v \u2202x + \u2202u \u2202y + [ \u2202u \u2202x \u2202u \u2202y + \u2202v \u2202x \u2202v \u2202y + \u2202w \u2202x \u2202w \u2202y ] \u23ab\u23aa\u23aa\u23ac\u23aa\u23aa\u23aa\u23ad (33) The displacement vector at node i is written: \u2212\u2192u i = \u23a7\u23a8\u23a9 ui vi wi \u23ab\u23ac\u23ad (34) The displacement field interpolation on element k is formulated: \u2212\u2192u = N1 \u2212\u2192u 1 +N2 \u2212\u2192u 2 +N3 \u2212\u2192u 3 = { N1 N2 N3 }\u23a7\u23a8\u23a9 \u2212\u2192u 1\u2212\u2192u 2\u2212\u2192u 3 \u23ab\u23ac\u23ad = \u2212\u2192N t\u2212\u2192U k (35) with \u2212\u2192 N t = \u23a7\u23a8\u23a9 N1 = a1 + b1x+ c1y N2 = a2 + b2x+ c2y N3 = a3 + b3x+ c3y \u23ab\u23ac\u23ad (36) The strain field can be written under the following matrix form: \u03b5 = [ BL + 1 2 BNL(\u2212\u2192U k) ]\u2212\u2192 U k (37) where BL is a (3\u00d7 9) constant matrix, BNL is a (3\u00d7 9) matrix dependent on the \u2212\u2192U k field" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003132_jmes_jour_1969_011_008_02-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003132_jmes_jour_1969_011_008_02-Figure2-1.png", "caption": "Fig. 2. Unstable zones with close juxtaposition of two natural frequencies", "texts": [ " These results differ little from equations (13), (14) and (15) respectively even when there is considerable disparity amongst the supposed nearly equal coefficients, 7, pn, y*,, v,, v*, being average values. The relationship between the asymptotic solutions and those applying with well separated natural frequencies can best be illustrated by considering the case where the coefficients p, v can all be expressed as multiples of a common factor E . As E varies, a stability map is drawn against w. Fig. 2 indicates such a map in the vicinity of w1 and +(w,+w,,) . The asymptotic solutions are seen to apply only when E is sufficiently large for overlapping of the separate instability zones. (2) Shaft on foundations stiff relative to all modes of practical interest This would correspond to a flexible shaft mounted in comparatively very stiff pedestals. Considerable asymmetry of bearing stiffness is then to be expected (11); however, provided that the shaft is relatively much more flexible than the bearings Further, high bearing stiffness relative to the shaft would ensure sufficient similarity between " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000718_1.1829068-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000718_1.1829068-Figure4-1.png", "caption": "Fig. 4 Wobble-plate compressor2", "texts": [ " The wobble plate is kept from rotating by pins A and A8, which ride in slots in the housing. Thus one rotation of the input shaft 1 causes link 2 to wobble, the four-bar is actuated, and the ratchet is indexed. Although these two mechanisms are very interesting applications of wobbling motion, probably the most widespread implementation of such motion is found in piston-cylinder or slidercrank mechanisms which, like the mowing machine of Fig. 2, are of the rotational-to-linear motion conversion type. Wobble-plate compressors and engines have an extensive history @3#. Figure 4 shows one of many examples of wobbling slider-crank mechanisms given by Artobolevsky @1#. Note the similarity of the piston rods here to members 6 and 68 in Fig. 3. Wobble plate 2 is attached to connecting rods 4 and 48 at spherical joints B and B8. Since rollers b and b8 are guided in slots in the housing, constrain- 005 by ASME MARCH 2005, Vol. 127 \u00d5 269 shx?url=/data/journals/jmdedb/27802/ on 03/07/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F ing the wobble plate from rotating out of the plane of the paper, revolution of the input shaft 1 results in actuation of the pistons 5 and 58" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000057_s0069-8040(08)70029-3-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000057_s0069-8040(08)70029-3-Figure2-1.png", "caption": "Fig. 2. The rotating disc. (a) Flow functions; (b) schematic streamlines.", "texts": [ " rotating hemisphere, see Table 3. (a ) Rotating disc and rotating ring electrodes The problem of laminar fluid flow to a rotating disc has been amply discussed in the literature [7, 101. We may describe the velocity components as [17,18] (see Fig. 1) u6 = r w G ( y ) U, = - ( c J u ) \u201d ~ H(y) u, = r w F ( y ) where y is a dimensionless distance from the electrode surface with CJ the rotation speed in rad s-l . The variation of the flow functions F, G , and H with distance and resultant streamlines are shown in Fig. 2. The particular form of the velocity components satisfies the Navier - Stokes equation and the equation of continuity. Note that u, is independent of the radial coordinate. We are interested primarily in the convection pattern close to the electrode surface in order to calculate the flux of electrons. Following Levich [19] , we say uz,z-*Q - cz2 U & , O 0 (16) u , , , , ~ Crz where C = 0.510 w3l2 u - ~ \u2019 ~ . The convective-diffusion equation in cylindrical polar coordinates in References p p . 434-441 362 the steadv state is where we neglect radial diffusion as being negligible compared with radia convection" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002567_robot.2007.364203-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002567_robot.2007.364203-Figure6-1.png", "caption": "Fig. 6. (a) The shaded regions represent tangential planes (b) The plane perpendicular to nf and containing the point q. The radius of the boundary of the cone is r = tan(\u03b8)(p \u00b7nf ). The angle \u03c6 equals to arccos(r/|AB|).", "texts": [ " Hence, every force in Pf \u2229 F generate torques lying not only in the same line but also in the same direction (sign). Since F is a convex set, all Pf yield torques being a convex set in Pp. Hence, T is a fan in this case. The boundary of the fan is the torque generated from two particular Pf that tangentially touch F . Formally,T is a fan H = {\u03b1a + \u03b2b|\u03b1, \u03b2 \u2265 0}, where a and b are the boundary vectors of the fan H . We can identify a and b by identifying the vectors fa and f b being on the boundary of F such that the plane containing the p and the vector tangentially touch the cone F (Fig. 6(a)). Let nf be the axis of F , let A be the point having p as its coordinate. We can calculate fa and f b by considering the plane \u03a0 perpendicular to nf and containing the point q. Let B be the point where the axis of F intersects \u03a0. Fig. 6(b) describes the calculation of fa and f b. A. R3-Positive Span of Three Torque Sets We have established that a torque set of any contact point is either a fan or a plane. Our problem is to check whether the three torque sets positively span R3. When any of the torque set is a plane, these sets can positively span R3 only when there exists two vectors that are on the different sides of that torque set plane. Let nt be the normal vector of the torque set plane. If the other sets are fans, we check whether the dot products of the boundary vectors with nt have different signs" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003077_asjc.196-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003077_asjc.196-Figure2-1.png", "caption": "Fig. 2. Missile divert control system.", "texts": [ " 1, the PWPF-type actuators try to output a signal that has almost the same frequency as the original input signal. Soon after the intercepting missile has been launched, the midcourse guidance [6] phase with thrust vector control (TVC) [6, 29] is initiated. When the missile approaches its pre-defined interception point at an altitude of around 150 km, the air density is low and the aerodynamic forces are difficult to exploit. Specifically, the divert control system (DCS) with thrust, as shown in Fig. 2, is assumed to be located near the missile\u2019s center of gravity and aligned with the two axes, bbx and bbz , perpendicular to the longitudinal axis, bby , of the kill vehicle, to generate motion. In contrast, the attitude control system (ACS) with thrusts, shown in Fig. 3, is located and aligned such that only three pure rotational moments about the principal axes are generated. III. ZERO-SLIDING GUIDANCE LAW A guidance system is generally designed to generate suitable motion commands for a missile to adjust its velocity and trajectory to fulfill the tactical goal" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002959_20090909-4-jp-2010.00076-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002959_20090909-4-jp-2010.00076-Figure3-1.png", "caption": "Fig. 3. Human model", "texts": [ " An excessive number of segments, however, increases the complexity of the resulting joint torque equations because of the complex coupling among all the segments; on the other hand a too small number of segments reduces the accuracy of the model. Thus, the seven segments, one for HAT (head, arm and trunk) and three for each leg (thigh, shank, and foot), is a good compromise between complexity and accuracy for the representation of the actual human body. Also, even thought the proposed model is planar on the sagittal plane, most of the gait motions can be explained with this model since the main walking motions happen in the sagittal plane. Fig. 3(a) shows the proposed human model, and Fig. 3(b) shows the notations and the directions of joint torques used in this paper. The global reference system (GRS) and the sign convention of the model are shown in the figure. The capital letters means the length of the part shown in the figure. A foot is modeled as a triangular shape to represent the actual foot shape. As stated in Section 2.1, the ground contact conditions are changed according to gait phases. The different ground contact conditions result in different constraints of the human model" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002171_j.triboint.2009.05.009-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002171_j.triboint.2009.05.009-Figure2-1.png", "caption": "Fig. 2. Geometry and reference systems. Left: illustration of the main", "texts": [ " The analysis of the results is focused on the behaviour of the journal orbits, maximum fluid film pressure and minimum fluid film thickness. In this section the formulation of representative equations describing the mathematical model for a hermetic compressor is developed. The motion equations for the piston-connecting rod\u2013crank system are formulated with the help of multibody dynamics theory, using Newton\u2013Euler\u2019s method and using a frame notation of common use in multibody dynamics [12\u201314]. Fig. 2 shows a sketch indicating the reference frames and the main angles of rotation (b, G, y and a). In order to be able to describe all the representative vectors, one inertial reference frame \u00f0IXYZ\u00de and four moving reference frames \u00f0Bi\u00de have been defined. The inertial reference frame is attached to the centre of the bearing (point O), the moving reference frames B1\u00f0X1Y1Z1\u00de, B2\u00f0X2Y2Z2\u00de and B3\u00f0X3Y3Z3\u00de are attached to the crank and the moving reference frame B4\u00f0X4Y4Z4\u00de is attached to the connecting rod. With the help of the geometrical transformation matrices Tb, TG, Ty and Ta any vector can be easily transformed from one reference frame to another. According to Fig. 2, the main constraint equation of the system is given by Ixp\u00feIl\u00bcIr\u00feIc (1) where Ixp \u00bc fxB;0; hpg T ; Il \u00bc f lcr cosa; lcr sina;0gT ; Ir \u00bc TT b T T C TT y B3 r; B3 r \u00bc frc ;0; hpg T ; and Ic \u00bc fxc ; yc ;0g T . The equations of motion for each body are given by Eqs. (2)\u2013(6). The force equations are described in the inertial reference frame and the moment equations of the crank and the connecting rod are described in the moving reference frames B3 and B4, respectively. Force equation\u2014crank:X If \u00bc mc I ac ) IfA \u00fe Ifub \u00fe Ifb \u00bc mcf\u20acxc ; \u20acyc;0g T (2) where Ifub is the vector of the crank unbalance force and Ifb is the vector of the dynamic journal bearing forces" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003917_19346182.2013.854799-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003917_19346182.2013.854799-Figure4-1.png", "caption": "Figure 4. Club head path differential between backswing and downswing in (a) sagittal plane and (b) frontal plane. Coaching swing plane is the two-dimensional projection of the shaft in the sagittal plane of the golfer (a); also shown in the frontal view (b).", "texts": [ " Kinematics and kinetics variables were calculated with respect to their local joint coordinate systems based on Cardan angle calculations: flexion\u2013extension, adduction\u2013 abduction, axial rotation and lateral bending were defined as angular rotations about each segment\u2019s local x-, y- and z-axes, respectively (Figure 3a and3b). There were four basic end-effector kinematics variables analysed in the backswing and downswing phases: (1) Club head path differential: distance between the club head position in downswing and the club head position at the same corresponding position in the backswing (Figure 4a and 4b). (2) Coaching swing plane: projection angle on sagittal plane of vector running through the golf shaft (Figure 4a); also shown in the frontal view (Figure 4b). (3) Dynamic swing plane: vertical elevation of golf shaft angular velocity vector with respect to the right horizontal projected on the sagittal plane of the golfer, the plane taken as orthogonal to the direction of the target line (i.e. viewed from behind the golfer) (Figure 5a). (4) Dynamic swing plane offset: horizontal offset angle of dynamic swing plane with respect to the target line (Figure 5b). Segmental sequencing was observed by examining the time histories of segmental angular velocities Figure 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003968_j.precisioneng.2012.05.005-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003968_j.precisioneng.2012.05.005-Figure1-1.png", "caption": "Fig. 1. Dimensions (mm) of specimens (a) bush and (b) pin.", "texts": [ " [14] developed a new approach to design nterference fits based on selective method assembly and FEM, to chieve more reliable interference fits between ring gear and wheel ssembly. In view of the above, in this work an effort has been made to tudy the effect of cylindricity error and surface modification along ith the interference on load bearing ability of interference fitted ssemblies. . Experimental details .1. Preparation of specimens To conduct experimental tests, the size of the pin and the bush re so chosen that it is easy to handle, assembling and testing [1]. he specimens were prepared using EN 8 steel as per the details ndicated in Fig. 1(a) and (b). The pins were turned using lathe and hen ground on a cylindrical grinding machine. The bushes were repared by drilling and reaming followed by internal grinding peration. During manufacturing of cylindrical components there s a chance of producing waviness/humps or nubs and other surface rregularities on the cylindrical surface. For any machining proesses, surfaces are produced which possess directional feed marks. enerally the feed marks are periodic in nature. To study the influnce of such deviations on load bearing ability of interference fitted ssemblies, the pins were prepared with nubs are as follows: (i) pin with no nubs (Fig. 1(a) and (b)); (ii) pin with one nub is at a distance of 6 mm from one end as shown in Fig. 2(a); iii) pin with two nubs, one at each end with a phase difference of 180\u25e6 (Fig. 2(b)) and iv) pin with three nubs, one each at ends and third one is at the centre with a phase difference of 120\u25e6 as shown in Fig. 2(c). one at each end with a phase difference of 180 and (c) pin with three nubs, one each at ends and third one is at the centre with a phase difference of 120\u25e6 . A large number of specimens of each category (i)\u2013(iv) as said above were prepared to facilitate selective assembly", " / Precision Further, the average undulation number [15] which provides a uantitative basis for specifying the Openness and Closeness of the oundness profile was computed using the Fourier coefficients as = \u221a\u221a\u221a\u221a\u221a\u221a\u221a\u221a\u221a m\u2211 k=1 k2[(ak) 2 + (bk) 2] m\u2211 k=1 [(ak) 2 + (bk) 2] (8) here m represents the cut-off point of the harmonics from the vailable harmonics. A computer program was developed to comute all the values of the above said parameters. .2. Analysis based on FEM The stress distribution at the contact/mating surface of pin and ush is analyzed by FEM and the model considered for the analysis s symmetric in respect of geometry and loading. The geometry of he pin and bush is shown in Fig. 1. The stress distribution at the ating surface of pin and bush is of main concern and the analyis was carried out on the assumption that the chosen material is omogeneous and isotropic in nature using ANSYS finite element nalysis package. The pin and bush were modeled using Quad 8 oded solid 82 elements. In the present analysis pin is considered as ontact surface and bush as a target surface and contac48 element s used to analyze the stress distribution at the interface of pin and ush. Fig. 7(a) and (b) shows the arrangement of nodes/elements f a finite element model used in the analysis of stress" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003271_s00170-011-3475-3-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003271_s00170-011-3475-3-Figure1-1.png", "caption": "Fig. 1 Radial cutting of the spur gear in three-axis CNC milling machine", "texts": [ " These are forming and generating methods, but in the previous study of the same author, the spur gears were machined by a method described as radial cutting method using the end mills in three-axis CNC milling machine differently from these cutting methods [22]. In the present study, to investigate the cutting errors of the tooth profile curves of the spur gears machined according to radial cutting method, a related method is briefly presented in the following sections. 2.1 Radial cutting method In the previous article [22], the gear blank was clamped by bolted mounting to a CNC milling machine as shown in the Fig. 1a. The spur gear was cut by the end mill in the parallel planes to XY-plane along the Z-axis of the gear blank. The tool paths were composed along the radial direction and circumference of the gear blank (see Fig. 1b). The gear material was chosen from a polyamide material because of easy machining. The dimensions of this gear blank were calculated by taking into consideration the IIx6DIN 3972 standard given for rack cutters according to ISO/R 53. The diameter of the end mill was calculated by considering the point wideness of the rack cutter. By machining the tooth flank of the gear by the CAM programme, the tool path expressions were derived by considering the generation principles of the involute curves for the right and left tooth flank" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001980_sisy.2008.4664900-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001980_sisy.2008.4664900-Figure8-1.png", "caption": "Fig. 8. Sketch of the robot with feet walking on its front feet edges (a and b) and on its side feet edges (c and d)", "texts": [ " It has to be emphasized that the notions of dynamic balance and ZMP are not connected with the robot's ability to perform walk. It is possible to synthesize robot's walk which is not dynamically balanced. However, such walk is little if at all resistant to disturbances (external and internal), and if a disturbance occurred the walk will mainly not be sustainable without corrective actions. Let us consider now the case of a robot with feet which are not flat on the ground surface, but rather rotated about the front edge (Fig. 8a) contacting the ground in a tiptoe manner or on the side feet edges (Figs. 8 c and d). In all the cases the contact is realized via both legs, hence we are speaking about the double-support phase. The support area is obtained by connecting with a straight line the front and hind edges of the ground and feet (Fig. 8b and Fig. 9). Besides, we have to clearly define here whether a static or dynamic case is in question. A static case corresponds to the maintaining of the immobile posture and dynamic one 5 Passive walkers move solely thanks to the moment generated by the body weight with respect to the ground so that they can move only on the slant surfaces. However, to this class we also include the robots having actuators and which can move on a horizontal support, like the Cornell robot [15], as they utilize the same principle as passive walkers", " It should be noticed again that the conclusions drawn in the above discussion hold in a static case all the time the posture considered is preserved, and if the case is dynamic they hold only for the time instant considered, since dynamic situation changes in time6. If the contact is made via the foot edge, it is clear that there is no support area in the single-support phase and neither there is ZMP, and hence the system is not dynamically balanced. However, in this case too, the system's motion is possible in the way described above. Moreover, the examples shown in Figs. 8 c and d are very similar to skate-shoe walking or roller walking, and the example shown in Fig. 8a is similar to tiptoe walk. All these sorts of walk require special skills, such motion is possible but, one has to bear in mind all the constraints imposed by such motions. During the single-support phase the system is not dynamically stable, whereas in the double-support phase it may be dynamically balanced, depending on whether the contact involving both feet is line-like or point-like. From this arises a general question: Why artificially constrain a robot to walk with its feet always flat on the ground when it is simply not necessary" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002211_j.humov.2008.11.001-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002211_j.humov.2008.11.001-Figure1-1.png", "caption": "Fig. 1. xC = hor height", "texts": [ " The effect on catching performance was assessed by systematically varying several variables, like the initial position of the catcher with respect to the landing spot of the ball, air friction, the take-off velocity of the ball and the catcher\u2019s sensitivity for detecting optical acceleration. The model was used to reproduce the running profiles that were observed experimentally by McLeod and Dienes (1996). Two coordinate systems were defined: a global coordinate system attached to the earth and a local coordinate system attached to the catcher\u2019s optical field that was set (arbitrarily) 1 m in front of the catcher\u2019s eyes (see Fig. 1). The equations of motion for a ball with air friction are \u20acxB;G \u00bc kB _xB;G ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi _x2 B;G \u00fe _y2 B;G q mB \u00f01\u00de \u20acyB;G \u00bc kB _yB;G ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi _x2 B;G \u00fe _y2 B;G q mB g mB ; \u00f02\u00de using the overdot notation to indicate differentiation with respect to time. The mass of the ball \u00f0mB\u00de was chosen to be 0.15 kg, equalling the official mass of a baseball. The air friction coefficient kB of a baseball was estimated to be 5", " xB;L \u00bc xB;G xc \u00f03a\u00de _xB;L \u00bc _xB;G _xc \u00f03b\u00de \u20acxB;L \u00bc \u20acxB;G \u20acxc: \u00f03c\u00de The position of the catcher at the beginning of each simulation was defined to be zero. The distance between the catcher and the launching position of the ball, calculated prior to the actual simulation, was the distance covered by the ball plus the distance to be covered by the catcher in each trial. This distance between the initial position of the catcher and the landing position will be referred to as \u2018initial distance\u2019. The projected height of the ball in the optical field (yf , Fig. 1) was expressed in local coordinates yf \u00bc yB;L xB;L : \u00f04\u00de The acceleration of the ball in the optical field was obtained by differentiating yf twice with respect to time _yf \u00bc _yB;L xB;L yB;L _xB;L x2 B;L \u00f05\u00de \u20acyf \u00bc \u20acyB;L x2 B;L 2 _yB;L _xB;L xB;L \u00fe 2 yB;L _x2 B;L yB;L \u20acxB;L xB;L x3 B;L : \u00f06\u00de First, the perceived optical acceleration (\u20acyfp) was made dependent on the optical acceleration (\u20acyf ) using a sigmoid function. Essentially, this equation transforms the optical acceleration into negative (\u20acyfp = 1), zero (\u20acyfp =0) or positive (\u20acyfp =1) perceived optical acceleration: For \u20acyf 6 0 \u20acyfp \u00bc y0 \u00fe 1 1\u00fe e \u20acyf \u00fethresh c " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002025_tmag.2007.915840-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002025_tmag.2007.915840-Figure4-1.png", "caption": "Fig. 4. Mesh of the variable reluctance machine studied.", "texts": [ " In this case, there are not numerical errors due the movement. However, in this case, it is well known that the locked-step introduces a very important constraint and limits the applications. This item essentially is devoted to the comparison of the methods from the numerical standpoint. A variable reluctance machine has been chosen because there are some difficulties to determine the derivative global values such as the electromotive force and the electromagnetic torque. The mesh (4623 elements) of the solid part is presented on Fig. 4. The air-gap mesh is regular with two layers to ensure accurate solutions. In our simulation to obtain a cogging torque the machine is fed with particular conditions. A direct current is imposed in one of the three coils and the machine rotates with a constant speed . The flux, their times derivatives that we call electromotive force (EMF) and the torque have been calculated. In Fig. 5, we can see the EMF waveform in function of time obtained by using MEM and PIM until order three, to simulate the movement" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003667_tmag.2010.2044875-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003667_tmag.2010.2044875-Figure2-1.png", "caption": "Fig. 2. Closed loop within a PMSM flux density distribution including an \u201cill-looping part,\u201d together with the equipotential lines of the corresponding magnetic vector potential.", "texts": [ " The error control by (7) is, in contrast to (5), independent from the displacement in curve direction, caused by the variable step size . The described algorithm should be used by direct user interactions in virtual reality to select the starting point interactively, as well as part of a seeding strategy algorithm; cf., Section III. In some cases, due to discretization errors, the field line never comes back in the vicinity of the seeding point. Usually, this happens at the singularities of the field; see Fig. 2. A further loop observation is then required to detect whether a seeding point is a valid point to find a closed loop. Such a situation is exemplified in Fig. 2, where the flux line algorithm is applied to a 2-D flux density distribution of a permanent magnet synchronous machine (PMSM) starting at the origin\u2014for comparison, the equipotential lines of the magnetic vector potential are given as well. To detect and remove this so-called \u201cill-looping part\u201d [3], the starting point in (5)\u2013(7) is substituted to its neighbor points and additionally tested on a closed loop. FE models require a huge amount of memory and are, especially in case of transient simulations, very time consuming" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000602_s1526-6125(05)70092-x-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000602_s1526-6125(05)70092-x-Figure6-1.png", "caption": "Figure 6 Thermocouple Measurement at Point of Interest", "texts": [ " This is because the characteristic conduction time, t = l2 / (where l is the distance from the center of the probe to the point of temperature measurement and is the thermal diffusivity of the material), for aluminum is much less than for tool steel of the sonotrode and anvil. Hence, heat dissipation to the surrounding aluminum mass is much faster than that into the tooling. Experimental Validation To calibrate the numerical thermal model, the experiments were conducted to record the temperature at a point of measurement located 1 mm anterior to the sonotrode edge in the direction of vibration and on the top side of the foil (Figure 6). First, a K-type thermocouple was used to obtain the temperature transient over the welding time. The temperature observed at the point of measurement (at t = 0.2 seconds) was then used to calibrate the contact thermal resistance factor, R, in the thermal model such that its predicted temperature matches to that of the thermocouple. The contact resistance, Rth, curve at the measurement point was, therefore, obtained between 0.1 seconds (trigger point) and 0.5 seconds (end of welding) and plotted in Figure 7", " Within a fraction of a second (at approximately 0.16 seconds), the temperature reaches its peak value of 45.1\u00b0C in the top foil and 44\u00b0C in the bottom foil. It is noted that this peak temperature rise is significantly less than the approximately 30% of melting temperature of the material (around 646\u00b0C in the case of Al 1100) reported in most literature (Hazlett and Ambekar 1970). We believe that this apparent discrepancy is primarily due to the distant location of the point of measurement from the center of the welding joint. As seen in Figure 6, the thermocouple tip was ultrasonically welded to the foil at the point of measurement\u2014the closest possible distance from the edge of sonotrode for successful measurement using a thermocouple. The experimentally obtained optimal distance of 1 mm kept the thermocouple bond intact\u2014a necessary condition for temperature measurement\u2014during the weld cycle. Other approaches of measuring the temperature directly at the welding interface under the sonotrode were tried but failed to give consistent results", " Also, the top foil has a smaller thermal capacity and thus responds with a higher temperature rise under the action of higher heat generation. In the plastic regime, the von Mises stresses actually rise slightly due to a minimal strain hardening of the material, but at the same time plastic strains decrease sharply. This causes less plastic work done, yielding less heat generation. The temperature appears to be gradually falling at around 0.35 seconds until it reaches asymptotically a uniform equalization room temperature (25\u00b0C) at the steady state. It should be noted in Figure 6 that the contact resistance at the welding interface drops steeply in the beginning until the foils yield plastically and then gradually over the rest of the welding duration. The other material properties of the aluminum foil (elastic modulus, yield point, thermal conductivity, etc.) are not significantly altered by the modest temperature variations and therefore are not adjusted (that is, they are assumed temperature independent) in the ultrasonic welding simulation. A numerical model for predicting transient temperature profile during one cycle of ultrasonic metal welding in a URM process was developed, calibrated, and verified experimentally within margin of errors as explained in the section above" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001202_iros.2006.281992-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001202_iros.2006.281992-Figure6-1.png", "caption": "Fig. 6 Two Robots Passing a Corridor", "texts": [ " These constrains are: -Vmax Vr Vmax (20) -amax 0 H ( X ) = 0 XE a P ( X ) = 0 H ( X ) > 0 X E R ' ( 3 ) 191 A second c o n d i t i o n t h a t has t o be s a t i s f i e d i s t h e f o r c e b a l a n c e e q u a t i o n : b ( X ) d x = f n ( 4 ) The i n t e r a c t i o n of t h e s e e q u a t i o n s i s i d e n t i c a l t o t h e l u b r i c a t e d c o n t a c t case , see f o r i n s t a n c e (11). I n t h i s d r y c o n t a c t a n a l y s i s t h e f i l m t h i c k n e s s i s set t o zero over t h e domain (a ) and t h e n o n l y t h e p r e s s u r e d i s t r i b u t i o n has t o b e c a l c u l a t e d . The problem h a s t r a d i t i o n a l l y b e e n s o l v e d b y d i r ec t m e t h o d s (Newton Raphson) , r e s u l t i n g i n l o n g computing t i m e s f o r l a r g e problems. AS a l a r g e number o f s o l u t i o n s i s r e q u i r e d i n a p p l y i n g t h e l i f e model t h e c o m p u t i n g t i m e c a n become e x c e s s i v e l y long . To a l l e v i a t e t h i s problem a n a l t e r n a t i v e i t e r a t i v e t y p e s o l u t i o n was used which h a s been d e s c r i b e d i n (12) . Convergence w a s f u r t h e r a c c e l e r a t e d by t h e a p p l i c a t i o n of novel m u l t i - l e v e l t e c h n i q u e s which have been a p p l i e d p r e v i o u s l y t o t h e s o l u t i o n o f d i f f e r e n t i a l e q u a t i o n s . The computing t i m e f o r t h e u s u a l i t e r a t i v e s o l u t i o n t e n d s t o be dominated by t h e i n t e g r a l c o m p u t a t i o n , s o t h e s o l u t i o n t i m e i s p r o p o r t i o n a l t o o r d e r n2 where n i s t h e number of p o i n t s , a s can b e seen i n Table 1. However it i s p o s s i b l e t o reduce t h e comput ing , t ime t o o r d e r n logn by t h e a p p l i c a t i o n of m u l t i l e v e l t e c h n i q u e s , s p e c i f i c a l l y M u l t i l e v e l M u l t i - I n t e g r a t i o n ( M L M I ) . The b a s i c i d e a was g i v e n i n (13) and worked out i n d e t a i l i n ( 1 2 ) . A s can be seen from Table 1, column 3 s i g n i f i c a n t t i m e s a v i n g s c a n be o b t a i n e d from l e v e l 6 onwards, s o t h e approach i s most u s e f u l f o r p r o b l e m s w i t h many g r i d p o i n t s . The c a l c u l a t i o n of t h e s u b s u r f a c e stresses i s a t a s k s i m i l a r t o t h e f i l m t h i c k n e s s s o l u t i o n , when one c o n s i d e r s o n l y one p a r t i c u l a r v a l u e of t h e d e p t h z, a t a t i m e . P l o t s of t h e d r y c o n t a c t s u r f a c e p r e s s u r e a n d a s s o c i a t e d s u b s u r f a c e o r t h o g o n a l s h e a r s t r e s s d i s t r i b u t i o n are shown i n f i g u r e 5 . 4 L I F E PREDICTIONS I n t h e t r a d i t i o n a l Lundberg a n d Pa lmgren b e a r i n g f a t i g u e l i f e model, t h e p r o b a b i l i t y of f a i l u r e c a n be e x p r e s s e d i n t e r m s o f t h e stressed volume and t h e magnitude and d e p t h below t h e s u r f a c e of t h e maximum o r t h o g o n a l s h e a r stress. I n ( 2 ) I o a n n i d e s and Har r i s p r o p o s e d a g e n e r a l i s e d model i n which t h e s t r e s s e d volume i s d i v i d e d i n t o d i s c r e t e volume e lements i n which t h e maximum stress i s c a l c u l a t e d a c c o r d i n g t o some stress r e l a t e d f a t i g u e c r i t e r i o n . I n common w i t h s t r u c t u r a l f a t i g u e l i f e p r e d i c t i o n s f o r s teels i n r e v e r s e d b e n d i n g or t o r s i o n , a t h r e s h o l d stress v a l u e i s d e f i n e d below which f a i l u r e w i l l n o t o c c u r . Each e l e m e n t i s w e i g h t e d a c c o r d i n g t o i t s depth below t h e s u r f a c e and t h e p r o b a b i l i t y o f f a i l u r e i s e x p r e s s e d i n terms o f t h e i n t e g r a l o f t h e e l e m e n t a l stresses over t h e e n t i r e volume. A modif ied l i f e c r i t e r i o n h a s been used t o compute t h e Ll0 b e a r i n g l i v e s i n t h e p r e s e n t c a s e . The maximum s h e a r stress ampl i tude z, i s c a l c u l a t e d d u r i n g t h e o v e r r o l l i n g of t h e d e n t . I n common w i t h s t r u c t u r a l f a t i g u e , t h e f a t i g u e stress t h r e s h o l d z, i s m o d i f i e d a c c o r d i n g t o t h e a b s o l u t e v a l u e of t h e s h e a r stress. z, i s assumed t o remain unchanged i f T,, d o e s n o t e x c e e d t h e y i e l d s t ress a n d t o d i m i n i s h l i n e a r l y t o z e r o f o r Tmax v a r y i n g between r e a n d t h e f r a c t u r e s t r e n g t h zf. A s t h e c rack might be e x p e c t e d t o b e c r e a t e d more e a s i l y i n t h e p r e s e n c e of a t e n s i l e r a t h e r t h a n compressive s t ress f i e l d , a n a d d i t i o n a l h y d r o s t a t i c weight ing was i n c l u d e d i n t h e model. I n t h i s t h e c r i t i c a l stress Ta was modi f ied to fa+ a.Hp, where Hp is e q u a l t o t h e h y d r o s t a t i c p r e s s u r e and a i s t a k e n as a=0.3.Using t h e s e v a l u e s , t h e p r o b a b i l i t y of s u r v i v a l of t h e i n n e r r i n g can be e x p r e s s e d a s : The e f f e c t i v e p e r t u r b a t i o n on t h e g l o b a l p r e s s u r e d i s t r i b u t i o n by t h e d e n t w i l l depend v e r y much on t h e r a t i o of t h e d e n t wid th t o t h e Her tz c o n t a c t s i z e . To a s s e s s t h i s e f f e c t f o u r d e n t / r o l l e r combinat ions w e r e chosen . An a r t i f i c i a l c i r c u l a r d e n t of 200 micron wid th and 3 micron d e p t h was o v e r r o l l e d by r o l l e r s of 2 , 4 , 8 and 16mm r a d i u s . The o v e r r o l l i n g of t h e d e n t is s i m u l a t e d u s i n g 9 d i f f e r e n t p o s i t i o n s o f t h e r o l l i n g element wi th r e s p e c t t o t h e d e f e c t , i n o r d e r t o p i c k up t h e maximum stresses. The p o s i t i o n of t h e c e n t r e x c of t h e r o l l i n g element i s g i v e n by: xc = b (n-5) /2 f o r n=1,2 ,..., 9 . The stress h i s t o r y i n e a c h p o i n t i s a n a l y s e d wi th r e s p e c t t o t h e s e n i n e p o s i t i o n s and t h e n t h e l i f e i n t e g r a l i s c a l c u l a t e d . A s t h e number of p o s i t i o n s i n t i m e and space a r e r e l a t i v e l y smal l , 9 and 49x17 r e s p e c t i v e l y , t h e v a l u e s of t h e l i f e i n t e g r a l s a re r a t h e r jumpy a n d c o n s e q u e n t l y t h e numer ica l r e s u l t s s h o u l d be i n t e r p r e t e d w i t h c a r e . 5 RESULTS The l i f e i n t e g r a l s were c a l c u l a t e d f o r f o u r d i f f e r e n t v a l u e s o f t h e r e d u c e d r a d i u s o f c u r v a t u r e a n d f o r s i x d i f f e r e n t l o a d s ( c o r r e s p o n d i n g t o H e r t z i a n p r e s s u r e s r a n g i n g f rom 2 . 0 t o 3 . 3 GPa) . T h r e e cases were examined, a smooth raceway, a raceway wi th one d e n t a n d a r a c e w a y w i t h o n e d e n t a n d a s s o c i a t e d r e s i d u a l stress f i e l d . The e f f e c t of t h e s e stress f i e l d s on p r e d i c t e d l i v e s c a n be g r a p h i c a l l y e x p r e s s e d i n t e r m s of r i s k maps. I n t h e s e a s e c t i o n of t h e x, z p l a n e i s drawn on a gr id w i t h t h e ' f a t i g u e c r i t e r i o n ' stress e x p r e s s e d as t h e y c o - o r d i n a t e . Each map i s n o r m a l i s e d t o t h e smooth c a s e by a s c a l i n g f a c t o r . The smooth c a s e r i s k map i s shown i n f i g u r e 6 ( a ) where a s e x p e c t e d t h e h i g h e s t r i s k o c c u r s a t t h e p o s i t i o n of t h e maximum o r t h o g o n a l s h e a r stress, 0 .8b below t h e b e a r i n g s u r f a c e . AS t h e d e n t ( f i g u r e 6 ( b ) ) and t h e d e n t p l u s r e s i d u a l stresses ( f i g u r e 6 ( c ) ) are i n c l u d e d t h e map i s m o d i f i e d , 192 p a r t i c u l a r l y a round t h e d e n t s h o u l d e r s . A s a r e s u l t t h e s c a l e f a c t o r , e f f e c t i v e l y a measure of t h e i n c r e a s e d r i s k , i n c r e a s e s d r a m a t i c a l l y , by a f a c t o r of a lmost 50 on t h e i n c l u s i o n of t h e r e s i d u a l stresses. The p r e d i c t e d l i v e s f o r each d e n t / r o l l e r combination a r e p l o t t e d i n f i g u r e 7 . From t h i s d a t a an approximate map of r e l a t i v e l i v e s can b e c o n s t r u c t e d i n t e r m s of t h e d e n t s i z e , c o n t a c t s i z e and r o l l e r r a d i u s o f c u r v a t u r e ( f i g u r e 8). A s can be s e e n t h e l i f e of t h e smooth r a c e w a y i n c r e a s e s r a p i d l y w i t h d e c r e a s i n g l o a d a n d t h e l i f e g e n e r a l l y i n c r e a s e s w i t h i n c r e a s i n g r a d i u s . The i n f l u e n c e of t h e dent wi thout t h e a s s o c i a t e d r e s i d u a l stresses on l i f e i s m i n i m a l . A s i g n i f i c a n t l i f e r e d u c t i o n o n l y occurs f o r t h e s m a l l e s t r a d i u s u n d e r t h e t h r e e l i g h t e s t l o a d s . T h i s c h a n g e s d r a m a t i c a l l y when t h e ( t e n s i l e ) r e s i d u a l stress f i e l d below t h e d e n t i s t a k e n i n t o a c c o u n t . The r e s i d u a l stresses were o b t a i n e d assuming no r a d i a l t r a c t i o n f o r c e a t t h e i n t e r f a c e , i e a f l a t p r e s s u r e d i s t r i b u t i o n . Under h i g h l o a d s t h e i n f l u e n c e o f t h e r e s i d u a l stress f i e l d s a r e r e l a t i v e l y s m a l l , b u t a s t h e l o a d i s reduced t h e l i v e s d e c r e a s e m a r k e d l y r e l a t i v e t o t h e smooth c a s e s . A s t h e r a d i u s o f t h e c o n t a c t i s i n c r e a s e d t h e i n f l u e n c e o f t h e r e s i d u a l s t resses a n d o f t h e d e n t g e o m e t r y i s diminished. I n s p e c t i o n o f t h e o r t h o g o n a l s h e a r stress c o n t o u r s ( f i g u r e 9) h e l p s e x p l a i n why t h i s might b e s o . A t h i g h e r l o a d s a l t h o u g h t h e o r t h o g o n a l s h e a r stresses of t h e m o d i f i e d H e r t z i a n f i e l d a r e of a h i g h e r magnitude, t h e maxima a r e s i t u a t e d w e l l below t h e s u r f a c e . The stress c o n c e n t r a t i o n s from t h e shoulder of t h e d e n t and p a r t i c u l a r l y t h e t e n s i l e stresses of t h e r e s i d u a l stress f i e l d l i e much more c l o s e l y t o t h e s u r f a c e and c a n n o t t h e r e f o r e combine w i t h them t o g e n e r a t e h i g h v a l u e s of t h e f a t i g u e c r i t e r i o n . A s t h e l o a d i s reduced t h e stress c o n t o u r s l i e more c l o s e l y t o t h e s u r f a c e a n d c a n combine w i t h t h e t e n s i l e r e s i d u a l stresses t o cause more damage. The r e s u l t s s h o u l d be i n t e r p r e t e d w i t h some c a u t i o n because of t h e s i m p l i f i c a t i o n s made and t h e c o a r s e g r i d numerics and s t r ic t q u a n t i t a t i v e c o n c l u s i o n s s h o u l d n o t b e made. However , t h e q u a l i t a t i v e r e s u l t s a r e i n t e r e s t i n g enough t o c o n t i n u e t h e r e s e a r c h i n t h i s d i r e c t i o n , i n c o r p o r a t i n g more r ea l i s t i c d e n t shapes , r e s i d u a l stress f i e l d s and more a c c u r a t e c a l c u l a t i o n s on f i n e r g r i d s . 6 CONCLUSIONS The e f f e c t of b o t h d e n t s i z e and s u b s u r f a c e r e s i d u a l stresses have been added t o t h e o r i g i n a l l i f e r e d u c t i o n work. The s l i p l i n e f i e l d a n a l y s i s c a n o n l y b e r e a l i s t i c a l l y a p p l i e d t o deep t r a n s v e r s e i n d e n t a t i o n s where t h e assumptions of p l a n e p l a s t i c s t r a i n can be j u s t i f i e d and t h e d r y c o n t a c t a n a l y s i s i s p r o b a b l y o n l y r e a s o n a b l e f o r r e l a t i v e l y t h i n f i l m c o n d i t i o n s . However t h e t r e n d s i n t h e l i f e r e d u c t i o n f a c t o r s are d i s t i n c t i v e and t h e e x p l a n a t i o n f o r them would s e e m reasonable and a p p l i c a b l e t o a n y d e n t p r o f i l e / r o l l e r combinat ion. The most s t r i k i n g outcome i s t h a t where f a i l u r e i s i n i t i a t e d t h r o u g h s u r f a c e i n d e n t a t i o n and a s s o c i a t e d r e s i d u a l stresses (and c o n s i d e r a b l e e v i d e n c e e x i s t s t o s u g g e s t t h i s i s s o ) t h e e x p e c t e d l i v e s may n o t i n c r e a s e w i t h d e c r e a s i n g l o a d a s r a p i d l y a s would b e p r e d i c t e d by convent iona l models. The r e d u c t i o n i n expec ted l i v e s i s v e r y s e n s i t i v e t o t h e s i z e of d e n t i n r e l a t i o n t o t h e r o l l e r r a d i u s a n d t h i s may w e l l have i m p o r t a n t consequences i n t e r m s o f c r i t i c a l p a r t i c l e s i z e and s a f e and u n s a f e l e v e l s of f i l t r a t i o n . To b e a b l e t o draw q u a n t i t a t i v e c o n c l u s i o n s f u r t h e r r e s e a r c h i s needed t h a t uses more r e a l i s t i c r e s i d u a l stress f i e l d s and f i n e r grids i n t h e l i f e c a l c u l a t i o n s . 7 ACKNOWLEDGEMENTS W e would l i k e t o thank D r Andrew Olver f o r h i s h e l p f u l a d v i c e i n t h i s work and t o register our g r a t i t u d e t o SKF-ERC, The Nether lands, who have s p o n s o r e d t h i s work, and t o D r I a n Leadbetter, Managing D i r e c t o r of SKF-ERC f o r permiss ion t o p u b l i s h . APPENDIX References 31 4 1 91 Webster, M . N . , Ioannides , E . and S a y l e s , R . S . , ( 1 9 8 5 1 , \"The E f f e c t o f T o p o g r a p h i c a l D e f e c t s on t h e C o n t a c t Stress and F a t i g u e L i f e i n R o l l i n g Element Bearings\" , Proceedings of t h e 1 2 t h LeedsLyon Symposium on T r i b o l o g y , Lyon, But te rwor ths , Vo1. 12, pp. 121-131. Lundberg, G . and Palmgren, A . , ( 1 9 4 7 ) , \"Dynamic C a p a c i t y o f R o l l i n g Bear ings \", Acta P o l y t e c h n i c a , Mechanical Engineer ing series , Royal Academy o f E n g i n e e r i n g Sc iences , Vol. 1, No 3, 7 . I o a n n i d e s , E . and H a r r i s , T . A. , (1985) , \"A N e w F a t i g u e L i f e Model f o r R o l l i n g B e a r i n g s \", ASME J o u r n a l of L u b r i c a t i o n Technology, , Vol. 107, pp. 367-378. Hamer, J . C . , Sayles , R. S . and Ioannides , E . , ( 1 9 8 5 ) , \"Deformation Mechanisms and S t r e s s e s C r e a t e d by 3 r d Body D e b r i s C o n t a c t s and T h e i r E f f e c t s on R o l l i n g Bear ing F a t i g u e \", Proceedings of t h e 1 4 t h Leeds-Lyon Symposium on Tr ibology, Lyon, But te rwor ths , Vol. 1 4 . Hamer, J. C . , Sayles , R. S. and Ioannides , E., \" P a r t i c l e Deformation and Counter face Damage When R e l a t i v e l y S o f t P a r t i c l e s a r e S q u a s h e d Between Hard A n v i l s \" , T r a n s ASME/STLE t o be p u b l i s h e d . Dumas, G . and B a r o n e t , C . N . , ( 1 9 7 1 ) , \" E l a s t o - p l a s t i c i n d e n t a t i o n of a h a l f - s p a c e b y a l o n g r i g i d c y l i n d e r \" , I n t e r n a t i o n a l J o u r n a l o f M e c h a n i c a l Sc iences , Vol. 13, 519. Olver , A . V . , (19861, \"Wear of Hard S t e e l i n L u b r i c a t e d , R o l l i n g C o n t a c t \" , Phd Thes is , I m p e r i a l Col lege . Olver , A . V . , Sp ikes , H . A . , Bower, A . and Johnson, K . L . , ( 1 9 8 6 ) , \"The R e s i d u a l S t r e s s D i s t r i b u t i o n i n a P l a s t i c a l l y Deformed Model Asper i ty\" , Wear, Vo1. 107, H i l l , R . , (1950) , \"The Mathematical Theory pp. 151-174. o f P l a s t i c i t y \", Oxford U n i v e r s i t y Press. 1 0 1 Ford , H . , ( 1 9 6 3 ) , \"Advance Mechanics of Mater ia l s\" , Longmans . 113 Lubrecht , A . A . , \"The Numerical S o l u t i o n of t h e Elas tohydrodynamica l ly L u b r i c a t e d L i n e and P o i n t C o n t a c t Problem, Using M u l t i g r i d Techniques\", Phd Thes is , Twente U n i v e r s i t y , l 9 8 7 , The Nether lands . 193 [121 Brandt, A. and Lubrecht, A. A., 'Multilevel Multi-Integration and Fast Solution of Integral Equations\", to be published in the Journal of Computational Physics. [131 Brandt, A. , \"Multigrid Techniques: 1984 Guide with Applications to Fluid Dynamics\", monograph available as GMD studien 85 from GMD postfach 1240 , Schloss Birlinghofen D5205 St. Augustin 1 BRD. coefficient (p) of: (a) p = 0 and (b) p = 0.2. %/A 000 ? Figure 4 Residual orthogonal shear stress (Q contour map. 194 shear stress (2,) distribution during the overrolling of the dent. 195 Figure 6(c) Risk map of indented surface plus residual stresses. Scale factor49.7 196 stresses. 197 n 9 17 33 65 129 257 513 1025 time 0.8 1.5 2.4 3.9 8.4 23.0 79.0 306.0 time * 6.2 12.3 24.3 48.4 Table 1: Computing time as a function of the level (L), the number of points n, in seconds on a VAX 785, for the smooth dry line contact problem, with (*) and without MLMI." ] }, { "image_filename": "designv11_20_0000622_cae.20045-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000622_cae.20045-Figure3-1.png", "caption": "Figure 3 The schematic representation of the DC motor module.", "texts": [ " ControlDesk is an experiment environment tool to manage Simulink and real-time experiments and for seamless controller development. With the integrated control of Simulink simulations, user can validate his/her controller models offline. It can be used for virtual instrumentation, automation, and handling parameter sets. Figure 2 shows the general structure of software interface and hierarchy. The DC motor module [14] consists of a DC motor, a potentiometer, a tachogenerator, and electronic circuits. Figure 3 depicts schematic representation of the DC motor module. The motor is capable of being driven at speeds of up to 2,500 rpm in either direction. The motor drives a slotted disk for the angular speed of the motor and a Gray-code disk for the angular position of the motor. An adjustable magnetic brake (eddy current brake) introduces a viscous term and it allows the user to change the parameters of the system in real-time. The motor output is geared down by a ratio of 9:1 to drive the output shaft. The output shaft has a calibrated disk to show the angular rotation of the shaft" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001952_3.49147-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001952_3.49147-Figure3-1.png", "caption": "Fig. 3 Notation used for pure bending.", "texts": [ " In either case it is necessary to determine the eigenvalue of the symmetric mode and the antisymmetric mode. The buckling load will always have the lower eigenvalue. This analysis can be easily performed by a digital computer, requiring a minimum of computer time and core memory. Appendix A, Formulation for Coefficients linearly with the number of the points on the meridional direction and with the cube root of the wave number. In the first case, the buckling loads and mode of buckling are calculated for R/t = 100, L/R = 1/2, Poisson's ratio being taken as 0.3 (Fig. 3). The prebuckling solution is worked out for the boundary conditions v = w = M^ = 0 and N^= \u2014 cos9. The number of terms in the Fourier series for the buckling state was limited to n = 0,1,2,..., 9 (n = 9), since the summation beyond the ninth term does not affect the final result of the computations. A uniform finite difference mesh with 30 and 46 points in the axial direction was selected. The difference between results for 30 and 46 points was about 1%. The critical bending stress was 1.016 greater than the critical stress of a cylinder under uniform compression" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002569_aim.2009.5230020-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002569_aim.2009.5230020-Figure4-1.png", "caption": "Fig. 4. DOF configuration of WL-16RV.", "texts": [ " Changing a position gain value is a feedforward approach, and a foot landing-impact force is reduced by a large position following error. However, it is difficult to realize a precise positioning control with a low position gain, and a walking robot becomes unstable in a stance phase. So we change back to a high position gain value after detecting a foot landing on a ground. On the other hand, the leg mechanism of WL-16RV consists of a parallel linkage mechanism called the Stewart Platform (see Fig. 4). Because it has a higher stiffness compared with a serial linkage mechanism, it is not sufficient to obtain a high compliance of the landing-foot only by changing a position gain value as mentioned above. Therefore, we realized a larger position following error by raising the foot\u2019s edge of the traveling direction and concentrating a landing-impact force to an actuator nearest to a contact area as shown in Fig. 5. As a result, we could obtain a higher compliance against ground reaction forces. C" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003259_185504-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003259_185504-Figure1-1.png", "caption": "Figure 1. Absorption spectra of (a) GNRs, (b) GOx , and (c) GOx\u2013GNR bioconjuate. The inset figure is a TEM image of GNRs synthesized by the seed growth method (aspect ratio = 3.5).", "texts": [ " The high energy band is induced by the oscillation of the electron perpendicular to the rod axis (transverse surface plasmon absorption). The other absorption band, which is red-shifted to lower energies, is caused by the oscillation of electrons along the rod axis (longitudinal surface plasmon absorption) [20]. Here, in order to obtain SPR of GNRs in the near-infrared (NIR) region (an optical transparency window, where biological tissue and water absorb minimally) so as to avoid GOx damage under low laser density, GNRs with an aspect ratio of 3.5 were particularly synthesized to modify the electrodes. The inset of figure 1 shows GNRs prepared via the seed-mediated method as mentioned in section 2.2. The average width of as-prepared GNRs is 15 nm and the average length is 53 nm. Figure 1 shows the absorption spectra of GOx , GNRs and GOx\u2013GNR aqueous solution. The transverse surface plasmon absorption of GNRs was excited at 520 nm and the longitudinal one at 762 nm (curve a). The absorption peak of pure GOx lies at 374 and 450 nm (curve b), consistent with the previous report [21]. After the conjugation of GOx with GNRs (curve c), the peaks are slightly shifted to 372, 445 and 767 nm; these shifts may be attributed to the dielectric environment changing after GOx absorption on GNRs [18]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.38-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.38-1.png", "caption": "Fig. 14.38 Truck mixer with drum driven by the truck\u2019s engine", "texts": [ "3 \u2022 Protection against agitator drive system overload: the use of hydraulic torque converters prevents drive system overloads and ensures a smooth start of the rotation of the agitators as the mixer is filled; also elastic suspension of the mixing blade arms protects the drive against overloading and deformation of the mixing blade arms.\u2022 Quick distribution of make-up water in the mass of dry components to reduce mixing time: several suitably arranged water nozzles ensure that the concrete mix quickly becomes homogenous.\u2022 The concrete mixer\u2019s dimensions should be suitable for transport on public roads.\u2022 The use of abrasion-resistant materials for the blades and the linings. The truck mixer (Fig. 14.38) is designed for producing homogenous concrete mix and transporting it over long distances. It consists of a pear-shaped drum (usually a freefall, reversing one) inclined at an angle of 15\u25e6 and a self-propelled chassis or a trailer. Its accessories include: a water tank, a water dosage unit, a charging hopper, and discharging chutes. The drum is supported by a cylindrical pin on the drive side and by two rollers mating with a rigid ring fastened on the drum. As con- crete mix components are loaded into the drum and mixed, the drum revolves in one direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003231_xst-2010-0242-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003231_xst-2010-0242-Figure3-1.png", "caption": "Fig. 3. Haptic device simulating a hand-piece with a bur. (A) 6D input of PHANToM Desktop simulating the bur center (C) and (X, Y and Z) axes. (B) A bur attached micro (pen-like) hand-piece.", "texts": [ " The method of deceasing voxel value however cannot reveal how the tool contacts tissues, therefore may roughly predict the force magnitude but cannot precisely predict the force direction. This study also uses the change of voxel distance-levels to obtain the contact information for accurate burring force calculations. Figure 2 shows the system architecture. Currently, a general PC with an IntelCore 2 Duo E8500 (3.16 GHZ) CPU, 3 Gbytes of main memory and a NVIDIA GeForce9600 GT graphics card is used. The simulator is equipped with a highly reliable haptic device, PHANToM Desktop (by SensAble Inc.) to provide 6D inputs and to output 3D translational forces [36]. As shown in Fig. 3, a 6D input simulates a bur position and (X, Y and Z) axes. The main (x, y and z) axes of the volume coordinate system indicate the directions of pixel rows and columns in parallel CT or MRI slices and the direction between the slices. The volume coordinates of all voxel centers are positive integers. The transformation between the bur and the volume coordinate systems includes a rotation, a translation and a scaling. The translation and rotation responds to the 6D device input, meanwhile the scaling indicate the slice interval and the pixel width in the slices" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001838_0094-114x(75)90072-5-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001838_0094-114x(75)90072-5-Figure2-1.png", "caption": "Figure 2. Reflection polygons in the elliptic pitch curves.", "texts": [ " In a similar way a crossed belt of length 4(a + e), encompassing the quadrangle E,F2E2F~ along its four sides, might be introduced. The second author raised the question whether there possibly exists an oval g~ between the ellipse p, and its focal segment E,F, which could be connected with a corresponding oval g2 by a closed crossed belt. Using the three-pole theorem of Aronhold-Kennedy[4, section 25] it becomes clear that in any case the crossing point of the straight parts of the belt must coincide with the pole P. To answer Zenow's question, let us begin with an arbitrary straight line through the pole P (Fig. 2). This line, representing an initial position of one straight part of the belt, cuts the ellipse p, in a second point P~, forming there with p, a certain angle a ' . After a certain time, P; coincides with the corresponding point P~ of p2, symmetric to P; with respect to the common ellipse tangent t at P. At this moment P~ = P~ is the new instantaneous center, and the new position of the segment PP ~ determines the other straight part of the belt. Its protraction P ~P',' forms at P ~ the same angle a ' with p~ and meets p2 in a second point P~. Proceeding in this way we see from Fig. 2 that there arises a reflection polygon PP ;P'I'... of p, whose sides are all tangents of the required curve g,, and an equal reflection polygon PP~P~.. . of p2 is circumscribed to the corresponding curve g2. Now a well-known theorem, due to J. V. Poncelet (1822), states that any reflection polygon of a conic is circumscribed to a confocal conic (which may degenerate to the pair of focal points). This theorem, usually derived by means of the projective theory of confocal conic systems[l, section 30], will be proved by elementary methods" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000489_bf00467722-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000489_bf00467722-Figure2-1.png", "caption": "Fig. 2. Voltammetric cell for automatic water analysis (for designations see text)", "texts": [ " The polarograph of own construction with the differential pulse mode as only voltammetric mode and an incorporated potentiostat is connected to the working electrode (WE), the reference electrode (RE) and the auxiliary electrode (AE) in the voltammetric cell. The automatic performance of the voltammetric analysis is controlled by the microprocessorized \"Programcontroller\". The resulting voltammogram is recorded by the \"Plotter-printer\" unit which also prints out the adjusted instrumental parameters of the polarograph and other experimental details of the measurement. The voltammetric cell depicted in Fig. 2 was designed to provide a rapid automatic rinsing and filling with the water to be analysed. In the cover of the cell are placed the holes for the auxiliary electrode (AE), the working electrode (WE), realized as a hanging mercury drop electrode, the reference electrode (RE) and the inlet tube for nitrogen T 3. The inlet tube for the addition of an acid T 4 and the inlet tube for the addition of the standard solution T s are placed in the wall of the cell. The sample flows into the cell via the inlet tube T1 and out of the cell via the overflow outlet tube T 2 serving also as gas outlet" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000124_ed082p712-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000124_ed082p712-Figure4-1.png", "caption": "Figure 4. (Top) Schematic diagram of the potential distribution as a function of distance, x, at the electrode\u2013electrolyte interface in the presence and absence of a supporting electrolyte (e.g., 0.1 M KCl). The driving force for electron transfer, \u03c6S \u2212 \u03c6PET, is a function of the electrostatic potential, \u03c6(r), at the plane of electron transfer (PET). Lower ionic concentrations results in smaller values of \u03c6S \u2212 \u03c6PET. (Bottom) Steady-state voltammograms corresponding to the oxidation of IrCl63 at a 160-nm radius Au electrode in the presence and absence of a supporting electrolyte. The drawn-out i\u2013V response in the absence of electrolyte is due to a decreased driving force for electron transfer.", "texts": [ "1 M KCl, the reaction is very fast (k \u223c 5 cm s) and is difficult to measure by conventional electrochemical methods. The value of k is also a function of the concentration of the supporting electrolyte, a consequence of the fact that the driving force for electron transfer depends on magnitude of the electrostatic potential drop between the electrode and the plane of electron transfer (PET), the latter corresponding to the distance between the redox molecule and the electrode at the moment of electron transfer (i.e., the tunneling distance). As shown in Figure 4, the driving force for electron transfer corresponds to (\u03c6S \u2212 \u03c6PET), which represents the difference between the electrostatic potential at the electrode, \u03c6S, and in the solution at the PET, \u03c6PET. Similar to the Debye length, \u03ba 1, the value of (\u03c6S \u2212 \u03c6PET) is a function of the supporting electrolyte concentration, decreasing as the ionic strength is lowered as shown in Figure 4. Although this is not completely understood, values of k for some redox reactions are very sensitive to (\u03c6S \u2212 \u03c6PET) while others are not. The oxidation of Ir(Cl)6 3 is an example of a redox reaction whose kinetic rate is particularly sensitive to (\u03c6S \u2212 \u03c6PET), as demonstrated in the lower part of Figure 4. 718 Journal of Chemical Education \u2022 Vol. 82 No. 5 May 2005 \u2022 www.JCE.DivCHED.org In this set of experiments, a 160-nm radius Au electrode was used to record the i\u2013V response corresponding to the oxidation of Ir(Cl)6 3 in the presence and absence of 0.1 M KCl. In the presence of the supporting electrolyte, the i\u2013V responses has the classic sigmoid shape that is characteristic of a reaction limited solely by molecular transport, i.e., k >> D r. Indeed, the electron-transfer rate is immeasurable fast for these solution conditions at a 160-nm radius electrode. On the other hand, when the supporting electrolyte is not present, the i\u2013V response becomes drawn out, suggestive of slow electron transfer, k << D r. Since D is not very sensitive to the electrolyte concentration, and since the same electrode is used in recording the two responses in Figure 4 (i.e., D r is constant), it follows that the change in the voltammetric response is due to a decrease in k when the electrolyte is removed. In more systematic studies, the i-V response can be used to determine k as a function of electrolyte concentration, as well as to estimate the corresponding values of (\u03c6S - \u03c6PET) (36). This type of data would be impossible to obtain using larger electrodes for several reasons. First, although the reaction appears sluggish in the absence of an electrolyte at the 160-nm radius electrode, it is still sufficiently fast that mass transfer would be rate limiting at any electrode of radius greater than ca" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000082_2004-01-2830-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000082_2004-01-2830-Figure5-1.png", "caption": "Figure 5: The aircraft parts are attached on the datum points", "texts": [ " The robot, the pick-ups and the Metrology Probe have a female Capto chuck. The robot chuck is locked automatically, whereas the probe and pick-up chucks are locked manually. THE PICK-UPS After the robot has positioned the Dynamic Modules to an accuracy of 50 \u00b5m, the pick-ups are attached on the top-plate of a Dynamic Module. The important accuracy however is the datum point. A datum point is the position in Cartesian space that defines where the aircraft parts are to be located, or \u201cpicked up\u201d, see figure 5. One issue is the fact that the position given by the robot during configuration of the Dynamic Module, ensures accuracy of the Capto interface and not the datum point. Figure 5 shows the relation between the male Capto interface on the Dynamic Module (DM) that is position by the robot and the datum point that is attaching the aircraft part. This relation was calibrated using a Coordinate Measurement Machine (CMM). From the measurements the CAD model of the Pick-up was updated. The updated value is in fact the difference between the nominal configuration and the measured relation. As a result the male Capto position in Cartesian space was updated in the simulation model, and the DM was re-positioned slightly, to compensate for the difference between the digital model of the Pick-up and the physical one" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002455_j.microc.2008.01.001-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002455_j.microc.2008.01.001-Figure2-1.png", "caption": "Fig. 2. Voltammograms of the solution on a glassy carbon electrode with a scan rate of 200 mVs\u22121 at pH 10.0; (a): the blank solution; (b): in the presence of 40.0 \u00b5mol L\u22121 Ru(bpy)2 +2; (c): in the presence of 40.0 \u00b5mol L\u22121 sulfide; (d, e and f): in the presence of Ru(bpy)2 +2 with 30.0, 40.0 and 50.0 \u00b5mol L\u22121 thiosulfate, respectively.", "texts": [ " This new oxidation wave is produced as a result of the interaction of the Ru(bpy)2 2+ and sulfite species at the surface of the electrode. Fig. 1B displays the effect of increasing the Ru(bpy)2 2+ concentration on the shape and magnitude of the peak currents. As it can be observed in high concentration of Ru(bpy)2 2+ in comparison with that of Fig. 1A, the peaks are well separated and the increase of sulfite concentration in low concentration regime results in the increase of new wave intensity. On the other hand, according to the results of Fig. 2, addition of high quantities of thiosulfate to the solution containing Ru(bpy)2 2+ causes to sharply increase the new formed anodic peak current at the surface of glassy carbon electrode. This dramatic increase of the peak intensity in the high concentration regime of the sulfur ions causes to the concomitant of the current peak of Ru(bpy)2 2+ and that of new wave obtained as a result of interaction of the sulfur anions and Ru(bpy)2 2+ ions. 3.2. Proposed mechanism of the interaction of Ru(bpy)3 2+ and sulfur species As it can be seen in Scheme 1, in the first step Ru(bpy)2 2+ can be oxidized at the surface of glassy carbon electrode to Ru(III)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000844_iemdc.2005.195712-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000844_iemdc.2005.195712-Figure10-1.png", "caption": "Fig. 10 Open-circuit field distributions with 10-pole rotor at peak cogging torque position.", "texts": [ "1 1st 2nd 3rd 4th 5th 6th 7th 8th 9th 10th 11th 12th 13th 14th 15th Harmonic order C og gi ng T or qu e (N m ) 1-slot 4-slot (b) Harmonics IV. 12-SLOT, 10-POLE MOTOR The synthesis technique has also been applied to a 12-slot, 10-pole motor, Fig. 9. Its outer stator diameter and axial length are 100mm and 50mm, respectively, while its airgap length and magnet thickness are 1mm and 3mm, respectively. The width of the stator slot openings is 2mm, whilst the magnets are again parallel magnetized, have a remanence of 1.2T and a relative recoil permeability of 1.05, a polearc/pole-pitch ratio of 1.0, Fig. 10 shows open-circuit field distributions for 1-slot and 12-slot motors, whilst Fig. 11 shows the associated cogging torque waveforms. Since the least common multiple, Nc, of the 1-slot, 10-pole motor is 10, the cogging torque periodicity is 36 degrees mechanical, while for the 12-slot 10-pole motor, Nc = 60, C = 2, and the cogging torque periodicity is only 6 degrees mechanical, i.e. its frequency is 6 times higher. Only the 6th harmonic and multiples thereof in the cogging waveform torque for the 1- slot motor contribute to the resultant cogging torque waveform of the 12-slot motor" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002872_ijhvs.2009.027138-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002872_ijhvs.2009.027138-Figure7-1.png", "caption": "Figure 7 Tractor front suspension with leaf springs (l) and drive axle (r) (see online version for colours)", "texts": [ " The global coordinate system in the models has the x-axis running along the vehicle longitudinal axis, the y-axis in the lateral direction and the z-axis in the vertical direction, forming an inertial system. The front suspension of the tractor is modelled as a leafspring with two dampers, 1050 mm apart. The distance between the leafsprings is 800 mm. Note that in the picture it seems like the dampers or leaf springs do not touch either of the axles, but this is just the appearance of the model (Figure 7). The trailer axles are made up from two pairs of dampers and one pair of springs. The damping of the leaning9 dampers are about one-fifth of the damping of the upright10 dampers. The distance between the upright dampers, as well as the distance between the springs, is 992 mm. The distance between the leaning dampers is 788 mm. The upright damper contributes to about 30% of the damping in vertical direction, and the leaning damper to about 70%. The tyres used in all models are models of the Michelin 315/80 R22" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002344_jf8018535-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002344_jf8018535-Figure3-1.png", "caption": "Figure 3. Changes in absorption of ascorbic acid and chlorogenic acid in acidified saliva. The reaction mixture (1.4 mL) contained 0.8 mL of saliva and 0.32 mL of 50 mM KH2PO4-KCl-HCl (pH 1.36). The pH and the concentrations of nitrite and SCN- of acidified saliva were the same as in Figure 1. (A) Oxidation of ascorbic acid (0.1 mM). Trace 1, no addition; trace 2, 0.1 mM chlorogenic acid; trace 3, 0.2 mM NaNO2; and trace 4, 0.1 mM chlorogenic acid + 0.2 mM NaNO2. (B) Oxidation of chlorogenic acid (0.1 mM). Trace 1, no addition, trace 2, 0.1 mM ascorbic acid; trace 3, 0.2 mM NaNO2; and trace 4, 0.1 mM ascorbic acid + 0.2 mM NaNO2. Upward arrows in A and B indicate the addition of ascorbic acid and chlorogenic acid, respectively.", "texts": [ " If SCN- enhanced the NO formation, O2 uptake by autoxidation of NO should be enhanced by the reagent. Rates of O2 uptake in the presence of 1 mM nitrite in 50 mM KCl-HCl (pH 2.0) were 0.72 ( 0.08 and 1.11 ( 0.21 \u00b5M/min [means ( SDs (n ) 5), p value ) 0.0049] in the absence and presence of 1 mM NaSCN, respectively, supporting SCN--dependent reduction of nitrous acid to NO. Spectrophotometric Measurements. If ascorbic acid and chlorogenic acid reduced nitrous acid to NO, oxidation of the two acids should be observed. Figure 3 shows time courses of the oxidation of ascorbic acid (A) and chlorogenic acid (B) in acidified saliva that contained 0.09 mM nitrite and 0.23 mM SCN-. By the addition of 0.1 mM ascorbic acid, absorption increased at 250 nm due to the absorbance of ascorbic acid, and the increased absorbance decreased slowly (trace A-1). The absorption decrease was enhanced by 0.1 mM chlorogenic acid (trace A-2). In trace A-2, the absorption increase observed by the addition of chlorogenic acid had been subtracted", " The result suggests that ascorbic acid reduced the oxidation intermediate or product of chlorogenic acid and that the oxidation of chlorogenic acid became to be observed after almost all ascorbic acid had been extinguished. No changes in absorption were observed when ascorbic acid (0.1 mM) was added to 50 mM KCl-HCl (pH 2.0). Chlorogenic acid has absorption peaks at about 320 nm, and the absorption decreases during the oxidation (15). By the addition of 0.1 mM chlorogenic acid to acidified saliva, slow absorption decrease following slow absorption increase was observed (Figure 3, trace B-1). The addition of chlorogenic acid in the presence of 0.1 mM ascorbic acid resulted in the continuous slow absorption increase (trace B-2). At present, it is not clear how the absorption increased. When 0.2 mM NaNO2 was added to acidified saliva and then chlorogenic acid was added, a clear absorption decrease due to the oxidation of chlorogenic acid was observed (trace B-3). The absorption decrease was inhibited by ascorbic acid, suggesting that ascorbic acid reduced the oxidation intermediate or product of chlorogenic acid to the original compound (trace B-4). The inhibition disappeared during incubation. The disappearance was related to the extinguishment of ascorbic acid from the reaction mixture as described above. The relation became clear by comparing time courses of absorption changes in Figure 3A,B. In 50 mM KCl-HCl (pH 2) that contains 0.01 mM ascorbic acid and 0.01 mM chlorogenic acid, preferential oxidation of ascorbic acid (about 4 \u00b5M/min) was observed at the initial period of incubation when 0.1 mM NaNO2 was added (data not shown). The oxidation of chlorogenic acid appeared when almost all ascorbic acid had been oxidized, and chlorogenic acid (0.02 and 0.4 mM) enhanced the oxidation of ascorbic acid, which was induced by 0.1 mM NaNO2, about 1.5- and 2-fold, respectively, supporting the oxidation of ascorbic acid by the oxidation intermediate or product of chlorogenic acid", " This result suggests effective reduction of ONSCN to NO by ascorbic acid without producing NO-ascorbic acid and supports the proceeding of reaction 3 in the mixture of saliva and gastric juice. Ascorbic acid-dependent oxygen uptake in acidified saliva was enhanced by 1 mM nitrite. According to the above discussion, the enhanced oxygen uptake could be attributed to the increase in rates of reactions 1, 3, and 4 to produce NO, which resulted in the increase in the rate of reaction 5. The enhancement of O2 uptake by chlorogenic acid (Figure 2) was due to the reduction of nitrous acid to NO by the acid, and the enhancement of oxidation of ascorbic acid by chlorogenic acid (Figure 3) was due to the reduction of the oxidation intermediate or product of chlorogenic acid by ascorbic acid. As the oxidation intermediate and product, chlorogenic acid radical (4) and the quinone form (7) were possible. When ascorbic acid had been extinguished, the quinone form might be accumulated. The quinone form can react with SCN-, producing a stable component as described in the Introduction, although SCN- did not enhance the reduction of nitrous acid to NO by chlorogenic acid as reported previously (4, 6)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001916_j.fusengdes.2008.05.042-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001916_j.fusengdes.2008.05.042-Figure5-1.png", "caption": "Fig. 5. Actual in-vessel situation (left) and virtual situation", "texts": [ " During actual operation, the operator sometimes needs addiional information that cannot be obtained by cameras even if they an be installed in the vacuum vessel. For example, during instalation of the blanket module, the module may interfere with the eys on the vacuum vessel. The operator may want to confirm the ituation behind the module, but the line of sight is interrupted by he next module. In this case, the simulator can provide such infor- ation by making the next module transparent on the screen, as hown in Fig. 5. .4. Testing a taught sequence (off-line) The simulator can also be used for testing a taught sequence efore actual operation. Using the teaching points in the sequence s inputs to the control system, the 3D model\u2019s motion can be eviewed to confirm the interaction between the manipulator nd other components. If the manipulator interferes with any omponents, the simulator will give an alarm showing the relvant parts in red, as shown in Fig. 6. The simulator can also how the distance between components that potentially interfere ith each other, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000883_tsmca.2006.878982-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000883_tsmca.2006.878982-Figure8-1.png", "caption": "Fig. 8. Parallel translation and rotation.", "texts": [ " 7 are obtained from \u03c81(t), \u03c81(t) and (13)\u2013(17) as follows: LU (t) LC(t) LB(t) = lC(t) \u00b7 sin(\u03c6B(t)+\u03c82(t)) sin(\u03b8U (t)\u2212\u03b8B(t)) lB(t) \u00b7 sin\u03c81(t) sin(\u03b8U (t)\u2212\u03b8C(t)) lC(t) \u00b7 sin(\u03c6U (t)+\u03c81(t)) sin(\u03b8U (t)\u2212\u03b8B(t)) . (19) Consequently, the positions Ug(t), Cg(t), Bg(t) in the global coordinate system are calculated from F and LU (t), LB(t), LC(t) as follows: Ug(t) = XF + LU (t) \u00b7 cos \u03b8U (t) LU (t) \u00b7 sin \u03b8U (t) 0 Cg(t) = XF + LC(t) \u00b7 cos \u03b8C(t) LC(t) \u00b7 sin \u03b8C(t) 0 Bg(t) = XF + LB(t) \u00b7 cos \u03b8B(t) LB(t) \u00b7 sin \u03b8B(t) 0 . (20) D. Estimation of Attitude Angles of the Golf Club Head Next, we consider how to measure the attitude angles. Fig. 8 shows how the points U(t), C(t), B(t) on the golf club head are transformed by the translation and rotation. The algorithm to estimate the angles is described by Fig. 8(a)\u2013(f). 1) Detection of the points U(t), C(t), B(t) on the local coordinate grid. The three points U(t), C(t), B(t) are detected from the 1-D image as the position on the local coordinate system. 2) Translation of the three points U(t), C(t), B(t). Translation is done so that the point is on the origin of the local coordinate. The position coordinates U \u2032(t), C \u2032(t), B\u2032(t) after translation are given by U \u2032(t) =U(t) \u2212 C(t) C \u2032(t) = 0 B\u2032(t) =B(t) \u2212 C(t). (21) 3) The three positions Ug(t), Cg(t), Bg(t) on the global coordinate" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000262_tmag.2004.824897-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000262_tmag.2004.824897-Figure8-1.png", "caption": "Fig. 8. Contours of eddy-current loss (1000 rpm). (a) Stator core. (b) Rotor core.", "texts": [ " This tendency becomes small as the rotation speed is high, and disappears at the synchronous speed (1500 rpm). Fig. 5 shows the distribution of the eddy-current density vectors in the secondary conductor. It is found that the eddy currents flow complexly in the bars through the end-ring. Fig. 6 shows the primary current characteristics. It is found that the results with skew are almost the same as those without skew. Fig. 7 shows the torque characteristics. It is found that the results with skew are larger than those without skew. The calculated results agree with the measured ones [7]. Fig. 8 shows the contours of eddy-current loss. It is found that there is a lot of eddy-current loss at the upper and lower parts in the stator teeth surface, and the upper part in the rotor surface. Fig. 9 shows the contours of hysteresis loss. It is found that the hysteresis loss at the upper part in the stator teeth surface and the upper part in the rotor surface are greater than the other part. Table II shows the discretization data and CPU time. We developed a new mesh modification method for the skewed motor analysis using the 3-D FEM" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001496_j.1467-2494.1980.tb00244.x-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001496_j.1467-2494.1980.tb00244.x-Figure5-1.png", "caption": "Figure 5. Example of a melting curve of sebum, showing the mode of determination of the parameters Tc and Sc/S. Sc = shaded area; S = total area.", "texts": [ " We tried to develop a mode of exploitation based upon the determination of the molten fraction at different temperatures. Until now, we are unable to correlate in a satisfactory manner the corresponding values with the results of clinical examinations. Characteristic parameters Examination of the melting curves recorded over an experimental population showed that there is a correlation between the seborrhoeic state of the subject and the characteristic signal which arises on the melting curve between -5OOC and -lODC. Two numerical parameters can be associated to that signal, as shown in Fig. 5: - Sc/S is the area of the characteristic signal (Sc), relative to the whole area ( S ) . - Tc is the temperature associated with this signal determined by extrapolation on the baseline. There is a good correlation between Sc/S and Tc, as shown on Fig. 6 , where the experi- mental points from all our subjects are reported. Correlation of Sc/S and Tc with clinical examination The modalities of clinical examination have been previously described (2). Here we improved the accuracy of that appreciation, using the I,iposcope@ (3), an instrument which operates according to the method proposed by Schaefer (4)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002859_20090630-4-es-2003.00151-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002859_20090630-4-es-2003.00151-Figure3-1.png", "caption": "Fig. 3. Quarter car model", "texts": [ " D M t t Following first principles, equations for the balance of torques at the mantle \u03b8\u03b8\u03c9 && ttDM1 '' dcTJ \u2212\u2212= (5) and at the wheel rim \u03b8\u03b8\u03c9 && ttM2 '' dcJ += (6) can be derived where J1 and J2 are the moments of inertia of the mantle and the rim, respectively. The twist angle \u03b8 between mantle and rim is given by . Applying Laplace Transformation and inserting equation (6) in (5) results in the transfer function \u03c9\u03c9\u03b8 \u2212= M & scJJsdJJsJJ csd sT ssG t21 2 t21 3 21 tt D ')(')( '' )( )()( ++++ + == \u03c9 (7) wb, vertical body acceleration or vertical wheel acceleration . Figure 3 shows the schematic of a vehicle suspension system which is called a quarter car model. Excited by variations in street height z bz&& wz&& h the vertical displacement z of the wheel with its mass mw w and the vertical displacement zb of the vehicle body with its mass mb are subject to oscillations. The wheel is connected to the vehicles body by the suspension system represented by the spring stiffness cb and the shock absorber\u2019s damping factor db. The wheels elastic properties can be expressed by the vertical tire stiffness ct" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003959_2157689.2157836-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003959_2157689.2157836-Figure4-1.png", "caption": "Figure 4: Definition of Viewpoint", "texts": [ " 1) Preparation First, to compute features, we developed a computer program to simulate users' views related to visibility of shops (Figure 3). The program builds 3D polygon data of the environment. It uses perspective transformation and z-buffering, well known 3Dgraphics algorithms, to convert 3D polygon data to 2D pixel data for the given viewpoint. It outputs perceived size of a shop as an amount of 2D pixels, denoted as )(vpvisibilityi for shop i where vp represents the current view point (x, y, ) (Fig. 4). The height of view (h) was set at 150 cm. For each traversed route (denoted as r), the program computed vp for every 125 ms. Regarding the width of view, it would be not a trivial issue to decide the simulated horizontal view because there are a couple of possibilities. Perhaps, people might focus on a narrow view while walking, as they would concentrate on the frontal direction. However, people might use a wider view, or even turn their heads, as they would search for directions to go in. For this problem, there is a successful modeling of pedestrian' decision-making behavior, in which a simulated agent was successful with a horizontal view angle ( h ) of 80 degrees [27]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001493_s0092-8240(78)80008-1-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001493_s0092-8240(78)80008-1-Figure2-1.png", "caption": "Figure 2. Unstable limit cycle for X; = 4.7544103884 and Y; = 4 represented in the phase plane t\" Z. Different trajectories are also represented starting from different initial conditions. When the initial conditions are chosen inside the limit cycle, Y' and Z evolve to the stable steady state (Y;, Z o = X ; + Y;). When the initial conditions are chosen outside the time-periodic solution, Y' is evolving quickly to negative values. The arrows on the line Y' =0 indicate the direction of the vector field (d Y'/dt', dZ/dt'), i.e. the slope of the integral curve Y' = Y'IZ) at Y' =0: (dY'/dZty.-o = - X ~ + 1 <0. This value is always negative, because X~ > 1 in order to insure a value B > 0, since B = Yo (Xo - 1 ). Full lines: temporal evolutions of Y', Z. Dashed line: unstable limit-cycle solution. S.S.T. : stable steady state (Y~,", "texts": [ " However, this branch is subcritical and hence, by a general theorem of bifurcation theory, is bound to be unstable (Sattinger, 1973). Numerical evaluation of the curve ZM(7) away from the bifurcation point 7= 0 shows that the branch remains unstable until a range of values of 7 above which the system admits a single solution. By applying on the steady state (5) perturbations corresponding to Z's higher than the ordinates of the curve ZM = ZM (7) one finds an explosive solution corresponding to negative values of Y', as shown in Figure 2. This confirms the inadequacy of Turing's model to yield physically reasonable time-periodic solutions beyond an instability of the steady-state solution when a differential inequality constraint is not artificially imposed. Zo =Xo + Y~) T h e p u r p o s e o f t h e f o l l o w i n g s e c t i o n is to e x t e n d t h e s e r e s u l t s to t he ca se o f s p a c e - d e p e n d e n t c o n c e n t r a t i o n s . 4. B([~4rcation of Space-Dependent Patterns. As in t h e p r e c e d i n g sec t ion , we sha l l f irst p r e s e n t t he l i n e a r s t a b i l i t y a n a l y s i s , w h i c h p e r m i t s to d e t e r m i n e the c r i t i c a l p o i n t s b e y o n d w h i c h the h o m o g e n e o u s s t e a d y s t a t e s o l u t i o n m a y b i f u r c a t e i n t o n e w t y p e s o f s o l u t i o n " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002907_icems.2009.5382927-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002907_icems.2009.5382927-Figure4-1.png", "caption": "Fig. 4. Experimental equipment of the primary side of the LIM", "texts": [ " PE-EXPERT3 is used for the switching control of voltage source inverter to generate the two different frequency components. The experimental system sets up variable transformer, diode rectifier, voltage source inverter, LC filter, a protection resistance and the LIM. The carrier frequency of the voltage source inverter is 10kHz. The cut off frequency of the LC filter sets up to 1.5kHz. The protection resistance uses not to be short circuit, when it is driven in the low frequency and the DC. The experimental equipment of the LIM is shown as in Fig.3 and Fig.4. Table.1 shows the principle parameters of the LIM. The primary side is the coil of the LIM. The secondary side consists of the iron and aluminum. The attractive force and thrust force is measured by the load cells. The four load cells are located under the primary side and measure the attractive force. The two load cells are located in front of the primary side and measure the thrust force. The air gap is 5mm. LIM is locked rotor test. V. EXPERIMENTAL RESULTS Fig.5 shows the characteristics of the current-attractive force of the LIM when LIM is driven by only the DC power source" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000513_2004-01-1058-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000513_2004-01-1058-Figure10-1.png", "caption": "Figure 10: Influence of TMD on Throttle-Off in a Bend", "texts": [ " The trace called \"VDM Control\" in Figure 9 shows the vehicle behavior tuned with the Vehicle Dynamics Module (VDM), which is part of GKN's control strategy for ETM. This performance was specifically tuned to the requirements of a car manufacturer. Generally it can be stated that this driveline configuration in combination with an intelligent control strategy allows for a tuning of the throttle-off behavior in the range between the trace \"3 Open Differentials\" and the trace \"FA+TC+RA\" shown in Figure 9. Figure 10 shows a principle sketch of the vehicle's driveline. With open differentials, the braking forces at the wheels resulting from the engine's drag torque after throttle-off are equal. With a TMD added to the center differential the braking forces at the rear wheels become smaller while the forces at the front wheels increase as the rear wheels are rotating slower than the front wheels. The outer front wheel has a longer lever arm around the vehicle's CoG (Center of Gravity) than the inner wheel. Therefore, with a TMD the increased force at the front outer wheel causes \u2013 in direct comparison to the version with open differential \u2013 an understeering yaw moment" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002814_elan.200900334-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002814_elan.200900334-Figure1-1.png", "caption": "Fig. 1. Installation of the disposable electrode into the cell body of an electrochemical detection cell (A); A carbon electrode sheet from a titanium and carbon radio frequency sputter deposition (B).", "texts": [ " All chemicals for preparation of standards were obtained from Sigma \u2013 Aldrich (St. Louis, MO, USA) and used without purification. Methanol was from Honeywell (HPLC Grade Solvent). Octanesulfonic acid (OSA) for the preparation of mobile phase was from Dionex (Certified Reagent).Water used for preparation of standards and eluents was of highest available purity (18 MW or better). Cyclic voltammetry was carried out under static conditions in a manually filled flow-through three-electrode detection cell (Fig. 1A). The cell was of the thin-layer type and consisted of a conventional glassy carbon or a disposable carbon working electrode, a silver/silver chloride reference electrode and a titanium counter electrode. A 25 mm-thick cell gasket with a cutout defined the thin-layer flow path (length x width of the flow path: 8.91 1.2 7 mm). Cyclic voltammetric scans were carried out using Electrochemical Workstation (Model 600C, CH Instruments, Austin, Texas, USA). A liquid chromatographic system (ICS3000, Dionex, Sunnyvale, California, USA) was configured with an electrochemical detector using the same cell (Fig. 1A) as specified for cyclic voltammetric measurements. The Chromeleon 6.8 software (Dionex) was used for system control and data acquisition. Separations were achieved on a reverse phase column (Acclaim 120, C18, 3 mm particle size, 100 mm 2.1 mm, Dionex). The mobile phase contained 57 mM citric acid, 43 mM sodium acetate, 0.10 mM ethylenediaminetetraacetic acid (EDTA), 1.0 mM OSA. Additionally, the chromatographic mobile phase contained 10% methanol in all experiments with the exception of those carried out to generate the long-term reproducibility data summarized in Table 2", "de 2010 Wiley-VCH Verlag GmbH & Co. KGaA, Weinheim Electroanalysis 2010, 22, No. 3, 325 \u2013 332 sputtering to improve adhesion of carbon to the polymeric substrate. A carbon layer of ca. 3000 , was then sputtered onto the titanium layer. The purity of utilized carbon target (carbon source) was 99.99%. A mask defining the electrode pattern was utilized during the titanium and carbon deposition steps. The resulting sheets, each holding 24 carbon electrode patterns, were cut to yield the individual electrodes (Fig. 1B). Two alignment holes were punched into each electrode for precise installation in the ED electrochemical detection cell as shown in Figure 1A. The geometric area of the circular working electrode was 0.785 mm2 (1 mm diameter). Plasma preparation kit from Recipe (P/N 1000, US distribution by IRIS Technologies International, Ltd, Olathe, Electroanalysis 2010, 22, No. 3, 325 \u2013 332 2010 Wiley-VCH Verlag GmbH & Co. KGaA, Weinheim www.electroanalysis.wiley-vch.de 327 KS), was used for catecholamine extraction. The kit contained pre-packed columns with a defined amount of activated aluminum oxide, in a buffer solution. A 1.0 mL aliquot of plasma was added to the pre-packed column along with 12" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002914_cdc.2010.5718021-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002914_cdc.2010.5718021-Figure2-1.png", "caption": "Fig. 2. Schematic of jth mobile sensor (not to scale) taken from [14].", "texts": [ " To form an estimated target track, a minimum of \u03b3 elementary detections are needed from \u03b3 distinct sensors within the network. A value of \u03b3 = 3 was found to give accurate tracking by proximity sensors subject to few false alarms, and errors normally distributed with a standard deviation of 20% [15]. Since the AUVs use omnidirectional sensors that follow the isotropic law (24), the field-of-view (FOV) of a sensor, denoted by Cj(t) = Cj [xj(t), rj ], at time t can be defined as a disk of constant radius rj and centered at the sensor position, xj . The sensor can then be viewed as a disk, illustrated in Figure 2 that moves in A consistent with (23). Let the FOVs of the full network of sensors be represented by the set S(t) = {C1(t), ..., CN (t)}, which is characterized by the sensor ranges, r1(t), ..., rN (t), and the AUV positions, x1(t), ...,xN (t). The sensors\u2019 states and the targets\u2019 speeds, headings, and initial positions are regarded as random variables governed by the joint PDFs \u2118(xj , t), fV (V, t), f\u03b8(\u03b8, t), and fT (xT0 , t), respectively. The PDF of the sensors\u2019 states is a function of time since the sensors move to optimize the network\u2019s track coverage, and the PDFs of the target tracks\u2019 variables are assumed to be known functions of time computed using the tracking methods in [15]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003435_j.mechatronics.2010.04.007-Figure12-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003435_j.mechatronics.2010.04.007-Figure12-1.png", "caption": "Fig. 12. Coordinate frame and force diagram of the conveyance system.", "texts": [ " Here, it may be desirable to insert much thinner shield plate than given 2 mm to lessen the air\u2013gap length between PMs and the conductive plate, leading to a stronger force density. But, the very high attractive force due to PMs acts on the shield plate. So, the plate is pulled toward PMs all the times. Therefore, the thickness of the shield plate should be decided by Table 2 Criterions to determine the sign of h3 and h4 from (18) and (19). x2 P 0 <0 y2 considering the structural stability not to be deflected, not by electromagnetic view. Fig. 12 describes a force diagram of the system, including coordinate frames. Z-directed force by each wheel is a repulsive force, and then has self-stability. Therefore, for out-of-plane motion, it is enough to consider only the moment arm from a mass center of the plate to each magnet wheel. Specially, as ascertained in Fig. 6, a thrust force is nearly constant irrespective of a rotating speed, in the speed range larger than 1200 rpm. Therefore, in this region, the normal force can be controlled varying a rotating speed of the wheel" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000131_0094-114x(87)90070-x-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000131_0094-114x(87)90070-x-Figure2-1.png", "caption": "Fig\u00b0 2. Dual vectors.", "texts": [ " s s x ask = - s o c ~ Yj = --S~'Si X a/~ = --(SykCq + CjkSi)\u00a2S) Z / = st's~ = c~c~ - sj~s#c t .~ s= s, 'z~j = ssks/ \u00a3 * = sk\" s j x a u = - sj, cj ~'s = s , ' s t x a u = - ( s ue s , + ctjss, cj) Z j = sk .s~ = ctscjk - suss, c s. Introducing the dual vectors and dual angles in the formulae (1), (2) and (3) gives the corresponding dual formulae, from which we can obtain the real parts which arc the same forms as (1), (2) and (3), and the dual parts that are named secondary formulae. Firstly, the law of the dual scalar product is introduced here. As shown in Fig. 2, the dual vectors t, = r~ + EOPI x rt ~2 = r2 + ~ O P 2 x r 2 where, ~ is the dual unit or operator, and E2= ~3= . . . = O, r I and r2 are unit vectors, Pt and P2 are arbitrary points on lines of r~ and r2 respectively. The scalar product of the dual vectors t, and P: can be expressed in primary part and secondary part. Let i=t \u2022 1~2 = F + ~Fo. F -- r I \u2022 r2 (4) Fo -- PiP2\"r2 \u00d7 rl. (5) Using (4) and (5) and introducing the shorthand notation presented by Duffy[l], we can easily obtain the secondary formulae corresponding to (1), (2) and (3), which can be expressed in the form f R0~" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002984_iros.2010.5650918-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002984_iros.2010.5650918-Figure3-1.png", "caption": "Fig. 3. Single stator", "texts": [ " When an AC voltage is applied to the piezoelectric vibrator, a standing wave is generated on the elastic body. By applying two AC voltages with a phase difference to the positive and negative sections of the piezoelectric elements, a traveling wave is generated due to combination 978-1-4244-6676-4/10/$25.00 \u00a92010 IEEE 3061 of the two standing waves[7]. The stators and the rotor are in pressure contact with each other, and the rotor is driven by the tangential force of the elliptical motion of the traveling wave. A single stator is shown in Fig. 3. Another piezoelectric element on the stator is used as a sensor detecting the resonance, and the signal is called the feedback signal. There are two inputs (AC voltage A and B), one output (Feedback) and FG (Frame Ground) terminals. The stators, namely vibrators, are located as shown in Fig. 2. Geometric parameters (stators\u2019 alignment) are \u03b81, \u03b82, \u03b83 and \u03c6. Using the parameters, the moment vector of each stator, mi, can be expressed as follows: mi = \u2212 cos \u03b8i cos\u03c6 \u2212 sin \u03b8i cos\u03c6 sin\u03c6 \u03c4i, (i = 1, 2, 3) (1) Here, \u03c4i is the generated torque of each stator" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000387_pesc.2004.1355595-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000387_pesc.2004.1355595-Figure1-1.png", "caption": "Fig. 1 Schematic ofopen-end winding induction motor drive fed by 2 -WO level inverters", "texts": [ " In [9]-[ IO], the open-end winding Configuration has been proposed for high power electric vehiclehybrid electric vehicle (EVAIEV) propulsion systems. In [ I I], space vector pulse width modulation technique is used to control the output voltage of both the inverters connected at both ends of the motor winding. It should be noted that the switching frequency capacity of both the inverters is same. This paper proposes ;I new scheme in which one of the inverters is switched al. high frequency, and the other at low switching frequency. A schematic of the open-end winding induction motor drive is shown in Fig.1. A two-level inverter, INVI, feeds the three ends of the stator winding R Y B, and the other three ends R\u2019 Y\u2019 and B\u2019 are connected to another two level inverter called INV2. INVl and INV2 are connected to separate dc sources (bmeries) of magnitude Vdc/2. This can be achieved by simply partitioning the set of batteries. This results in a significant reduction in space, which is a very important aspect in on-board ship propulsion applications. Induction motor voltage equations can be derived in the satne manner as if the machine was connected to a single two-level inverter as follows [9], [14]: Phase voltages in tenns of switch position of respective phases are given by: where SR,S, & S , = 1 or 0 and it depends on the inverter leg switching state" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000132_53.3.305-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000132_53.3.305-Figure2-1.png", "caption": "Fig. 2 A schematic diagram of bending of a flagellar hook. The original structure of the hook ABCFED bends to form a curved structure A\u2032BCFED\u2032. XY is considered as an arbitrary line that passes through the longitudinal axis of the flagellar hook. This arbitrary line is considered as the small section that does not get deformed, even on the bending of the hook. The dark strip on the upper portion of XY represents a small section of area \u03b4A at distance z from XY.", "texts": [ " Grids were examined in an FEI Tecnai 12 Biotwin transmission electron microscope. Calibration was done using catalase crystal with an alternate lattice plane spacing of 8.75 and 6.85 nm (Agar Scientific Ltd, UK). Measurements of the contour length, straight length and diameters of the flagellar hooks, the filaments and the bacterial cells were done using analySIS software (SIS GmbH, Germany). First, we consider the bending deformation of a cylindrical flagellar hook, caused by random thermal forces or fluctuation. Figure 2 shows a schematic diagram of the bending of a flagellar hook. We assume that the hook is initially (i.e. before the deformation) straight (ABCFED in Fig. 2) and unstressed. The bending deformation does not bring any change in the plane of the cross section and the elastic limit of the hook fiber is never exceeded during the deformation. A neutral axis (XY in Fig. 2) passes through the hook, which does not get deformed, even on the bending. Every cross section of the cylindrical hook fiber is symmetrical about an axis perpendicular to the neutral axis of the hook [7]. The segment on the upper portion (i.e. above XY in Fig. 2) of the hook (ABC) will get elongated (A'BC) due to the bending. Let \u03b4A be the sectional area of a small strip under consideration (thick black line in Fig. 2), placed in the upper portion (A'BC), and f be a force that comes to play across the section of the strip to resist its elongation. The magnitude of f will be [7] f = E z \u03b4A / R (1) where E is Young\u2019s modulus, z is the distance of the segment from XY and R is the radius of curvature of the neutral axis. The resisting force f arises from the elastic properties of the hook. Likewise, due to the bending, the lower portion (DEF, below XY in Fig. 2) will get deformed and shortened (D\u2032EF). For a strip (not shown in Fig. 2) of the same sectional area (\u03b4A), situated at the lower portion (D\u2032EF) at a similar distance (z) away from XY, a resisting force of similar magnitude will act on it and will oppose its shortening. The two forces of similar magnitude act oppositely to each other and together generate an internal resisting couple. Taking into consideration the different pairs of equidistant strips and summing up their respective resisting forces, the magnitude of the total resisting force in the complete hook will be \u03a3f = \u03a3E z \u03b4A / R" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.22-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.22-1.png", "caption": "Fig. 14.22 Road pavers", "texts": [ " In order to produce good installation of the hot material, it is necessary to heat the screed\u2019s compaction devices that are in contact with the material. This is done by electric systems powered by a generator that is run by a diesel engine. Furthermore gas and liquid heating systems are used for heating the screed using burners. The working width of road pavers is 1\u201316 m, with deposition layers of up to 0.35 m being possible. Given appropriate material logistics, large road pavers can install more than 1000 t of material per hour [14.21] (Fig. 14.22). The paver tractor consists of a frame to which the chassis is mounted. Furthermore, it contains the driving unit, consisting of a diesel engine, the gearbox, the cooler, and the hydraulic system. The driver stand is situated in the upper part of the vehicle, giving the driver a good view of the material container and the screed. The chassis is either a wheel or a caterpillar chassis. The caterpillar chassis is composed of two hydrostatically driven chain tracks which are electronically controlled" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003828_s00453-010-9399-8-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003828_s00453-010-9399-8-Figure6-1.png", "caption": "Fig. 6 A polygon P with a hole Q can be folded so that P forms one axis and Q forms another, and the interior of P lies between the two axes. The distance between the axes is the width of the ribbon (gray) around Q", "texts": [ " To repair the construction, we first require that the tree of molecules respect the containment relations of boundary components. Thus the molecules interior to a simple polygon, strictly interior to the sheet of paper, must form a proper subtree of the tree of molecules. And if the polygon contains a hole within an island within a hole, then each successive level of molecules must form a proper subtree within the tree above it. We surround each boundary component with a ribbon of width \u03b5 as in [4] or Fig. 6. The deepest molecules then fold to books as in Fig. 5(c). The second-deepest molecules fold to books containing the deepest molecules as \u201cchapters\u201d, contiguous set of flaps within the larger books. The ribbons, not considered part of any molecules, offset the shared axis of the deepest molecules so that it is \u03b5 above (and parallel to) the axis of the second-deepest molecules. We continue in this manner, building books within books, with each level of containment offset from the previous one by \u03b5, something like a layered wedding cake", " The next lemma folds a polygon with holes into a book with two axes. The complexity of the folding increases to (some constant times) the area of the polygon divided by its minimum feature size. We can define the minimum feature size to be the minimum distance between any two non-adjacent edges. Lemma 4 A polygon with holes can be embedded as a book folding so that the outer boundary embeds to one axis, the hole boundaries all embed to a parallel axis, and the interior of the polygon embeds between the two parallel axes. Proof As shown in Fig. 6, we surround each hole with a thin \u201cribbon\u201d (offset polygon) of width \u03b5, where \u03b5 is smaller than polygon\u2019s minimum feature size. In the disk packing step of the construction, we pack the polygon minus the ribbons, and then project the tangency-point folds from the molecules perpendicularly across the ribbons. Thus the interior boundaries of the ribbons fold to the axis, and the boundaries of the holes fold to a parallel line, a new axis, distance \u03b5 from the original axis. By pleating down all the molecules (Fig", " Then as in the case of a topological sphere, we open one more path e to serve as the outer boundary, thereby obtaining a PL manifold M \u2032 homeomorphic to a disk with g pairs of holes, one pair for each handle loop. See Fig. 8. We surround each hole by a ribbon of suitable width \u03b5 > 0. These ribbons will be folded and taped to recreate the handles in the flat folding. In order that the foldings agree so that Q and Q\u2032 can be taped together, we require that the disk packings agree at Q and Q\u2032. We can think of the disk packing as taking place on the uncut 2-manifold with ribbons of width 2\u03b5, as shown in Fig. 8. Each disk at Q projects perpendicularly across the ribbons (as in Fig. 6) to a matched counterpart at Q\u2032, so that the disk at Q and the disk at Q\u2032 agree on the creases crossing the ribbons. The book foldings for the ribbons at Q and Q\u2032 will have top and bottom axes like the folding shown in Fig. 6; the top edges will be used for taping Q to Q\u2032, and the bottom edges will be used to join the handle formed by the two folded ribbons to the body formed by the rest of the folding. The right side of Fig. 8 gives an exploded view of the entire folding, which is a book folding with two parallel axes; handles (gray) form chapters within the body (white). In order for this plan to work, we must be sure that no other pairs of holes will interfere with the taping of Q to Q\u2032. This requirement can be fulfilled by adding some further conditions to the earlier steps of the one-straight-cut algorithm of Sect" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000575_sice.2006.315833-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000575_sice.2006.315833-Figure2-1.png", "caption": "Fig. 2 Flight of golf ball after impact", "texts": [ " To overcome the problem of image capturing speed, the DLT method5) is proposed that uses multiple cameras with a synchronizing capturing system, as illustrated in Fig. 1. However, such a system is expensive and requires a large space for conducting the measurements and analysis. Furthermore, DLT method does not accurately estimate the initial three parameters from multiple image sequences4). To estimate the three parameters accurately, we introduce a novel one-line CCD camera for measuring high-speed flying objects such as a golf ball, since the capture rate of a one-line CCD camera is very high compared with high-speed video camera. Figure 2 shows how the ball flies after impact at the speed of 40 m/s to 80 m/s in the measurement plane. First, the ball is set at initial position (a) just in front of the measurement plane. After impact, the golf ball enters the measurement plane and passes through the plane from position (b) to (c). The rotational speed of the flying golf ball can be determined from the impact point between the golf club and the golf ball. However, it is very difficult to estimate the impact point and impact time accurately" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.156-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.156-1.png", "caption": "Fig. 14.156 Robot for laying tiles on building elevation", "texts": [ " Nevertheless, because of the large share of lining work in finishing work, attempts are being made to automate this field. Progress has been achieved by employing a rubber belt conveyor with suckers so that several tiles can be laid simultaneously. Tile-laying robots are mainly used for covering with large surfaces of exterior walls with tiles. A tile-laying robot moves on rails fixed to scaffolding. A robot for laying tiles on the building\u2019s facade, developed jointly by several Japanese companies, is shown in Fig. 14.156. The robot is intended for laying 227 \u00d7 60 mm (8\u201315 mm thick) tiles on traditional mortar. Its daily capacity is 14 m2. For comparison, a craftsman is able to lay 7 m2 in this time. More information about the robot can be found in [14.53]. Automated building construction systems (ABCS) for high- and medium-rise buildings were developed in Japan as a measure to alleviate the labor shortage in the construction industry. Fourth-generation robots, designed according to the principle of design for robotic construction (DfRC) integrating robot design, building erection, and the materials, are employed for the realization of buildings in such systems" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002924_s11012-010-9279-y-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002924_s11012-010-9279-y-Figure1-1.png", "caption": "Fig. 1 Configuration of the bearing system", "texts": [ " The analysis of this paper was improved upon by Patel, Deheri and Patel [24] to investigate the performance of a squeeze film between rough truncated conical plates in the presence of a magnetic fluid as a lubricant. The effect of transverse surface roughness was studied by Deheri, Andharia and Patel [10] concerning the slider bearing with squeeze film based on a magnetic fluid. It was found that the effect of transverse surface roughness turned out to be considerably adverse. It has been sought to analyze the behavior of hydromagnetic squeeze films between conducting porous rough conical plates. The Fig. 1 describes the geometry and configuration of the bearing system. The lower plate with a porous facing is assumed to be fixed while the upper plate moves along its normal towards the lower plate. The plates are considered electrically conducting and the clearance space between them is filled by an electrically conducting lubricant. An external transverse uniform magnetic field B0 is applied between the plates resulting in the current to flow in the perpendicular direction. The flow in the porous medium obeys the modified form of Darcy\u2019s law (Ene [13]) while for the film region the equations of hydromagnetic lubrication theory hold" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002284_20080706-5-kr-1001.00797-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002284_20080706-5-kr-1001.00797-Figure2-1.png", "caption": "Fig. 2. Waypoints for straight and turning flight.", "texts": [ " Section 6 describes the design of the robust multivariable H\u221e controller and discusses robustness of the controller across the flight envelope. Simulation results are presented and discussed in section 7; section 8 concludes the paper. Fig. 1. A photograph of the experimental UAV. The mission of the UAV is planned in advance and loaded into the Flight Control Computer before flight. The mission plan consists of a number of waypoints which define the path the UAV is going to take. The entire mission can be represented by a series interconnection of straight line segments, arcs and circles (for loiter). Fig. 2 978-3-902661-00-5/08/$20.00 \u00a9 2008 IFAC 4737 10.3182/20080706-5-KR-1001.2734 shows on the left hand side part of a mission comprising of a straight line segment. In this case the UAV is required to fly through the waypoint WP2. The right side of Fig. 2 shows part of a mission with a turn. Here the UAV is not required to go through the central waypoint (WP2) but rather fly on a circular arc AB close to WP2. Point A is where the turn starts and B is where it ends. Similarly loiter missions are also possible where the vehicle flies in a circular orbit around a given waypoint. Here we shall assume that a mission has been planned and the mission data in terms of waypoints is available to the Flight Control Computer. We now formulate the problem in more specific terms", " Waypoints are defined as geographic positions in terms of latitude and longitude. For any three consecutive waypoints WP1, WP2 and WP3, if the turn angle (defined as the difference of azimuths of WP1 and WP3 at WP2) at the central waypoint (WP2) is nearly 180\u25e6, then the path from WP1 to WP2 will be considered a straight line, and WP2 will be referred to as a straight waypoint. On the other hand if the turn angle is significant, then WP2 will be referred to as a turning waypoint and the path from WP1 to WP3 will consist of straight and curved parts (Fig. 2). Flight path computations are worked out for both straight and curved paths; however only straight path formulae will be presented here due to space limitation. Let WP1(\u03c61, \u03bb1), WP2(\u03c62, \u03bb2) and WP3(\u03c63, \u03bb3) be three consecutive waypoints (\u03c6 and \u03bb denote the latitude and longitude, respectively), and let M(\u03c6M , \u03bbM ) be the current position of the vehicle. We define an Earth-Centred Earth-Fixed (ECEF) frame with its origin at the centre of mass of the earth, the z-axis directed north along the polar axis, the x-axis in the equatorial plane and passing through the Greenwich Meridian, and the y-axis also in the equatorial plane and passing through 90\u25e6 east longitude" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.128-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.128-1.png", "caption": "Fig. 14.128 High-pressure (airless) painting unit", "texts": [ "127 Scheme illustrating the operation of a highpressure (airless) painting unit by electric motors, pneumatic engines or combustion engines, with a power as high as 4.1 kW. A high-pressure painting unit is shown schematically in Fig. 14.127. The piston causes the pulsation of a membrane and the suction of paint from the reservoir through a filter and its forcing into the hose and spraying gun. A valve in the oil pump system is used to set the spraying pressure. A typical high-pressure (airless) painting unit design is shown in Fig. 14.128. Paint is sucked in from a reservoir by the paint pump and forced (at a maximum pressure of 17.5 MPa) into the spraying gun with a nozzle 0.43 mm in diameter. The rate of painting is about 4 m2/min. It should be noted that in paint work, besides the application of paint coatings, preparation of surfaces and paints plays a vital role. Industrial vacuum cleaners, sliding grinders, and corner grinders are used for preparing the base [14.40]. Paints are mixed by means of hand-operated and mechanical mixers and strained using screens to remove impurities" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002770_1.3609138-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002770_1.3609138-Figure4-1.png", "caption": "Fig. 4 Beams used to determine mode of contact", "texts": [ " Levy [3] disregarded any distribution of the reactive forces in his analysis of flat face flanges and assumed that line contact occurs where the slope of the beam is calculated to be zero under the imposed loads. It appeared reasonable to expect, however, that the position he assumed for line contact and the location of the centroid of distributed contact forces would be essentially the same. The validity of this proposition was established by means of several tests. In one test, two carbon steel beams were arranged as shown in Fig. 4. The surface of one beam was instrumented with 1/8-in. gage length electrical resistance strain gages. Load was increased in ncrements in a compression machine and strains were measured at each increment of load. The reaction at the shim, for each increment of applied load, and the position of the centroid of the reactive forces were determined from these data. The position of the centroid as determined by tests, designated bexp, and the position calculated according to Levy's hypothesis, designated bcaic, are compared in Table 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000683_00423118208968699-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000683_00423118208968699-Figure3-1.png", "caption": "Fig. 3. Yaw plane free body diagram of the bicycle model for directional response.", "texts": [ " a in the usual way in the traditional equations of motion. This model is a good approximation at low values of & [15]. A Model for TLB Directional Response A simple directional response model will be useful to investigate the effects of lateral compliance. In this model, the effects of pitch and roll are neglected and the forward velocity of the tractor is assumed constant. YAW PLANE FREE BODY DIAGRAM D ow nl oa de d by [ U ni ve rs ity o f W at er lo o] a t 1 2: 33 2 8 O ct ob er 2 01 5 DIRECTIONAL DYNAMICS OF A TRACTOR-LOADER-BACKHOE 257 Figure 3 presents a free-body diagram. The state variables and parameters are defined in the nomenclature listing. The equations of motion are: m (+ + ur) = C Fy (9) I,,; = C M, (10) The linear formulation for lateral forces yields Z F, = - Capr - (1 1) C = - Capf a + Cars!, b (12) where the compliant slip angle a! is related to the kinematic slip angle a!, by equation 8, and Thus the equations of motion are TLB Testing Several tests were conducted to provide parameters for the directional response model" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000413_robot.2003.1242208-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000413_robot.2003.1242208-Figure1-1.png", "caption": "Fig. 1: Metal spinning", "texts": [ " The author proposes applying hybrid position/force control for shear spinning, which is free from fine adjustment of the cleamnce. The effectiveness of the proposed method was experimentally verified. 1 Introduction This study seeks to exploit robot control techniques such as force feedback control for metal spinning. We aim to develop flexible and intelligent forming processes, and to expand a new application area for robot control. Metal spinning is a plasticity forming process that forms a metal sheet or tube by forcing the met,al onto a rotating mandrel using a roller or a paddle tool (Fig. 1). It is widely used for producing round hollow metal parts and products, e.g. tableware, kitchen- ware, ornaments, lighting fixtures, parabola antennas, boilers, tanks, gas canisters, nozzles, engine parts, and tire wheels. This forming process is also known as a highly-skilled manufacturing craft by artisans that requires decades of experience. Even nose cones for H2 space rockets launched by NASDA in Japan are produced by such manual metal spinning. Metal spinning has several merits over other metal forming processes as follows" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002815_1.4002527-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002815_1.4002527-Figure6-1.png", "caption": "Fig. 6 Magnitude of the disk inertia moment in the x direction", "texts": [ " After performing differentiations of the kinetic energy required y Lagrange equations of the second kind and after neglecting the mall terms of higher orders, the relationships for components of he inertia forces and moments Eqs. 8 \u2013 12 can be rewritten into he following forms: FSyT = \u2212 mDy\u0308T 14 FSzT = \u2212 mDz\u0308T 15 MS T = \u2212 JP\u0307 16 MS 1 = \u2212 JD\u03081 \u2212 JP\u03072 \u2212 \u0307JP 2 17 MS 2 = \u2212 JD\u03082 + JP\u03071 18 he generalized coordinates 1, 2, and T and corresponding omponents of the inertia moment acting on the disk are referred o axes y , , and , which are not orthogonal Fig. 4 . Projections f the component of inertia moment MS 1 T acting in plane x y or into axes y , , and x are denoted MS 1, MS T, and MSx , espectively Fig. 5 . Using Fig. 6, it holds ig. 5 Decomposition of the inertia moment into the y , , and directions 21001-4 / Vol. 78, MARCH 2011 om: http://appliedmechanics.asmedigitalcollection.asme.org/ on 01/28/201 MSx = MS T \u2212 MS 1 \u2212 MS T 2 2 19 As evident from Fig. 4, y and z components of the disk inertia moment can be expressed as MSy = MS 1 20 MSz = MS 2 \u2212 MSx 1 21 After substitution of Eq. 19 into Eq. 21 and taking into account small rotations, one obtains MSz = MS 2 \u2212 MS T 1 22 From the physical point of view QTy, QTz, Q 1, and Q 2 represent the applied and constraint forces and moments acting on the liberated disk" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001923_j.mechatronics.2008.11.006-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001923_j.mechatronics.2008.11.006-Figure3-1.png", "caption": "Fig. 3. Sliding force~S computed from the perception of ~R vector.", "texts": [ " Consequently, the intensity of each signal changes dynamically as the robot is moved along. We detail in Section 4.3.1 the computation of the signal intensities used in the experiments. Fig. 2 shows a flowchart of the paralyzed agent\u2019s behavior. 3.2. The pusher robots: a force field approach These robots explore the environment until they perceive attractive signals emitted by the paralyzed one. Their behavior then consists of moving towards the origin of the signal, whose direction can be easily computed from the set of sensors that encircle the robot\u2019s body (see Fig. 3). As a consequence, they will arrive near the carriage and may collide with it in order to apply a pushing force. However, this task needs to be precise so that the agents are placed correctly and push against the arms of the carriage (see Fig. 1). For this purpose two simple reactive behaviors are defined: (i) attraction towards the carriage by following the signal (see Fig. 4a), (ii) sliding along the carriage sides (see Fig. 4b). The objective is to point the arms in a direction so they can be pushed (see Fig", " \u2013 An avoiding force from other mobile robots in proximity is integrated only outside the sliding area. Three force vectors can be derived from these perceptions: \u2013 An attraction force towards the signal source~A \u00bc pp0 ! . The intensity of the signal reception allows to approximate the norm kpp0 ! k. Note that the attraction force ~A decreases with the distance to p0 and is bounded to a maximum value. k~Ak \u00bc max\u00f0k ~pp0k;Amax\u00de. \u2013 A sliding force~S along the carriage such that~S:~R \u00bc 0, where ~R is the vector from the closest point of the carriage to the robot (see Fig. 3). The intensity of k~Sk \u00bc Rmax k~Rk, which increases when the robot goes near the carriage. The direction of ~S is given by the perceived signal. Considering~R as reference, when the robot is inside the left signal, ~S is oriented on the right-hand side, or else on the left-hand side (see Fig. 4b). \u2013 A force ~D to avoid collisions with other nearby robots. A sliding force is computed for each of them. ~D is a weighted sum of these vectors as a function of their respective distance, see details in [15]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002271_s11434-008-0370-x-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002271_s11434-008-0370-x-Figure4-1.png", "caption": "Figure 4 The piezoelectric high-speed on/off valve. (a) Structure; (b) PZT actuator from PI.", "texts": [ " If there is no temperature compensation for the PZT actuator, when the PZT actuators are powered off, some leakages will happen because the expanded length of the PZT actuator lifts the poppet a small distance. Considering the temperature expansion coefficient of the PZT actuator[9], a piece of special metal is selected as the compensation piece. Because the function of the PZT actuator 2 is to strike the poppet close, temperature rising does not affect its task, and therefore only the PZT actuator 1 needs temperature compensation piece. The structure of the PZT valve is shown in Figure 4. The PZT actuators are selected from PI Company, and their parameters are shown in Table 1. The piezoelectric actuators 1 and 2 are driven respectively by a pulse DC of 120 V with a period of 2 ms, and some basic parameters are given in Table 2. The jet angle of the poppet \u03b1 varies from 45\u00b0 to 65\u00b0 by a step of 10\u00b0 and other parameters are not changeable. OUYANG XiaoPing et al. Chinese Science Bulletin | September 2008 | vol. 53 | no. 17 | 2706-2711 2709 A R TI C LE S Table 1 Parameters of the PI PZT actuator[10] PZT actuator Size A\u00d7B\u00d7L (mm\u00d7mm\u00d7 mm) Displacement (\u03bcm) Max output force (N) Stiffness (N/\u03bcm) Resonant frequency (kHz) 1 7\u00d77\u00d736 32 1850 50 40 2 10\u00d710\u00d736 32 3800 100 40 Table 2 Basic parameter of the PZT valve The pressure p2 and flowrate qv,2 of the digital valve are shown in Figure 5" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003015_oceanssyd.2010.5603906-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003015_oceanssyd.2010.5603906-Figure2-1.png", "caption": "Fig. 2. Auto-Ballasting System (ABS). The ABS consists of a motor driven piston mechanism that either draws water into the piston tube, or flushes water out, changing the float\u2019s buoyancy.", "texts": [ " One mission could be primarily based on near\u2013bottom work, while the float maintains a constant altitude. Another mission could consist primarily of rate controlled profiling with CTD readings between two user\u2013defined depths. These new mid\u2013depth Lagrangian profiling floats are designed to be a very versatile platforms that can support many oceanography studies. To control buoyancy and therefore vertical movement, the float is equipped with an Auto\u2013Ballasting System (ABS) that can drive water into or out of a piston tube within the float (Fig. 2). The volume of the piston tube with respect to the float is quite large so that the ABS can maintain performance in dynamic environments often associated with shallow waters. Essentially, by adjusting the volume of water within the piston tube, the effective volume of the float changes, thereby changing the buoyancy. The total active volume (Va) for buoyancy control is a function of the piston stroke (h) and plunger diameter (dp) and area. This large volumetric change is required so that the float can handle dynamic environments and also so that the float can properly lift its iridium antenna high enough out of the water", " The system dynamics provide a foundation for specifying all system components in order to ensure the Lagrangian float\u2019s physics are matched with the onboard Auto-Ballasting System (ABS). The main forces acting on the float include buoyancy, gravity, and drag as shown in Fig. 4. Added mass, which affects accelerating objects underwater, is also considered. This term is not actually derived, but is found empirically. The equation of motion considers these factors, and the influence of the ABS, which uses a piston to pull water into or out of the system, as illustrated in Fig. 2. By changing the volume of water contained within the float, the float\u2019s volume actually remains the same. Realistically, the float\u2019s mass is the only variable changing because the water pulled into the float must move with the float as well. Intuitively, it is more helpful to think of the ABS as a volume changing mechanism. It will be shown through the following derivation that this assumption is actually quite valid when the ABS volume is small compared to the float\u2019s volume. Additionally, this assumption is critical to a more simplified control system design", " The SM2315DT power consumption is shown as a function of ABS flow rate and applied pressure in Fig. 6. A similar model was developed for the SM2316DT\u2013PLS2 motor as well. The efficiency is then found by comparing the power consumed with the theoretical mechanical power required to flush water at particular pressures and velocities. The mechanical power required to drive water at a given pressure with a constant velocity is a product of the constant torque the motor is subject to and the angular velocity at which it is driven. Referring back to the ABS in Fig. 2, L is the leadscrew lead, p is the applied pressure, Ap is the area of the piston, Ffs is the force of friction due to the seal, els is the leadscrew efficiency, egb is the gearbox efficiency, Nr is the gearbox ratio, and the constant torque acting on the motor, \u03c4c is: \u03c4c = L (pAp + Ffs) 2\u03c0elsegbNr (6) The mechanical power (Pm) required from the motor is then a product of the constant torque and the motor\u2019s angular velocity (\u03c9): Pm = \u03c4c\u03c9 (7) Next, the power results in Fig. 6 are divided by the theoretical values found in (7)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000283_0020-7403(85)90082-7-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000283_0020-7403(85)90082-7-Figure7-1.png", "caption": "FIG. 7. Continuous strut including axial inertia.", "texts": [ " We now have a kinetic energy expression which is also a function of displacement T = \u00bdML zcos zQQ2 (41) the potential energy remaining unchanged. Application of the general theory leads to A = Q/sin Q (42) and W-' = (1 - A cos Q)/cos \"~ Q (43) which is represented graphically in Fig. 6(b). It can be seen that the first-order solution is AA = QZ/6 144) and W 2 ~ Q2/3 (45) which gives A A / W ~ = \u00bd, (46) where terms greater than quadratic have been neglected. Therefore axial inertia has no effect on the initial slope but may be significant for highly buckled systems. (iii) Continuous systems Consider the strut shown in Fig. 7 with distributed mass m per unit length. A particte on the strut moves from position A to B, described by u and w in the x and y directions, respectively. The velocity of this displacement is v z = ,~z + ~2. (47) The kinetic energy of the particle is 6)\" = \u00bd(max)(,': +,~2) and adding these up we get where o si2dx dx, d O u =- w '2 dx. If we use a modal solution of the form w - Q M ( x ) then w' = QM'(x ) and w,2 = Q2 M,2(x). Therefore, and (48) (49) (5o) (51) (52) '~ = Q0 f~ M '~ dx (53) so that we can see the axial inertia of the distributed mass is of second order" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002012_cans.2008.13-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002012_cans.2008.13-Figure1-1.png", "caption": "Figure 1. Workspace of a leg Figure 2. Workspace of mobile platform", "texts": [ " The in-out equations of the 6-PGK parallel robot are in the form ,0KZKYKXKZYX oipzipyipxi 2 p 2 p 2 p =++++++ ( )6,1i = (1) Substituting dZp = , these are in the form 0KXKX oipxi 2 p =\u2032++ , ( )6,1i = (2) which is a system of 6 equations with 6 unknowns, from which cross-sections of the workspace with horizontal planes is obtained. In paper [6] are developed in complete form the coefficients from equations (1) and (2). While the 6-PGK parallel robot is at work, points iC ( )6,1i = attached at the mobile platform are moving on spherical surfaces of radius iR (Figure 1). The centres of the inferior and superior spheres are points i iA , respectively s iA , that are situated on straight line ic ( )6,1i = . The generalized coordinates of the robot in these configurations are m iq and M iq . Let is and iS be the inferior, respectively the superior positions of the two spheres. In conclusion, the workspace of point iC is a volume bounded superior by the sphere iS and inferior by the sphere is . When the manipulated object has a constant orientation, its trajectory is situated in the plane [ ]pQ (Figure 2)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000139_0020-7403(86)90032-9-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000139_0020-7403(86)90032-9-Figure2-1.png", "caption": "FIG. 2(a) Position of centre of mass of spinning sphere relative to centre and direction of motion. (b) Deviation as measured using screens and firing from (c) a musket with a deliberately bent", "texts": [ " T h e Civ i l o r L e g a l d a t e 6 F e b r u a r y 1671, w a s the s a m e as t he H i s t o r i c a l d a t e 6 F e b r u a r y 1672. A d a t e s u c h as 6 F e b r u a r y , b e t w e e n 1 J a n u a r y a n d 25 M a r c h w a s t h u s wr i t t en , 6 F e b r u a r y 167 ~. barrel [la]. center of pressure acting through the geometric center and the inertial reaction acting through the center of mass. When fired, the ball should drift to the right--it did'. (The latter quotation may be helpfully represented as in Fig 2(a).) Below, the contents of the first section of Magnus' first paper [2a and b]--which is the significant portion for our purposes--are described in some detail and it is rather different from that just quoted. Robins does not of course refer to 'unsymmetrical pressure distributions produced by the Bernoulli effect': ideas about the joint effect of velocity and pressure were then just being elucidated by D'Alembert and later Euler; see Note 2, in the Appendix. That the association of Magnus' name with this effect is widespread is soon found after consulting standard English classical text and reference books", " Also, briefly, he showed subsequently in experiments that when he fired a 'piece from some fixed notch' so that the projectile might pass through successive screens of thin paper they were deflected from the vertical plane of flight. In another test, Robins arranged paper screens at fixed distances from a wall and fired several (two lots of five) balls (17 to the lb, in the first set of tests, windage + being extremely small) with \u00bc oz of good powder. Typically the lateral deviations of one test may be accepted as representative, see Fig. 2(b), and he demonstrates incurvation from the original plane of flight. And finally and conclusively, Robins took 'a barrel . . . and bent it at about three or four inches from the muzzle to the left, the bend making an angle of 3 \u00b0 or 4 ~ with the axis of the piece. This piece thus bent was fired with a loose ball . . . . It was natural to expect that if this piece was pointed in the general direction of its axis the ball would be canted to the left of that direction by the bend at its mouth. But the bullet ", " that the bullet might be expected to incurvate towards the r i g h t . . , and this upon trial did most remarkably h a p p e n . . , it deviated to the right and a little beyond the second screen crossed that track from which it before diverged' (by having been canted) \"and on the wall was deflected 14 inches, as I remember, on the contrary side . . . . (it) i s . . . the strongest confirmation that it is brought about in the very m a n n e r . . , which we have all along d e s c r i b e d . . . \", see Fig. 2(c). * There may be evidence that this innovatory piece of measuring equipment was in fact introduced by one of the Cassinis in 1707. Unreferenced statement on p. 103 in Ref. [20]. * Windage is the clearance between the projectile (ball) and the internal diameter of the gun barrel; for 32pr. guns of diameter 6.375 in. this was 0.198 in [19, p. 584]. The Magnus Effect and priority 865 Robins' clear appreciation of the contributions of the velocities causing the deviations lead him to venture to say, 'I think I am so much master of this subject as to undertake refutation of whatever objections shall be hereafter started on this head'" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001410_0301-679x(77)90021-4-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001410_0301-679x(77)90021-4-Figure10-1.png", "caption": "Fig 10 (below) No wear part o f the f lank around the tooth trace", "texts": [ " However, examinations of surface replicas show that the flanks have been slightly worn: the amount of wear increases with increasing pressure (Fig 8). The amount of wear of the gears run at 2 000 and 4 000 rpm respectively was approximately equal. However, a test series was made to check the influence of running speed on the initial wear. The results obtained show a slight increase of wear with increasing running speed (Fig 9). In all cases studied, both with as hobbed and shaved tooth flanks, there was no wear around the tooth trace of the flanks (Fig 10). The tooth trace is the line of intersection between the tooth flank and the reference cylinder. The no-wear widths were about 1 and 2 mm for the tests of as hobbed gear teeth at 4 000 and 2 000 rpm respectively and about 3 mm for the shaved gear teeth. In none of the test series was there any noticeable difference in the amount of wear between the gears made of SIS 2216 or SIS 2240. Discussion The results obtained showed that the changes of surface roughness are greater with increasing speeds despite the increase in theoretical film thickness with increasing speed" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000202_s0218127404010874-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000202_s0218127404010874-Figure1-1.png", "caption": "Fig. 1. Existence of the Poincare\u0301 map \u03a0A ++ in a neighborhood of the contact point p+ when n = 2.", "texts": [ " Let \u03a0+\u2212 be the Poincare\u0301 map which transforms point of LI + into points of LO \u2212, and \u03a0\u2212+ the Poincare\u0301 map which transforms points of LI \u2212 into points of LO +. Since both maps are defined by the flow of the linear system x\u0307 = Bx, we refer to them by \u03a0B +\u2212 and \u03a0B \u2212+. Proof of Theorem 1.1. (a) The statement follows immediately from Lemmas 3.2 and 3.5(b). (b) From Lemma 3.5(b), there exists exactly one contact point p+ \u2208 L+ of order n\u22121. Hence, the orbit \u03b3p+ through p+ satisfies the following local behavior. If n is even, from Proposition 2.3(a), then \u03b3p+ does not cross the hyperplane L+, see Fig. 1. We can consider a tubular neighborhood U of \u03b3p+ contained in a flux box surrounding a piece of \u03b3p+ in a neighborhood of p+. According to the Continuous Dependence Theorem of the solutions of a differential equation with respect to the initial conditions, U intersects LI + and LO +. Take q1 \u2208 LO + \u2229 U . The In t. J. B if ur ca tio n C ha os 2 00 4. 14 :2 84 3- 28 51 . D ow nl oa de d fr om w w w .w or ld sc ie nt if ic .c om by W E ST V IR G IN IA U N IV E R SI T Y o n 11 /0 5/ 14 . F or p er so na l u se o nl y" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003908_978-94-007-4201-7_1-Figure1.9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003908_978-94-007-4201-7_1-Figure1.9-1.png", "caption": "Fig. 1.9 Force acting on a body", "texts": [ " That is, oi $i \u00bc \u00f0o1S1;o1S01\u00de \u00fe \u00f00; v1S1\u00de \u00bc \u00f0o1S1;o1S01 \u00fe ho1S1\u00de \u00bc o1 S1; S 0 1 ; (1.52) or oi $i \u00bc o1 S1; S 0 1 \u00bc \u00f0v1; v 0\u00de; (1.53) where v1 is the angular velocity of the body and v0 is the velocity of a point in the body coincident with the origin. Note that the directions of v0 andv1 are different in general, unless the axis of the screw passes the origin. The pitch is h \u00bc S1 S01 S1 S1 \u00bc v1 v0 v1 v1 : (1.54) Analogous to instantaneous rotation, unit line vectors can be used to express the action of a force on a body, as shown in Fig. 1.9. A force f can be expressed as a scalar multiple fS of the unit vector S bound to the line. The moment of the force C0 about a reference point O can be expressed as a scalar multiple fS0 of the moment vector S0 \u00bc r S . The action of the force upon the body can thus be elegantly expressed as a scalar multiple f$ of the unit line vector where $ is unit line vector, S S \u00bc 1 and S S0 \u00bc 0. The line vector can be used to express the magnitude, direction and the acting line in space. The Pl\u20acucker coordinates of the force line vector are f \u00f0S; S0\u00de, \u00f0fS; fS0\u00de or \u00f0f ;C0\u00de, where C0 is the moment of force f about the origin, that is, C0 \u00bc fS0 \u00bc f r S0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000518_imc.1990.687271-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000518_imc.1990.687271-Figure4-1.png", "caption": "Fig. 4 Mathematical model of two dimensional manipulator", "texts": [ " We adopted the membership function, like Fig. 3, of the error and of the sum of weight changes, and the rule map, like Table 1 , experimentally. 0 'DoOo( 1 Fig. 3 Membership function In this paper, in order to investigate the ability of the proposed Time Delay Neural Network as a servo controller, simulations are carried out in case of the position/force (mainly force) control of two dimensional robotic maniplatar which handles unkncrwn objects and so vihich has the strong nonlinearlity and the canplex dynamics. Figure 4 shows the mathematical model of two dimensional robotic maniplator. The handling object is supposed to be modeled as a linear spring, a linear damper and a linear mass. Figure 5 shows the neural servo control system of two di\"iona1 robtic manipulator. The PID control is applied for the position control ard the PI oontrol for the force control. The PID/PI control is done as the Hykid cmtrol employing the hybrid ratio sl. ?he neural network which is applied to the hybrid control consists of two hidden layers and input/output layers" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001872_physreve.76.051905-Figure14-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001872_physreve.76.051905-Figure14-1.png", "caption": "FIG. 14. Initial state of a symmetric microtubule pair: force acts on the microtubules perpendicular to the bisector; a attachment point is equidistant from the centers of mass; b inflexible tubules connected by a finite-size motor; c a vector in an orthonormal frame.", "texts": [ " The filaments are assumed to be of the same length, with the motor attached at an equal distance from the minus end of each filament even attachment and oriented transversally to the bisector of the filament pair symmetric attachment . The general idea of symmetrization of motor attachment is captured in the case of rigid tubules connected by a stiff but flexible motor of length h, where h L. The motion then reduces to a system of ordinary differential equations governing the overdamped motion of a system of two rigid rods microtubules connected by flexible inextensible link of the length h motor as depicted in Fig. 14. In a fixed coordinate system the centers of mass of the tubules are at c 1 and c 2 , respectively, while the motor attachment points are at c 1 + t 1 and c 2 + t 2 . In terms of the distances from the respective centers s1,2 we have t 1,2 =s1,2t\u0302 1,2 . The motor is represented by the vector ; n 1 , n 2 , and are normal to t 1 , t 2 , and , respectively; and hats denote unit vectors. The dynamics of the system are determined as in the semiflexible case by the balance of forces and the kinematic constraint d dt s1,2=v, d dt c 1,2 = \u00b1 f \u22121 t\u0302 1,2 , \u0302 t\u0302 1,2 + \u22121 n\u0302 1,2 , \u0302 n\u0302 1,2 , d dt t 1,2 = vt\u0302 1,2 \u00b1 f r \u22121s1,2 2 t\u0302 1,2 , \u0302 n\u0302 1,2 , A1 where \u22121, \u22121, and r \u22121 denote the inverses of tangential, transversal, and rotational viscosities; the motor force of magnitude f is directed along \u0302, the sign depending on the direction from the tubule. To obtain the force f and the motor motion, we observe the geometric constraint = c 2 \u2212c 1 + t 2 \u2212 t 1 Fig. 14 . Since the motor is rigid, its velocity must be directed along the normal \u0302. Since the motor is rigid, its velocity is directed along the normal: \u0307=h\u0307\u0302. Here \u0307 denotes the angular velocity, which is independent of the choice of an orthonormal coordinate system. We obtain the relations h\u0307\u0302 = \u2212 f \u22121 t\u0302 1 , \u0302 t\u0302 1 + t\u0302 2 , \u0302 t\u0302 2 \u2212 f \u22121 n\u0302 1 , \u0302 n\u0302 1 + n\u0302 2 , \u0302 n\u0302 2 + v t\u0302 2 \u2212 t\u0302 1 \u2212 f r \u22121 s1 2 t\u0302 1 , \u0302 n\u0302 1 + s2 2 t\u0302 2 , \u0302 n\u0302 2 . A2 Denote by vu\u0302= v , u\u0302 u\u0302 the component of a vector v along the unit vector u\u0302", " 173, 733 2006 . 24 Expression 20 is in fact similar to the critical height hc of self-buckling condition of vertical column of density and radius r, hc= 2.5Er2 / g 1/3, where g is gravity acceleration and E Young\u2019s modulus. 25 D. M. Smith, F. Ziebert, D. Humphrey, C. Duggan, M. Steinbeck, W. Zimmermann, and J. A. K\u00e4s, Biophys. J. to be published . 26 In two dimensions a vector product of a vector with \u0302 is a scalar and can be replaced by a scalar product of the normal with \u0302 see illustration on the right of Fig. 14 . 27 We have t\u0302 1 + t\u0302 2 = t\u0302 t\u0302 2 1 + t\u0302 n\u0302 2 1 + t\u0302 t\u0302 1 2 + t\u0302 n\u0302 1 2 = t\u0302 n\u0302 2 1 + t\u0302 n\u0302 1 2 + t 1 , t 2 t\u0302 1 + t\u0302 2 . 28 G. Strang and G. Fix, An Analysis of the Finite Element Method Prentice-Hall, New York, 1973 . 051905-12" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002288_j.mechmachtheory.2008.03.007-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002288_j.mechmachtheory.2008.03.007-Figure3-1.png", "caption": "Fig. 3. Generation with two fixed axes. Case with S \u00bc T\u00f00\u00de \u00bc G\u00f00\u00de.", "texts": [ " Application 1: generation with fixed axes In order to grasp the main features of the general formulation previously presented, we start the application sections by employing it to the study of the simple case of generation between two fixed axes. The layout considered refers to the generation of a hypoid pinion on a Gleason cradle style machine without supplemental motions. Therefore, the relative motion between the tool surface Rt and the to be generated gear surface Cg consists only of two rotations about fixed axes, as illustrated in Fig. 3. 6.1. Formal definition of some fundamental quantities As customary, let us introduce in the Euclidean space E3 a fixed frame S, and two moving frames T and G, stationary with respect to the tool and the gear blank, respectively. Then, let nt1 and ng1 be the (unit) twists associated with the rotations, with twist components given by 5 The nt1 \u00bc x\u0302t1 qt1 xt1 ; ng1 \u00bc x\u0302g1 qg1 xg1 ; \u00f094\u00de where xt1 and xg1 are the unit vectors of the axes, and qt1 and qg1 are the coordinates of any point on each axis", " It is worth observing that, till now, no explicit choice of the position and orientation of S has been made, and therefore the components in all the expressions are still floating values. Moreover, from (104), beside the rotational parts of the twists xt1 and xg1 , the role played by the two vectors (difference of point coordinates) \u00f0pt\u00f0u; v\u00de qt1 \u00de and \u00f0qg1 qt1 \u00de is automatically highlighted, no matter where the actual S will be set. 6.2. Example calculations with S \u00bc T\u00f00\u00de \u00bc G\u00f00\u00de and Og \u00bc qg1 To illustrate further the concepts presented above, explicit calculations of some interesting quantities are reported in the case of Fig. 3 where the three frames S, T and G coincide in the initial condition, that is, S \u00bc T\u00f00\u00de \u00bc G\u00f00\u00de, or equivalently, gst\u00f00\u00de \u00bc gsg\u00f00\u00de \u00bc ggt\u00f00\u00de \u00bc I. Moreover, we let the origin Og of G be coincident with the chosen point on the gear blank axis qg1 . The unit vectors xt1 and xg1 , and the points qt1 and qg1 chosen on the tool and gear axes, respectively, have components xt1 \u00bc \u00f00;0;1\u00de; qt1 \u00bc \u00f0 D0 cos c0; E0; B0 D0 sin c0\u00de; \u00f0105\u00de xg1 \u00bc \u00f0cos c0;0; sin c0\u00de; qg1 \u00bc \u00f00;0;0\u00de: \u00f0106\u00de The coordinates of the twists nt1 and ng1 are therefore nt1 \u00bc \u00f0E0;D0 cos c0;0;0;0;1\u00de; \u00f0107\u00de ng1 \u00bc \u00f00;0;0; cos c0;0; sin c0\u00de: \u00f0108\u00de The components n1 and n2 are simply n1 \u00bc ng1 and n2 \u00bc nt1 , since gsg\u00f00\u00de \u00bc I, and the transformation matrix gsg in (95) is given by the product of the two following matrices: en\u03021h1 \u00bc 1\u00fe \u00f0cos h1 1\u00de sin2 c0 sin c0 sin h1 sin c0 cos c0\u00f0cos h1 1\u00de 0 sin c0 sin h1 cos h1 cos c0 sin h1 0 sin c0 cos c0\u00f0cos h1 1\u00de cos c0 sin h1 1\u00fe \u00f0cos h1 1\u00de cos2 c0 0 0 0 0 1 266664 377775; \u00f0109\u00de en\u03022h2 \u00bc cos h2 sin h2 0 D0 cos c0\u00f0cos h2 1\u00de \u00fe E0 sin h2 sin h2 cos h2 0 E0\u00f0cos h2 1\u00de \u00fe D0 cos c0 sin h2 0 0 1 0 0 0 0 1 26664 37775: \u00f0110\u00de The proximal and distal Jacobians Jg gt and Jt gt , respectively, have the following expressions: Jg gt\u00f0h1\u00de \u00bc 0 E0 \u00fe E0\u00f0cos h1 1\u00de sin2 c0 \u00fe D0 cos c0 sin c0 sin h1 0 D0 cos c0 cos h1 E0 sin c0 sin h1 0 cos c0\u00f0E0\u00f0cos h1 1\u00de sin c0 \u00fe D0 cos c0 sin h1\u00de cos c0 cos c0\u00f0cos h1 1\u00de sin c0 0 cos c0 sin h1 sin c0 1\u00fe cos2 c0\u00f0cos h1 1\u00de 26666666664 37777777775 ; \u00f0111\u00de Jt gt\u00f0h2\u00de \u00bc sin c0\u00f0E0\u00f0cos h2 1\u00de \u00fe D0 cos c0 sin h2\u00de E0 sin c0\u00f0D0 cos c0\u00f0cos h2 1\u00de E0 sin h2\u00de D0 cos c0 cos c0\u00f0E0\u00f0cos h2 1\u00de D0 cos c0 sin h2\u00de 0 cos c0 cos h2 0 cos c0 sin h2 0 sin c0 1 2666666664 3777777775 : \u00f0112\u00de The computation of the remaining expressions such as, e" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001378_ijmmm.2007.015474-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001378_ijmmm.2007.015474-Figure5-1.png", "caption": "Figure 5 Model proposed by Astakhov to study the thermal phenomenon in metal cutting", "texts": [ " The equivalent cutting geometry together with other input parameters such as cutting speed, feed, physical and mechanical properties of the work and tool materials, geometry of the workpiece are used to calculate the relevant similarity numbers as described in Astakhov (1998, 2004). Then analytical models for the orthogonal cutting are employed to predict the forces, temperatures and thermal energy distribution in the deformation zone. These models were presented by Astakhov (1998, 2004) and described as follows: Figure 5 shows the model presented by Astakhov (1998, 2004) to calculate the temperatures and thermal energy distribution in the deformation zone. Since practically all of the mechanical energy associated with chip formation is converted into the thermal energy, the heat balance equation is of prime concern in metal cutting studies (Astakhov, 1998). This equation can be written as: C C c w tF v Q Q Q Q\u03a3= = + + (4) where Fc is the cutting force, N; vc is the cutting speed, m/s; Q\u03a3 is the total thermal energy generated in the cutting process, J/s; Qc is the thermal energy transported by the chip, J/s; Qw is the thermal energy conducted into the workpiece, J/s; Qt is the thermal energy conducted into the tool, J/s", "5 Pe Pe Pe 1 sin Pe 1 sin cos sin cos sin cos n n n n n n n n B B B B B B FD B F D B B B B B ( ) ( )\u03b3 \u03b3 \u2212 \u23a1 \u23a4+ \u2212 +\u23a2 \u23a5\u23a3 \u23a6 0.4eq eq eq 0.3 1 sin Pe 1 sin 0.225n nB FD (13) Concerning to the maximum temperature at chip formation zone, and at tool-chip and tool-workpiece interfaces, the following equations were derived by Silin and presented in the corrected form by Astakhov (1998, 2004): \u2022 The maximum temperature reached in the plane of maximum combined stress (also known as shear plane), TA, is supposed to be reached close to the tool cutting edge (see point A in Figure 5) and given by the following equation: ( ) Pe erf (11) 4 f A p w B T c \u03c4 \u03c1 = (14) where cp e\u03c1 are the specific heat (J/Kg\u00b0C) and the density (Kg/m3) of the work material, respectively. \u2022 The maximum contact temperature at the tool-chip interface, TM, (the cutting temperature) is supposed to be reached at the middle of tool-chip contact (see point M in Figure 5) and is given by the following equation: ( )1M A MT T \u03c8= + (15) where ( )( ) ( )eq eq eq eq 0 eq eq cos sin cos sin0.9675 1/ 1 Pe cos sinerf Pe / 4 n n n n M n n Bb B BB \u03b3 \u03b3 \u03b3 \u03b3 \u03c8 \u03b3 \u03b3 + \u2212 \u2212+ = + (16) where ( ) 0.3 eq eq 0 0.2 0.3 eq eq eq eq 0.25 cos sin Pe cos sin cos sin n n n n n n FD B b B B \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 + = \u23a1 \u23a4+ \u2212 \u2212\u23a3 \u23a6 (17) \u2022 The maximum contact temperature at the tool-workpiece interface, TN, is supposed to be reached at the middle of tool-workpiece contact (see point N in Figure 5) and is given by the following equation: 0.25 eq 1.25 0.36sin 0.5 0.53 Pe n N A NT T B E \u03b1 \u03c8 \u239b \u239e = + +\u239c \u239f \u239d \u23a0 (18) where ( )( ) \u03b1 \u03c8 \u03b1 + = 1.25 eq 1 0.25 eq 0.6 1 Pe cos sin erf Pe / 4 n N n p B E B (19) and 0.3 0.1 eq 1 0.2 0.1 0.24 sin Pe nFD p E B \u03b1 = (20) Figure 6 shows simplified diagram of the forces generated in orthogonal cutting, as proposed by Astakhov. In this diagram R is the resultant cutting force, F\u03c4 the force along the plane of the maximum combined stress, FC and Ft are the power and trust components of the cutting force, respectively and Fn\u03b3eFt\u03b3 are the normal and the friction forces on the tool rake face" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure5-1.png", "caption": "Fig. 5. Common frame.", "texts": [ " The crossing angle between the tangent of the directrix at some point and the generatrix of the pitch cylinder or the pitch cone passing this point is named as the spiral angle b, and generally it is the function of the point at the directrix, whereas the absolute values of the spiral angles of a pair of conjugate directrixes at conjugate point must be equal. To the outer gearing, the hands (left-hand or right-hand) of two meshing gears are opposite, so the signs (\u2018\u2018+\u2019\u2019 or \u2018\u2018 \u2019\u2019) of the spiral angles of conjugate directrixes are also opposite; to inner gearing, the hands of two meshing gears are same, so the signs of the spiral angles of conjugate directrixes are also same. (3) In Fig. 5, point P is a conjugate point at line of action (instantaneous axis), {P,a1a2a3} is a common frame of two directrixes, wherein a1 is unit vector of common tangent of the directrixes, a3 is unit common normal vector of two datum surfaces, a2 is in common tangent plane and forms the orthogonal right-hand coordinates. For paralleled-axis drive (R = 0,p), then there is q = \u00b1AIa. In Fig. 5(a), a3 is parallel to unit vector a, so there exists q2 \u00bc AIa \u00f0N\u00f01\u00de v\u00f021\u00de p \u00de \u00bc 0, noninterference condition is satisfied naturally; for intersected-axis drive (A = 0), then there is q = IsinR(a \u00b7 P(0)). In Fig. 5(b), a3 \u00c6 a = a3 \u00c6 P(0) = 0, so a3 is parallel to a \u00b7 P(0), there exists q2 \u00bc I sin R\u00f0a P\u00f00\u00de\u00de \u00f0N\u00f01\u00de v\u00f021\u00de p \u00de \u00bc 0, noninterference condition is also satisfied naturally. Paralleled-axis drive is applied widely in the engineering, and we can acquire basic properties of this type of drive from general principles of normal circular-arc gear drive concluded above. In Fig. 6, R\u00f01\u00dep and R\u00f02\u00dep are a pair of pitch cylinders, IA is instantaneous axis (line of action), R1 and R2 are the radii of two pitch cylinders" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001544_iros.2007.4399373-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001544_iros.2007.4399373-Figure2-1.png", "caption": "Fig. 2. Simulated 3DOF arm", "texts": [ " If an accurate inverse model is used, the composite controller can achieve compliant, fast and accurate movement. One effect of the composite control approach is that the more accurate the inverse model g, the smaller are the errors and the error-correcting PD signals. Thus, the total amount of feedback control is a measure of the accuracy of the inverse predictive model. The ability of LWPR to learn non-linear dynamics that can be used for control was verified. A simulated 1 3 DoF arm was used (see Fig. 2). The first joint allows up and down movements and the next two allow left and right movements. The task of the arm was to follow a smooth trajectory planned in joint angle space. The trajectory was a superposition of different phase-shifted sinusoidal trajectories for each joint. Twenty iterations of the trajectory were 1Simulations performed using ODE and OpenGL executed. The arm was controlled by a PD controller in the first 3 iterations of the trajectory and then switched to a composite controller using the model being learned" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000632_s00419-005-0435-0-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000632_s00419-005-0435-0-Figure1-1.png", "caption": "Fig. 1 Small dynamic deformation superimposed on large static deformation", "texts": [ " LSVM was derived through linearization of Simo\u2019s model and reference transformation. The SDIF was introduced empirically to take into account the influence of prestrain on the relaxation function. It has been observed that the K\u2013Y model agrees well with the results of dynamic compression tests [7]. In this section, we review the K\u2013Y model. Then the modified K\u2013Y model considering the Payne effect will be proposed. 2.1 Notation A small deformation superimposed on a large static deformation is depicted in Fig. 1. Let \u03be denote the configuration of the body B at instant \u03be . The t \u2032 , t0 , and t refer respectively to the undeformed, the statically deformed, and the current configurations. The \u03be T(\u03b7) represents a kinematical tensor T at time \u03b7 with respect to configuration \u03be . For convenience the following simplified notations are also used : T(\u03b7) = t \u2032 T(\u03b7), \u03be T0 = \u03be T(t0), 0T(\u03be) = t0 T(\u03be), \u03be T = \u03be T(t). The deformation gradient tensor and the volume-preserving deformation gradient tensor are respectively denoted by ( \u03be F(t) ) ij = ( \u2202xi(t) \u2202Xj (\u03be) ) and \u03be F\u0304 = J\u22121/3 \u03be F, where J is det ( \u03be F ) " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000752_00207720600798320-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000752_00207720600798320-Figure5-1.png", "caption": "Figure 5. Fleet of eight identical Trainer 60 aircraft used in the multi-UAV testbed at MIT.", "texts": [ " The 7 ft diameter blimps used in the testbed provide a simple means to include vehicles with different dynamics and capabilities. The blimps were designed to perform tasks in conjunction with the rovers, such as rapidly performing reconnaissance or classification beyond the obstacles blocking the rovers. They were scaled to carry the VAIO and have an identical control architecture. While these vehicles are more complicated to operate than the rovers, they are far easier to fly indoors than fixed or rotary wing vehicles. 3.2 Tower Trainer 60 UAV Figure 5 shows the second testbed, which started as a fleet of eight UAVs. In order to make successful demonstrations of multi-vehicle flights, the logistics require that all the vehicles have adequate minimum flight durations to ensure that there is sufficient time to perform the required ground operations. For a fleet of four vehicles, flight times greater than 40min are needed in order to have sufficient time in the air to perform experiments. In addition, the vehicles must have sufficient wing loading capacity to carry the additional weight of the sensors and batteries" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001242_12.659594-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001242_12.659594-Figure3-1.png", "caption": "Figure 3: Illustration showing the unknown \u03b6.", "texts": [ " Also note that equation (11) is always under constraint as (I3-Ra) has rank 2, regardless to the number of in-plane rotations, meaning there is no single solution for tx and the general solution will have exactly one (the number of unknowns minus the rank) arbitrary scale factor \u03b6. So the solution can take the form: nx ttt *)( \u03b6\u03b6 += o where \u03b6 is the unknown scale factor and t\u03bf is a unique solution in the plane of motion (2-dimensional), since (I3-Ra) has rank 2. tn is the normal to the plane of motion (Figure 3). In our case, if the plane of motion is the US image plane (the x-y plane), tn may equal (0, 0, 1)T, which is a unit vector in the z-direction and thus perpendicular to the plane of motion. Figure 4: Experimental system and an illustration for the two suggested special motions. \u2018Motion I\u2019 indicates a planar motion, and \u2018Motion II\u2019 indicates a rotation about an axis. Proc. of SPIE Vol. 6141 61412N-7 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 06/25/2016 Terms of Use: http://spiedigitallibrary" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000400_jra.1986.1087061-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000400_jra.1986.1087061-Figure1-1.png", "caption": "Fig. 1. Zero reference position description for typical revolute axis i, prismatic axis j , and end-effector.", "texts": [ " This research was conducted K. Kazerounian is with the Department of Mechanical Engineering, University of Connecticut, Storrs, CT 06268, USA. K. C. Gupta is with the Department of Mechanical Engineering, University of Illinois at Chicago, Chicago, IL 60680, USA. IEEE Log Number 8610327. II. ZERO REFERENCE POSITION DESCRIPTION In the zero position description 1141, 1151, a suitable configuration of the manipulator is designated as the zero reference position where all of the joint variables have zero values (Fig. 1 ) . In this zero reference position the unit vectors along the revolute or prismatic joints (a& along the kth joint, k = 1 to 6) as well as the position vector of a point on the axis of each joint ( Qok, k = 1 to 6 ) are given in the base coordinate system. In Fig. 1 , k i for a typical revolute joint and k = j for a typical prismatic joint. In addition, the position vector poh of a reference point h on the hand and two perpendicular unit vectors through the point h (preferably an axial and a transverse unit vector, u,, and uOr) are also given. All of the joint variables are set to zero at this reference position. The unit vectors U&(k = 1 to 6) , u,, and uot, and position vectors Qok(k = 1 to 6 ) and P o h completely define the kinematic structure of the manipulator" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003319_j.ijmecsci.2010.10.005-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003319_j.ijmecsci.2010.10.005-Figure9-1.png", "caption": "Fig. 9. Photographs of the experimental set-up and the two deformation patterns", "texts": [ " We conclude that this line-contact deformation becomes unstable via a limit-point bifurcation at p. It is noted that the two deformations in Figs. 7 and 8 buckle to deformations via different paths. In the secondary buckling, the line-contact region bulges up before the point of separation, as demonstrated in Fig. 7(a). In a limit-point bifurcation, the linecontact segment \u2018\u2018peels\u2019\u2019 off the wall from the point of separation, as demonstrated in Fig. 8(a). In order to examine the above theoretical predictions, we design an experimental set-up, as shown in Fig. 9(a), to observe the when P\u00bc6 and sp\u00bc0.64. (a) Pattern . (b) Pattern . deformation patterns. The elastica is made of carbon steel (type SK5) with Young\u2019s modulus 205 GPa and mass density 7830 kg/m3. The cross section of the elastica is 46 mm 0.1 mm. The original length of the metallic strip is 35 cm. In order to minimize the effect of gravity, the metallic strip is placed with the width in the vertical direction. A vertical flat plate with a long slit in the longitudinal direction of the elastica serves as the plane wall constraint in the model", " In the case when the loading point is close to the constraining wall, the metallic ring is allowed to hide in the longitudinal slit on the flat plate to avoid interference. We use a micro-stage to measure the displacement of the midpoint. On the micro-stage we attach a metallic probe connected to a multimeter. An alarm will sound off when the probe touches the elastic strip. A caliper is used to measure the vertical movement of the dead weight. The accuracy limits of the midpoint and loading point movement measurements are 0.01 and 0.02 mm. The photographs in Fig. 9 show two stable deformations for the case when P\u00bc6 and sp\u00bc0.64, one in pattern (Fig. 9(a)) and the other (Fig. 9(b)). The static measurements are plotted with cross marks in Figs. 2 through 5 as the point load moves from the left to the right. At the last point before a jump occurs, the cross mark is circled. For convenient reference, we add the physical units on the right and top sides of the diagram. The physical magnitudes corresponding to dimensionless forces P\u00bc2, 3.4, 6, and 10 are 0.13, 0.22, 0.38, and 0.64 N, respectively. It is found that very good agreement between the experiment and theory is achieved" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001832_07ias.2007.187-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001832_07ias.2007.187-Figure2-1.png", "caption": "Figure 2. Finite Element Mesh", "texts": [ " For the case of a 5kHz switching frequency and 60Hz fundamental, this results in 500 steps per cycle. The voltage wave form is assumed to be an ideal PWM waveform, i.e. dead-time, rise-time and device voltage drop are all neglected and the input voltages to the This work has been supported by the Natural Sciences and Engineering Research Council of Canada, under a Discovery Grant entitled \u201cEnergy Efficient Power Conversion\u201d. 0197-2618/07/$25.00 \u00a9 2007 IEEE 1193 machine are assumed to be constant throughout the simulation time step. The mesh has 1986 nodes per slice and is shown in Fig. 2. The total iron loss due to changing flux density in an element of a finite element mesh may calculated [16],[17] using 1 0.65\u02c6 1 \u02c6 n h h i i p l k fB B B \u03b1\u03b4 = = \u2206 + \u2211 (1) 22 1 12e l d dBp dt T dt \u03c3\u2206 = \u222b (2) 3 2 e x T l k dBp dt T dt \u03b4\u2206 = \u222b (3) where ph, pe, px are hysteresis, eddy and excess loss components, respectively, \u2206 is element area, l is stack length, adjusted for stacking factor, d, \u03b4, \u03c3 are the lamination thickness, mass density and conductivity , kh, ke, \u03b1 are material constants, T is the fundamental time period, Bi are flux reversals" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003444_j.proeng.2010.04.132-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003444_j.proeng.2010.04.132-Figure4-1.png", "caption": "Figure 4: Hammer model", "texts": [ " The vertical lines on the same axis represent the low points during double-support phase and the high points during single-support phases. Although the velocity and the angular velocity increased gradually, they both increased not monotonically but in an oscillating way over the course of the turns. The local maximums of the angular velocity did not coincide with one of the velocity and phase delays were seen between the angular velocity profile and the linear velocity one. To study the hammer movement and energy pumping, a hula-hoop model is employed (Fig. 4). The pendulum moves on an inclined plane at a fixed angle \u03b1 and has mass m and an inextensible weightless wire of length l. The hammer moves in the Cartesian coordinates \u03a3xy on the plane. The motion degrees of freedom of the pendulum are represented by the position vector of the handle p(t) = [x0(t), y0(t)]T and the counterclockwise angle of deviation of the pendulum q(t) from the x axis. The hammer handle moves with circular movement similar to a hula-hoop motion during the turns. The position vector of the hammer is X(t) = p(t) + l el(t) (2) and it\u2019s velocity vector is X\u0307(t) = p\u0307(t) + l q\u0307(t)eq(t), (3) where el(t) = [cos q(t), sin q(t)]T is a unit vector in the normal direction and eq(t) = [\u2212 sin q(t), cos q(t)]T is the tangential unit vector" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003863_b978-0-12-417049-0.00002-x-Figure2.9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003863_b978-0-12-417049-0.00002-x-Figure2.9-1.png", "caption": "Figure 2.9 (A) Kinematic structure of a car-like robot, (B) A car-like robot prototype. Source: http://sqrt-1.dk/robot/robot.php", "texts": [ " This Jacobian is again noninvertible, but we can find the inverse kinematic equations directly using the relations: _\u03c6=v5 \u00f01=D\u00detg\u03c8 or \u03c85 arctg\u00f0D _\u03c6=v\u00de \u00f02:47a\u00de and _\u03b8 w 5 vw r 5 1 r ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v2 1 \u00f0D _\u03c6\u00de2 q \u00f02:47b\u00de The instantaneous curvature radius R is given by (Figure 2.8): R5Dtg\u00f0\u03c0=22\u03c8\u00f0t\u00de\u00de \u00f02:48\u00de From Eq. (2.45) we see that the tricycle is again a 2-input driftless affine system with vector fields: g1 5 r cos \u03c8 cos \u03c6 r cos \u03c8 sin \u03c6 \u00f0r=D\u00desin \u03c8 0 2 664 3 775; g2 5 0 0 0 1 2 664 3 775 that allow steering wheel motion _\u03b8w, and steering angle motion _\u03c8, respectively. The geometry of the car-like mobile robot is shown in Figure 2.9A and the A.W.E. S.O.M.-9000 line-tracking car-like robot prototype (Aalborg University) in Figure 2.9B. The state of the robot\u2019s motion is represented by the vector [20]: p5 \u00bdxQ; yQ;\u03c6;\u03c8 T \u00f02:49\u00de where xQ, yQ are the Cartesian coordinates of the wheel axis midpoint Q, \u03c6 is the orientation angle of the vehicle, and \u03c8 is the steering angle. Here, we have two nonholonomic constraints, one for each wheel pair, that is: 2 _xQ sin \u03c61 _yQ cos \u03c65 0 \u00f02:50a\u00de 2 _xpsin\u00f0\u03c61\u03c8\u00de1 _ypcos\u00f0\u03c61\u03c8\u00de5 0 \u00f02:50b\u00de where xp and yp are the position coordinates of the front wheels midpoint P. From Figure 2.9 we get: xp 5 xQ 1D cos \u03c6; yp 5 yQ 1D sin \u03c6 Using these relations the second kinematic constraint (2.50b) becomes: 2 _xQsin\u00f0\u03c61\u03c8\u00de1 _yQcos\u00f0\u03c61\u03c8\u00de1D\u00f0cos \u03c8\u00de _\u03c6 The two nonholonomic constraints are written in the matrix form: M\u00f0p\u00de _p5 0 \u00f02:51a\u00de where M\u00f0p\u00de5 2sin \u03c6 cos \u03c6 0 0 2sin\u00f0\u03c61\u03c8\u00de cos\u00f0\u03c61\u03c8\u00de D cos \u03c8 0 \u00f02:51b\u00de The kinematic equations for a rear-wheel driving car are found to be (Figure 2.9): _xQ 5 v1cos \u03c6 _yQ 5 v1sin \u03c6 _\u03c65 1 D vwsin \u03c8 5 1 D v1tg \u03c8 _\u03c85 v2 \u00f02:52\u00de These equations can be written in the affine form: _xQ _yQ _\u03c6 _\u03c8 2 664 3 7755 cos \u03c6 sin \u03c6 \u00f01=D\u00detg\u03c8 0 2 664 3 775v1 1 0 0 0 1 2 664 3 775v2 \u00f02:53\u00de that has the vector fields: g1 5 cos \u03c6 sin \u03c6 \u00f01=D\u00detg\u03c8 0 2 664 3 775; g2 5 0 0 0 1 2 664 3 775 allowing the driving motion v1 and the steering motion v2 5 _\u03c8, respectively. The Jacobian form of Eq. (2.53) is: _p5 Jv; v5 \u00bdv1; v2 T \u00f02:54\u00de with Jacobian matrix: J5 cos \u03c6 0 sin \u03c6 0 \u00f0tg\u03c8\u00de=D 0 0 1 2 664 3 775 \u00f02:55\u00de Here, there is a singularity at \u03c85 6 \u03c0=2, which corresponds to the \u201cjamming\u201d of the WMR when the front wheels are normal to the longitudinal axis of its body", "46) [20]: _p5 Jv; J5 cos \u03c6 cos \u03c8 0 sin \u03c6 cos \u03c8 0 \u00f0sin \u03c8\u00de=D 0 0 1 2 664 3 775 \u00f02:56a\u00de In this case the previous singularity does not occur, since at \u03c856\u03c0=2 the car can still (in principle) pivot about its rear wheels. Using the new inputs u1 and u2 defined as: u1 5 v1; u2 5 \u00f01=D\u00desin\u00f0\u03b62\u03c6\u00dev1 1 v2 the above model is transformed to: _xp _yp _\u03c6 _\u03b6 2 664 3 7755 cos \u03b6 sin \u03b6 \u00f01=D\u00desin\u00f0\u03b62\u03c6\u00de 0 2 664 3 775u1 1 0 0 0 1 2 664 3 775u2 \u00f02:56b\u00de where \u03b65\u03c61\u03c8 is the total steering angle with respect to the axis Ox. Indeed, from xp 5 xQ 1D cos \u03c6 and yp 5 yQ 1D sin \u03c6 (Figure 2.9), and Eq. (2.56a) we get: _xp 5 _xQ 2D\u00f0sin \u03c6\u00de _\u03c65 \u00f0cos \u03c6 cos \u03c82 sin \u03c6 sin \u03c8\u00dev1 5 \u00bdcos\u00f0\u03c61\u03c8\u00de v1 5 \u00f0cos \u03b6\u00deu1 _yp 5 _yQ 1D\u00f0cos \u03c6\u00de _\u03c65 \u00f0sin \u03c6 cos \u03c81 cos \u03c6 sin \u03c8\u00dev1 5 sin\u00f0\u03c61\u03c8\u00de\u00bd v1 5 \u00f0sin \u03b6\u00deu1 _\u03c65 1 D sin\u00f0\u03b62\u03c6\u00de\u00bd v1 5 1 D sin\u00f0\u03b62\u03c6\u00de u1 _\u03b65 _\u03c61 _\u03c85 1 D sin\u00f0\u03b62\u03c6\u00de v1 1 v2 5 u2 We observe, from Eq. (2.56b), that the kinematic model for xp, yp, and \u03b6 (i.e., the first, second, and fourth equation in Eq. (2.56b)) is actually a unicycle model (2.28a). Two special cases of the above car-like model are known as: Reeds-Shepp car Dubins car The Reeds-Shepp car is obtained by restricting the values of the velocity v1 to three distinct values 11, 0, and 21" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003423_1.4002259-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003423_1.4002259-Figure10-1.png", "caption": "Fig. 10 Meshed plane strain/stress model", "texts": [ " The nite element analyses were performed using finite element analy- 31205-6 / Vol. 133, JUNE 2011 om: http://pressurevesseltech.asmedigitalcollection.asme.org/ on 01/27/20 sis code ANSYS9.0 . Two-dimensional eight nodded isoparametric element PLANE82 and one-quarter geometry model with symmetrical boundary condition are adopted. Contact 172 and Target 169 elements are used to simulate the surface-to-surface contact between the deformable surface of the tube and tubesheet. The meshed FE planar model is shown in Fig. 10. The nonlinear of materials and the nonlinear of geometry are considered in the FEA. An example of calculated residual contact pressure distribution on the contact surface is shown in Fig. 11. Residual contact pressure versus expansion pressure obtained by FEA in plane strain and plane stress states, respectively, and the corresponding results predicted by present model case 1 are plotted in Fig. 12. Transactions of the ASME 16 Terms of Use: http://www.asme.org/about-asme/terms-of-use F a F f a T a w fi t s F h p m p p T d i s s f t i f m c e c i c c h J Downloaded Fr EA results in plane strain and plane stress states, respectively, nd the results predicted by present model case 2 are plotted in ig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002061_6.2007-6525-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002061_6.2007-6525-Figure4-1.png", "caption": "Figure 4. First two bending modes for HiLDA model.", "texts": [], "surrounding_texts": [ "American Institute of Aeronautics and Astronautics\n7\n3 Rockwell International Corporation, \u201cDesign and Development of a Structural Mode Control System\u201d. NASA CR-143846. October 1977.\n4Gregory, I. M. \"Modified Dynamic Inversion to Control Large Flexible Aircraft \u2013 What\u2019s Going On?\" AIAA Guidance, Navigation and Control Conference, No. AIAA99-3998, Portland, OR, August 9-11, 1999.\n5Gregory, I. M. \u201cDesign and Stability Analysis of Integrated Flight/Structural Mode Controller for Highly Flexible Advanced Aircraft Utilizing a Novel Nonlinear Dynamic Inversion.\u201d Ph.D. Thesis, California Institute of Technology, 2004.\n6Cao, C. and Hovakimyan, N. \u201cDesign and Analysis of a Novel 1L Adaptive Controller, Part I: Control Signal and Asymptotic Stability.\u201d In Proc. of American Control Conference, 2006, pp. 3594\u20133599.\n7Cao, C. and Hovakimyan, N.. \u201cDesign and Analysis of a Novel 1L Adaptive Controller, Part II: Guaranteed Transient Performance.\u201d In Proc. of American Control Conference, 2006, pp. 3397\u20133402.\n8Cao, C. and Hovakimyan, N. \u201cGuaranteed Transient Performance with 1L Adaptive Controller for Systems with Unknown Time-Varying Parameters and Bounded Disturbances: Part I\u201d In Proc. of American Control Conference, 2007, pp. 3925-3930.\n9Cao, C. and Hovakimyan, N. \u201cStability Margins of 1L Adaptive Controller: Part II\u201d In Proc. of American Control Conference, 2007, pp. 3931-3936.\n10Cao, C., Patel, V.V., Reddy, C.K., Hovakimyan, N., Lavretsky, E., and Wise., K. \u201cAre Phase and Time-delay Margins Always Adversely Affected by High-Gain?\u201d AIAA Guidance, Navigation, and Control Conference, No. AIAA2006-6347, Keystone, CO, 2006.\n11Roger, K.L., \u201cAirplane Math Modeling and Active Aeroelastic Control Design\u201d, AGARD-CP-228, 1977, pp.4.1-4.11. 12Cao, C. and N. Hovakimyan. \u201cAdaptive Output Feedback in the 1L Framework.\u201d Submitted to 2008 American Control\nConference, Seattle, WA, June 11-13, 2008\nIX. Figures", "American Institute of Aeronautics and Astronautics\n8", "American Institute of Aeronautics and Astronautics\n9" ] }, { "image_filename": "designv11_20_0000849_sensor.1995.717235-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000849_sensor.1995.717235-Figure3-1.png", "caption": "Fig. 3 SEF measurement part", "texts": [ "2B) Glucose oxidase solution, bovine serum albumin solution, and gultaraldehyde solution, were mixed and spin-coated on the wafer at 1,500 r.p.m. 4) Lift-off (Fig2C) After 30 minutes, the enzyme membrane, other than that on the enzyme ISFET, was lifted off in acetone by ultrasonic vibration. The thickness of the enzyme membrane was very uniform, except at the edges. Albumin membranes were also deposited on both ISFET surfaces, which canceled the influence of potential change due to protein adhesion on the surface (5). A measuring cell, where the ISFET glucose sensor is mounted, and an SEF collecting cell are integrated as shown in Fig. 3. Figure 4 shows a top view and a cross-sectional view of the measuring cell. 448 TRANSDUCERS '95 * EUROSENSORS IX The 8th International Conference on Solid-state Sensors and Actuators, and Eurosensors IX. Stockholm, Sweden, June 25-29, 1995 Figure 5 shows a cross-sectional view of the SEF collecting cell. It consists of a rotary valve, and a suction cell whose diameter is about 4.5 cm. R o t a r y v a l v e H In an animal experiment, the fur on a part of a rabbit's back was trimmed and shaived off" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002288_j.mechmachtheory.2008.03.007-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002288_j.mechmachtheory.2008.03.007-Figure2-1.png", "caption": "Fig. 2. Schematic of the upper body of the machine in the reference configuration.", "texts": [ " 1 in frame S in the basic configuration, and let hti \u00f0u\u00de \u00f0i \u00bc 1; . . . ;4\u00de be the joint motion functions, casted as ht\u00f0u\u00de \u00bc \u00f0ht1 \u00f0u\u00de; . . . ; ht4 \u00f0u\u00de\u00de, all depending on a common parameter of motion, denoted as u. Therefore ht : R! Qt , where Q t is the configuration space of the joints controlling the motion of the tool. The common feature of the joint motion functions employed in the product of exponentials formula (23), is that they must be zero for a zero value of u, that is hti \u00f00\u00de \u00bc 0. Since the reference configuration, depicted in Fig. 2, differs from the basic configuration (Fig. 1) by a constant posture, we can take it into account by introducing the (constant) joint offset values h0 i . Due to the conventions adopted by Gleason, and the basic configuration chosen, the non-zero joint offset values are only h0 t3 and h0 t4 : this are the constant components of the swivel and tilt motions, respectively. From this, it follows imme- diately that only the coordinates of the last twist in the basic configuration, n0 t4 , differ from the correspondent ones in the reference configuration, nt4 " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.78-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.78-1.png", "caption": "Fig. 14.78 Portable single-mast climbing platform with rack-and-pinion lifting gear", "texts": [], "surrounding_texts": [ "General Elevating work platforms form a wide class of cranes for elevating persons and equipment for repair, maintenance, and assembly purposes. Depending on their design and application, elevating work platforms can be divided into the following groups: \u2022 Portable mast-climbing platforms\u2022 Mobile (mounted on special chassis) elevating work platforms\u2022 Hanging scaffolds Portable Mast-Climbing Platforms Portable mast-climbing platforms are used for masonry plastering, assembly, insulation, and facade works in housing and industrial construction. Portable mast-climbing platforms (Figs. 14.78 and 14.79 and Tables 14.9 and 14.10) are made up of the following units: Part B 1 4 .5 Fig. 14.74 Facade, frame scaffold in atypical version. The atypicality consists of the lack of anchoring \u2022 A base\u2022 A mast or multisectional masts\u2022 A work platform lifting unit\u2022 A work platform The work platform moves on a mast fixed to a base and anchored to a wall of a building structure. The mast is +66.00 m +64.00 m +56.00 m +48.00 m Fig. 14.76 Facade, frame scaffold in atypical version. The atypicality consists of the shape of scaffold support scheme, which differs greatly from the typical version Part B 1 4 .5 Table 14.9 Specifications of typical portable two-mast climbing platforms Parameters Platform\u2019s length/lifting capacity (m/kg) 4.1/1300 4.2/1300 4.2/2000 4.2/2700 7.1/800 7.4/1000 7.4/1700 7.3/2300 10.1/500 10.5/700 10.5/1400 10.5/1900 12.5/1200 13.7/1500 16.9/1000 Maximum platform elevation without anchoring \u2013 Protractible beams protracted on both sides (m) 6 15 10 20 \u2013 Protractible beams protracted on one side of mast (m) 6 15 15 12.5\u201317.5 Maximum platform elevation with 11.5 25 25 25 one mast anchoring point located at top (m) Maximum platform elevation with 100 200 200 200 anchoring along entire length of mast (m) Spacing between anchors (m) 6 12.5 12.5 12.5 Max. length of platform\u2019s protractible part (m) 1.0 0.3 1.4 2.5 Max. loading of struts (kN) 15 50 60 65 Lifting speed (m/min) 6 6 6 6 Transport mass (kg) 3700 2 \u00d7 3500 2 \u00d7 4000 2 \u00d7 4000 Mast section: length/weight (mm/kg) 1508/48 1256/82 1256/82 1256/82 Electric specifications of platform lifting gear 400 V/50 Hz 400 V/50 Hz 400 V/50 Hz 2 \u00d7 400 V/50 Hz 3 kW, 16 A 3 kW, 16 A 3 kW, 16 A 3 kW, 16 A 2.50 2.00 2.00 2.00 2.00 1.50 0.72 2.50 Fig. 14.77 Atypical suspended scaffold. The atypicality consists of the suspending main part of scaffold made up of sections added up to the required height. In the case of a free-standing mast, the work platform can be usually lifted to a height of 20 m. The mast\u2019s base can have the form of a carriage or be stationary. If a carriage is used as the base, the free-standing mast-climbing platform can be moved without it being it necessary to disassemble the mast completely. The carriage can be towed by a tractor or be self-propelled and so able to move along the wall of the building structure. A small-sized stationary base is used when the space for the mast-climbing platform is restricted and no carriage can be employed, e.g., in a street with a narrow pavement. The work platform is elevated by a rack-and-pinion gear. Thanks to its sectional design the mast-climbing platform can be configured to fit the building structure\u2019s shape. In addition, protractible struts, with length adjustable from 0 to 2500 mm, can be attached to the mast-climbing platform\u2019s sections. Planks or wooden boards are placed on the beams, thereby creating additional working surface. The combination of the mast-climbing platform\u2019s sectional structure and the system of protractible struts makes it possible to obtain work access on walls of any shape (straight, curved, and with slants and bevels) and architectonic form (balcony, Part B 1 4 .5 loggia, niche, bay). The whole mast-climbing platform is fenced in with railings to protect persons working on it from falling out. It is also possible to combine two single-mast climbing platforms to form one platform (up to 40 m long) climbing two masts (Fig. 14.79 and Table 14.10). In many cases, mast-climbing platforms may replace stationary construction-assembly scaffolds. Similarly to material hoists and person and material hoists with a rack-and-pinion drive, mast-climbing platforms are equipped with the following safety devices: \u2022 An emergency lowering system\u2022 A braking device\u2022 A safety device preventing the driving gear wheel from disengaging from the mast\u2019s gear rack\u2022 Electric-shock protection\u2022 Overload protection for the electric motors\u2022 Work-platform-slanting cutouts (in two-mast climbing platforms)\u2022 Work platform terminal position cutouts\u2022 Sensors signalling a platform loading which may result in overturning of the mast-climbing platform or its damage Mobile Elevating Work Platforms Mobile elevating work platforms have a similar range of applications (elevating persons and equipment) as the portable mast-climbing platforms described above, except that their use in one work place is short. Mobile elevating work platforms form a class of devices varied in their design. The basic types of mobile elevating work platforms are listed in Table 14.10. Mobile elevating platforms are usually equipped with protractible struts. Truck-mounted platforms are the most popular mobile elevating work platforms used in construction. Table 14.10 Types of mobile elevating work platforms Description of mobile elevating work platforms Design schematics A mobile elevating work platform featuring a telescopic boom. The boom can be raised, lowered, and slewed relative to the vertical axis. 1 2 3 A platform mounted on a telescopic column or a hydraulic servomotor. The working motions are: the raising and lowering of the platform in the vertical plane. 1 2 3 4 A mobile elevating platform with a scissor extending structure. 2 3 1 A truck-mounted elevating platform whose extending structure is usually in the form of a two-stage boom. The elevating platform performs the following motions: \u2022 Raising and lowering \u2022 Slewing relative the vertical axis perpendicular to the base 1 2 3 Key: 1 platform; 2 extending structure; 3 chassis; 4 struts Truck-Mounted Elevating Platform The truck-mounted elevating platform consists of the following parts shown in Fig. 14.80. The supporting frame is a body to which protractible struts, a hydraulic feeder, and a slewing gear are fixed. The supporting frame is secured to a truck chassis. Part B 1 4 .5 The slewing column is attached to the supporting frame through a crown-bearing. The column is a slewing welded-construction frame with a transmission gear and a brake mounted on it. The lower boom stage and the cylinder are attached to the column by articulated joints. The boom consists of two stages connected by an articulated joint. Each stage is a welded box structure. The work platform consists of a floor made from thin-walled steel sections, railings, and curbs. The straight-line system is a tension-member structure which keeps the work platform in a horizontal position regardless of the angles at which the boom\u2019s stages are positioned. The work platform positioning system consists of a pump, filters, distributors, valves, steel pipes, hoses, and hydraulic cylinders. The hydraulic pump is powered from the truck\u2019s gearbox via a lay shaft. Oil is supplied to the working cylinders through a rotary joint. The struts are controlled through distributors mounted on the supporting frame. The boom\u2019s cylinders and slewing can be controlled through a distributor mounted near the slewing column or on the work platform. The system is protected against the eventuality of simultaneous steering of the struts and the boom\u2019s cylinders. The boom\u2019s cylinders and those of the struts are equipped with valves to prevent a pressure failure and the resulting uncontrolled shift of the piston rod. The maximum angle of elevation of the boom\u2019s upper stage is limited by a limit switch in the form of a hydraulic valve controlled by a cam on the boom. The hydraulic system\u2019s circuits are protected against excessive pressure rise by overflow valves. In elevating platforms capable of large elevations (20\u201340 m), proportional control, ensuring that the working motions controlled by the operating lever are fluid and quick, is used instead of the typical hydraulic control. Truck-mounted elevating platforms have the following safety measures: \u2022 Limit switches preventing the boom\u2019s upper section from being excessively raised and the boom\u2019s lower stage from being lowered while the boom\u2019s upper stage is maximally raised\u2022 A hydraulic system lock preventing the working circuit and the struts from being simultaneously fed\u2022 Overload protection in the form of overflow valves Part B 1 4 .5 \u2022 Emergency lowering of the cradle while the pump drive is switched off The truck-mounted elevating platform\u2019s basic operating parameters are: \u2022 Lifting capacity\u2022 Maximum elevation\u2022 Maximum radius\u2022 Work area, specifying the allowable position of the elevating platform in the vertical plane\u2022 Angle of rotation of the body The work area with specified positions of the work platform in the vertical plane is shown in Fig. 14.81. Mobile elevating platforms can be mounted on mass-produced truck chassis. The type and size of chassis depends on the specifications of the elevating platform. The characteristics of selected chassis for the particular groups are detailed in Table 14.11." ] }, { "image_filename": "designv11_20_0000980_1-4020-3559-4-Figure16-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000980_1-4020-3559-4-Figure16-1.png", "caption": "Figure 16. Primary suspension model of the ML95 trailer bogie: a) Threedimensional spring-damper elements; b) Suspension model with springs and dampers.", "texts": [ " The multibody model of the trailer vehicle of the train, developed in the work by Pombo [13], is composed of the car shell suspended by a set of springs, dampers and other rigid connecting elements on the bogies. This assembly of connective elements constitutes the secondary suspension, sketched in Fig. 15, which is the main one responsible for the passenger\u2019s comfort. The connections between the bogies chassis and the wheelsets, also achieved by another set of springs, dampers and rigid connecting ele- ments, constitute the primary suspension represented in Fig. 16. The primary suspension is the main suspension responsible for the vehicle running stability. The simulation results of the vehicle, running in a circular track with a radius of 200 m with velocities of 10 and 20 m/s, show that the prediction of flange contact is of fundamental importance. Fig. 17 shows that contact forces obtained with the Kalker linear theory originate the lift of the outer wheel of the front wheelset at the entrance of the curve. Despite this wheel lift, derailment does not occur and the analysis proceeds up to end" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.147-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.147-1.png", "caption": "Fig. 14.147a\u2013c Working dimensions of a mini-crane for erecting aerated concrete block walls: (a) spacing of supports; (b) vertical radius; (c) dimensions in transport mode", "texts": [ "53] one can find more examples of the automation of crane control systems such as a crawler hydraulic crane for dam construction and a tower crane for the automatic distribution of concrete mix during the erection of a high-rise building. Part B 1 4 .8 14.8.5 Automation of Materials Handling and Elements Mounting by Mini-Cranes and Lightweight Manipulators Mini-cranes and lightweight manipulators are used for materials handling, fitting building finishing elements, and transferring heavy construction equipment that cannot be moved manually. A radio-controlled mini-crane is shown in Fig. 14.147. The machine is intended for putting up aerated concrete block walls. This mini-crane with a three-sectioned telescopic boom has a lifting capacity of 500 N at an operating radius of 3 m. In its working mode the mini-crane is supported by four outriggers, whereas for relocation or transport it is mounted on a crawler chassis. The crane can move up and down stairs and for transport can be fitted into a delivery truck. The crane\u2019s working dimensions are given in Fig. 14.147. A lightweight manipulator designed for assembly work and transporting construction equipment inside buildings is shown in Fig. 14.148. Equipped with all kinds of attachments it can be used instead of interior work scaffoldings as a staging for ceiling work such as the fitting of lightweight beams, soffit, lightening, and heavy items. Its main advantage is that the position of the fitted element can be adjusted by means of the traversing and lifting gears operated from the handle attached to the manipulator\u2019s tip" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003183_b978-0-12-374227-8.00004-3-Figure4.1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003183_b978-0-12-374227-8.00004-3-Figure4.1-1.png", "caption": "Figure 4.1: Left: Schematic side view of a H -driven fl agellar motor, with the proposed location and copy number of proteins involved in torque generation. MotA and MotB are thought to form stator complexes with stoichiometry A 4 B 2 , and FliN a tetramer that has 1:1 stoichiometry with FliM. The motor spans the three layers of the cell envelope: outer membrane (OM), peptidoglycan cell wall (PG), and cytoplasmic membrane (CM). Right: detail of proposed location and orientation of rotor proteins. X-ray crystal structures of truncated rotor proteins, FliG (PDB ID 1lkv), FliM (PDB ID 2hp7) and FliN (PDB ID 1yab), are shown docked into the rotor structure. N- and C-termini and missing amino acids are indicated. Molecular graphics generated using PyMol ( http://www.pymol.org )", "texts": [ " In particular, we focus on (1) new structural information from X-ray crystallography and electron microscopy; (2) measurements of torque and speed; (3) numbers and dynamics of stators, from single molecule fl uorescence microscopy of green fl uorescent protein (GFP)-labeled motor proteins and careful analysis of rotation speeds; (4) the dependence of motor function on ion-motive force (IMF), from manipulation and measurement of IMF in single cells; and (5) stepping motion in fl agellar rotation revealed by particle tracking with nanometer and submillisecond resolution. Like any rotary motor, the bacterial fl agellar motor consists of a rotor and a stator. The rotor spins relative to the cell and is attached to the helical fi lament by a universal joint called the hook; the stator is anchored to the cell wall. Figure 4.1 shows a schematic diagram of the bacterial fl agellum of gram-negative bacteria, based on an EM reconstruction of the rotor from S. typhimurium ( Figure 4.2A ) ( Francis et al., 1994 ; Thomas et al., 1999, 2006 ). The core of the motor is called the \u201c basal body \u201d and consists of a set of rings up to 45 nm in diameter that spans three layers of the cell envelope ( DePamphilis and Adler, 1971a,b,c ). The L- and P-rings are embedded in the outer lipopolysaccharide membrane and peptidoglycan cell wall, respectively, and are thought to work as a bushing between the rotor and the outer parts of the cell envelope", " FliG interacts with MotA to generate torque ( Garza et al., 1995 ; Lloyd et al., 1996 ); FliM binds the active form of the response-regulator CheY, altering the probability of CCW rotation; and FliN may be responsible for the intrinsic bistability of the rotor that gives it relatively stable CW and CCW states. Atomic structures of the middle and C-terminal domains of FliG, middle part of FliM, and C-terminal part of FliN have been resolved by X-ray crystallography ( Brown et al., 2002, 2005 ; Lloyd et al., 1999 ; Park et al., 2006 ). Figure 4.1 (inset) shows these crystal structures and a model for where they fi t into the C-ring based on a range of biochemical studies ( Brown et al., 2007 ; Lowder et al., 2005 ; Paul and Blair, 2006 ; Paul et al., 2006 ). The overall structure of the C-ring is determined by single-particle reconstruction and cryo-EM of the fl agellar basal body. This work has revealed an interesting symmetry mismatch within the rotor. The MS-ring symmetry was fi rst reported to be 26 ( Suzuki et al., 2004 ). A partly functional fusion protein between FliF and FliG is strong evidence that FliG has the same copy number as FliF, the main MS-ring protein ( Francis et al", " No single mutation in these residues completely abolishes torque generation, and chargereversing mutations in both proteins can compensate each other, indicating an electrostatic interaction at an interface between the two proteins. PomA and FliG of the Na -driven fl agellar motor of V. alginolyticus show a similar pattern ( Yakushi et al., 2006 ), but with differences in which conserved charged residues are most important for function ( Yorimitsu et al., 2002, 2003 ). The stoichiometry of stator complexes deduced from targeted disulfi de cross-linking studies and chromatography appears to be A 4 B 2 ( Braun et al., 2004 ; Kojima and Blair, 2004b ; Sato and Homma, 2000 ) ( Figure 4.1 ), with the membrane-spanning helices of MotA subunits surrounding a suspected proton-binding site at residue Asp32 of MotB ( Sharp et al., 1995a,b ). This is the only conserved charged residue in MotA or MotB that is absolutely essential for function, and it is postulated that each stator contains two ion channels, each containing one MotB Asp32 residue ( Braun and Blair, 2001 ). Different patterns of protein digestion indicate a conformational change in MotA between wild-type stators and stators containing the mutation MotB Asp32 to Asn32, which mimics the protonation of Asp32 ( Braun et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003711_pime_conf_1967_182_341_02-Figure29.2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003711_pime_conf_1967_182_341_02-Figure29.2-1.png", "caption": "Fig. 29.2. Schematic view of vibration tester", "texts": [ " Vibration measuring equipment, which is known as VKL, is frequently mentioned in this paper. It contains essentially the same mechanical and electronic parts as the VKR and VKK equipment. In particular the spindle, pick-up, and preamplifier are the same and, generally, the same type of filters are used. As previously mentioned, the VKL equipment contains a device for obtaining a purely axial load over the bearing, and the pick-up measures the radial velocity at a point at the outer perimeter of the non-rotating outer ring. This is illustrated in Fig. 29.2. The pick-up will now be discussed, since it is the most questionable part of the test equipment at the present time. We shall not attempt a general discussion on pos- Proe Insrn Mech Engrs 3967-68 sible types of pick-up, but concentrate upon the SKE electrodynamic pick-up. The main technical data of the pick-up arc as follows : almost linear frequency characteristic from 0 to 10 000 Hz a t a measuring load of 0-2 newton; the weight of the moving system is 0-17 g; the radius of the pick-up tip is 0", " The circular frequency flJk and the effective values of the amplitudes ( l /v\u20192)bk constitute the frequency spectrum at x = zo of our surface S when rotating with constant angular velocity w around the fixed z-axis. Calling k the order of waviness we define the waviness spectrum at z = zo of the surface S as the set of numbers and velocity amplitudes 1 k = 1,2,. . ., m \\Waviness spectrum of Ly at z = zo ,! - b, 4 2 The careful reader at once observes that we have defined the waviness spectrum from a pick-up constrained to move radially while we, according ro Fig. 29.2, measure it along the theoretical normal to the surface S . It is not difficult to show that the waviness velocity amplitudes when measuring along the normal become Proc Instn iWech Engrs 1967-68 where y is the angle between the normal and the x-3\u2019 plane. In general y is a small angle so we may put cos y = 1 and make use of the above definition of waviness spectrum. Turning to the practical problem of determining the waviness spectrum of a real surface in a measuring equipment it is necessary to introduce an upper \u2018cut-off\u2019 waviness order nl", " value 20, = - ( jv- iv+l)- l 2 b;,2 2\u2018 \u201c 2 k = & More precisely this is the effective value of the waviness level in the I& waviness band at z = z, measured at the angular velocity cu. The unit of w\u2019, is taken to be micrometres per second. We assume that the ammeters or recorders used are all r.m.s. instruments. Frequently the amplitude scales of our instruments are logarithmic, i.e. decibcl scales. This is of no importance to us here and does not influence the previous statements. APPENDIX 29.2 V I B R A T I O N SPECTRUM The following results are valid only for the particular kind of mounting obtaining in the VKL tester (Fig. 29.2). Let u be the number of rolling bodies in the test bearing and let q >, 0 and p >, 1 be positive integers. Then the following holds true for the vibrations in radial direction measured at a point of the outer perimeter of the outer ring: (1) Inner ring waviness of orders k = q u f p w = P u ( W i - - W c ) ~ P W . where wz, wc are the inner ring and cage angular velocities as observed from a frame of reference fixed in space. For p = 1 the vibrations are of the rigid body type, i.c. the outer ring moves as a rigid body while for p = 2, 3, etc" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001378_ijmmm.2007.015474-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001378_ijmmm.2007.015474-Figure1-1.png", "caption": "Figure 1 Model of the rounded tool cutting edge", "texts": [ " This means that the layer to be removed by the moving tool is deformed and separates into two different directions. One part of this layer flows along the tool rake face forming the chip and the other part is pressed by the round cutting edge against the workpiece surface, contributing to the ploughing (or burnishing) phenomenon. In this work, it is assumed that the total uncut chip thickness, tr, is separated into the actual uncut chip thickness, t1 and the layer to be burnished, hp, by the round part adjacent to the tool flank face (see Figure 1). Because the actual uncut chip thickness, t1, is smaller than the total uncut chip thickness, tr, the cutting feed corresponding to the actual uncut chip thickness can be thought of as the apparent cutting feed, f1, which is smaller than the real cutting feed, f, measured as the velocity of tool along the workpiece as shown in Figure 2. As seen, the apparent (or corrected) cutting feed can be calculated as ( ) ( ) 1 1 sin sin p r r ht f f \u03ba \u03ba = = \u2212 (2) where \u03bar is the tool cutting edge angle. This equation is valid when tr > hp, or, as it follows from Equation (2), when f > hp /sin(\u03bar) (to keep f1 > 0). The use of small RTS\u2032 also induces changes in the tool rake angle. Figure 1 shows the method used to calculate the corrected normal rake angle, .n\u03b3 \u2032 This angle corresponds to the angle between the trace of the reference pane (the vertical direction in Figure 1) and the tangent line to the rake face profile at point A. According to this method, the corrected normal rake angle calculates as: ( ) ( ) \u03b3 \u03b3 \u03b3 \u03b3 \u23a7 \u239b \u239e \u2212 < +\u23aa \u239c \u239f\u2032 = \u23a8 \u239d \u23a0 \u23aa \u2265 +\u23a9 arcsin 1 if 1 sin if 1 sin r r n n nn n r n n t t r r t r (3) where \u03b3n is the normal tool rake angle. Figure 3 shows the flowchart of the proposed analytical approach to study the influence of RTS on the forces, temperatures and thermal energy distributions in the deformation zone. A detailed description of this approach can be found in Outeiro (2002)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001071_3.58552-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001071_3.58552-Figure6-1.png", "caption": "Fig. 6 Seal land bleed tubes amplified instability to confirm predicted AP effect.", "texts": [ " The analysis showed as the pressure differential across this cavity was increased statically or dynamically, the entire system tended to become aeroelastically unstable. The mechanisms causing an increased pressure change in real life situations would be clearance changes on the subsequent knife-edges, either static or dynamic. The front to back pressure influence on the first two knife-edges prompted the need for an invention. The invention was necessary to turn on the instability during the next test and then, when the fix was implemented, demonstrate enough margin to handle unpredictables and growth. The invention is shown in Fig. 6. It is simply a bleed tube pierced into the second cavity, between the second and third knife-edges. By valving this bleed tube to a sink of any desired pressure, the pressure change across the first cavity could be varied almost at will, with the condition most conducive to instability being with the bleed vented to ambient, which increased the normal static pressure drop by 260%. During the next test of the engine, it was possible to turn on the instability at will, whenever the engine was operating at maximum rotor speed donditions" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002971_j.mechmachtheory.2009.11.001-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002971_j.mechmachtheory.2009.11.001-Figure2-1.png", "caption": "Fig. 2. Model of force diagram for pinion and gear.", "texts": [ " Section 2 derives dynamic models for the gear-bearing system with a nonlinear suspension effect, strongly nonlinear gear mesh force and strongly nonlinear oil-film force. Section 3 describes the techniques used in this study to analyze the dynamic response of the gear-bearing system. Section 4 presents the numerical analysis results obtained for the behavior of the gear-bearing system under various operational conditions. Finally, Section 5 presents some brief conclusions. The dynamic model to stimulate the gear-bearing system under the assumptions of nonlinear suspension effect and strongly nonlinear fluid film force effect is established in Fig. 1. Fig. 2 presents a schematic illustration of the dynamic model considered between gear and pinion. Applying the principles of force equilibrium, the forces acting at the center of journal 1, i.e. Oj1 (Xj1, Yj1) and center of journal 2, i.e. Oj2 (Xj2, Yj2) are given by Fx1 \u00bc fe1 cos u1 \u00fe fu1 sin u1 \u00bc Kp1\u00f0Xp Xj1\u00de=2; \u00f01\u00de Fy1 \u00bc fe1 sin u1 fu1 cos u1 \u00bc Kp1\u00f0Yp Yj1\u00de=2; \u00f02\u00de Fx2 \u00bc fe2 cos u2 \u00fe fu2 sin u2 \u00bc Kp2\u00f0Xg Xj2\u00de=2; \u00f03\u00de Fy2 \u00bc fe2 sin u2 fu2 cos u2 \u00bc Kp2\u00f0Yg Yj2\u00de=2; \u00f04\u00de in which fe1 and fu1 are the viscous damping forces in the radial and tangential directions for the center of journal 1, respectively, and fe2 and fu2 are the viscous damping forces in the radial and tangential directions for the center of journal 2, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.93-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.93-1.png", "caption": "Fig. 14.93 Scaffold mountable winch", "texts": [ "92 shows the use of scaffold cranes, portable cranes, and small-capacity gantries during building erection. Machinery of this type is intended for lifting and transferring loads of up to 200 kg to a height of 80 m. The design and technical specifications of these winches make them a highly effective means of vertical transport in construction work involving scaffolds as well as the assembly and disassembly of scaffolds. Winches in scaffold cranes can be mounted in two ways: \u2022 Outside the crane, to the lowest (from the ground) scaffold upright (Fig. 14.93)\u2022 On the crane\u2019s boom (Fig. 14.92 item 5, and Fig. 14.94) In the case of winches mounted using the former method, a limit switch, functioning also as a load limiter and a block upper position switch, is incorporated into the winch\u2019s housing. The way in which a winch is mounted onto the boom is shown in Fig. 14.94 and Fig. 14.92 item 6. The working radius of the boom with a mounted winch can be changed by protruding the load-bearing tube. There is a series of holes in the inner tube for a blocking pin" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001988_09544070jauto67-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001988_09544070jauto67-Figure3-1.png", "caption": "Fig. 3 (a) Brake ODSs (5739 Hz) at five distinct times during squeal obtained by laser vibrometer and (b) out-of-plane modes induced by in-plane excitation versus different pressures and phases near squeal frequency (reprinted from SAE paper 2004-01-2798 [35] with permission from SAE International)", "texts": [ " brake pressure was at 100 lbf/in2. The scanning laserfrequency range to 250 kHz. When a microscope-type laser vibrometer is used, the minimum measurement vibrometer was set on measure/scan-on-trigger mode. velocity and displacement will be improved. A microphone, which on receiving the squeal signal The laser vibrometer has also been arranged to would send it to the vibrometer to trigger the scan, become a three-dimensional or six-degree-of-freedom was located near the brake. The measured ODS system [32\u201334]. is shown in Fig. 3(a). It can be seen that there is transition from in-plane vibration induced out-ofplane motion/vibration to out-of-plane vibration generated out-of-plane motion/vibration. For further2.1 Applications confirmation of the observation, a special-design 2.1.1 Brake ODS and mode shape measurement and mode shape measurement test is shown in Fig. 4 analysis in which a shaker was attached to the rotor edge in the tangential direction. In this way, the in-planeAs discussed in the introduction, brake rotor ODS vibration/mode induced out-of-plane vibration/modeand mode shapes are important for CAE model correlation and squeal root cause identification. One can be measured. The results are shown in Fig. 3(b). JAUTO67 \u00a9 IMechE 2007Proc. IMechE Vol. 221 Part D: J. Automobile Engineering at UNIV OF MICHIGAN on June 22, 2015pid.sagepub.comDownloaded from As can be seen, at zero brake pressure, there is an in-plane mode induced out-of-plane vibration (compression and extraction type) that matches the through and be reflected from the contact surface of squeal ODS at 0\u00b0. When a light brake pressure of the pad. Figure 6 shows the test results, from which 1 lbf/in2 was applied, an out-of-plane mode induced it can be seen that the contact mode shape changes out-of-plane vibration (bending type) was observed and the resonant frequency increases as the weight [Fig. 3(b)] from 0 to 90\u00b0 phase angles owing to the increases [36]. rotor being coupled with a caliper and other brake components. It should be noted that, since vibration 2.1.2 Panel acoustic contribution analysis (PACA) \u2013 shapes are the same at 0\u00b0 phase between zero mastic optimization pressure and 1 lbf/in2, the one with zero pressure is shown in Fig. 3(b) with a slightly different resonant As mentioned in the introduction, vehicle body panel vibration can generate vehicle interior noise. Panelfrequency. With the above results available, it suggests that there is significant in-plane mode/vibration damping, stiffening, or mass loading treatments are often used to reduce vibrations. Firstly the scanningduring squeal. How the in-plane vibration contributes to the squeal will be the subject of further study, and laser vibrometer (SLV) is used to identify the locations with large vibration amplitudes and highsome details can be found in reference [35]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000420_robot.2004.1308859-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000420_robot.2004.1308859-Figure1-1.png", "caption": "Fig. 1. Passive one-legged hopper,", "texts": [], "surrounding_texts": [ "Pmusdlnga of me 2004 IEEE Internatltlonal Conkrsnce on Robouu 6 AutOmatlon\nNew Orleans. LA * npdl2004\nRunning Control of a Planar Biped Robot based on Energy -Preserving Strategy\nSang-Ho Hyon Department of Bioengineering and Robotics\nGraduate School of Engineering, Tohoku University Aramaki-&a-Aoba 01, Sendai 980-8579, Japan\nEmail: sangho@ieee.org\nAbslmct-In this paper, we extend running controller of passive one-legged hopper to a planar biped robot with torso, and evaluate the controller on simulations. The controller is derived based on energy-preserving strategy and it actually preserves mechanical energy at touchdown. Interestingly, zero dynamics of decoupling controller (dynamics about pair of controlled leg) i s found to be stable. Combining simple attitude controller at stance phase generates stable periodic running gaits of arbitral period. The control performance is shown to be better than a simple PD-feedback control of leg placement.\nI. INTRODUCTION A. Backgmund\nAfter the Raiben\u2019s excellent works [I], running robots have been widely studied both experimentally [21[31[41[51[61 and theoretically [7][8]. On the other hand, recently there are many studies on biped humanoid robots with the aim of practical application.. Therefore, enhancing the mobility of them is important target to be reached. Running control of biped humanoid robots is included in such targets. Energy-efficient control of fast running is especially crucial for autonomous humanoid robots because it directly extends operation time.\nIn this connection, there are some remarkable researches on energy-efficient running control. Tompson and Raiben showed that spring-driven one-legged bopping robot can hop without any inputs, provided if the initial conditions were appropriately chosen [9]. Ahmadi and Buehler applied Raiben\u2019s algorithm to this robot and realized energy-efficient hopping in simulation and experiment [IO]. Franpis and Samson derived a rather systematic controller based on the general control method used in nonlinear oscillatory system [ I I].\nE. Energy-preserving controller for one-legged hopper Motivated from their works, we proposed, in [12], altemative controller based on its energy analysis for a planar onelegged robot shown in Fig. I .\nThe underlining principle is energy-presenring contml strategy. This means the controller preserves system energy as much as possible. The most important reason why we use this strategy is: if the system energy is preserved, it is expected that the system autonomously generates natural periodic gaits, just as some class of Hamiltonian system exhibit natural periodic orhit [13]. Instead of depending on some pre-calculated\nTakashi Emma Department of Bioengineering and Robotics\nGraduate School of Engineering, Tohoku University Aramaki-ha-Aoba 01, Sendai 980-8579, Japan\nEmail: emura@emuramech.tohoku.ac.jp\nperiodic solutions, or target (desired) dynamics, analysis on energy change of the original nonlinear hybrid system were utilized as shown below.\nFirst, we choose desired touchdown angle @ d and angular velocity Bd at the moment of lift-off to meet the following energy non-dissiparion condition:\nPtd- := k t d - COS Btd + &d- Sin 6 t d + TO = b t d - = 0 ( I ) Subscript \u201ctd\u201d means just the moment of touchdown and \u201ctd-\u201d or \u2018\u2019td+\u201d mean just before I after touchdown hereinafter. Since the energy change between just before and after touchdown is calculated as\ncondition ( I ) means there is no energy exchange between the robot and the ground, provided if no control +put applied during stance phase. Having determined 6\u2019d and @ d . finally we can apply simple linear dead-beat controller because the flight dynamics is integrable.\nAs a result, interesting quasi-periodic gaits, which can be seen in some Hamiltonian system, were found, and both period\n0-7803-8232-3/04/$17.00 @ZOO4 IEEE 3791", "stabilization and one-periodic passive running were achieved in simulation. It was also found that an adaptive control of touchdown angle, which is similar to the delayed feedback controller for chaotic system [15], can asymptotically stabilize these quasi-periodic gaits to desired periodic ones. Especially for I-periodic gait, by using some additional adaptive controllers, the robot eventually hops without any control inputs, that is, complete passive running is obtained. Fig. 2 is an example of simulation results, which shows an adaptation control law achieves complete passive running. Fig. 3 is the stick pictures of high-speed passive running at 5 [ d s ] . Complete description and results can be found in the literature [141.\nC. Paper organization\nThe purpose of this paper it to extend the controller of passive one-legged hopper to biped robots. This was partially done on a 3D biped model in [17], where a rotor rotating around yaw-axis of torso was introduced. In this paper, we consider a highly nonlinear planar biped model having massive legs and torso and try to achieve stable periodic running gaits of it. Specifically, we derive dead-beat controller at flight phase based on the energy-preserving strategy to preserve mechanical energy at touchdown. Then, we combine some stance-phase controllers to get stable running gaits.\nThe paper is organized as follows: Section II introduces our new biped running robot and the equations of motion of simplified model are given. Section IU extends the controller described in Section I to a planar biped robot with torso, by introducing nonlinear decoupling control and target dynamics. Section N shows simulation results of biped running. Section V concludes this paper.\n11. A PLANAR BIPED ROBOT\nFigure 4 shows a CAD model of newly developed planar biped robot, Skipper//. The robot has two springy telescopic legs swinging around hip joints. Leg actuators are mounted parallel to the leg spring. Overall height of the robot is 0.75 [in] and the total weight is about 7 [kg]. The hardware design and experiments will be presented elsewhere.\nFigure 5 shows the definition of mathematical model. The generalized coordinates are defined as the position of center of gravity (COG), z = (z,,~,)~ E R2, the attitude of the torso, 4 E RI, and joint angles, $ = ( $ I , $ z ) ~ E RZ. A l b and m are the mass of torso and leg respectively. I and J are the moment of inertia about COG of the torso and COG of the leg respectively. All principal axes of each rigid part are coincident with their center axes. Table I shows the physical parameters, together with the values used in later simulations. This model is highly nonlinear because it has massive legs and torso, whose COG are located away the hip joint.\nAdditionally, the following assumptions are imposed on the model: (A) Mass of the foot (unsprung mass) is negligible (B) The foot does not bounce back, nor slip the ground\n(inelastic impulsive impact)" ] }, { "image_filename": "designv11_20_0000980_1-4020-3559-4-Figure27-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000980_1-4020-3559-4-Figure27-1.png", "caption": "Figure 27. The SAR antenna: a) half unfolded state b) folded antenna; c) multibody model.", "texts": [ " Each unfolding system has two degrees of freedom, driven individually by actuators located in joints A and B. In the first phase of the unfolding process the panel 3 is rolled out, around an axis normal to the main body, by a rotational spring-damper-actuator in joint A, while the panel 2 is held down by blocking the joints D and E. The second phase begins with the joint A blocked, next the panels 2 and 3 are swung out to the final position by a rotational spring-damped-actuator. The model used for one half of the folding antenna, schematically depicted Fig. 27, is composed of 12 bodies, 16 spherical joints and 3 revolute joints. The central panel is attached to the satellite, defined as body 1, which has much higher mass and inertia. The data for this antenna is reported in the work of Anantharamann [24]. In the first phase of the unfolding antenna, the rotational springdamped-actuator is applied in the revolute joint R3. For the second phase, the revolute joint R3 is blocked and the system is moved to the next equilibrium position by a spring-actuator-damped positioned in joint R1" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000507_2006-01-0582-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000507_2006-01-0582-Figure3-1.png", "caption": "Figure 3. Components and tri-axial force coordinates inside the tripod CV joint", "texts": [ " Experiments were repeated at different articulation angles, these being 0\u00b0 representing idealistic conditions with minimal internal friction and GAF, and 15\u00b0 representing aggressive conditions with significant CV joint friction and GAF. All CV joint friction experiments were performed under grease lubricated conditions with the typical CV joint grease (petroleum hydrocarbon plus solid EP additives). Using this set up, the in-situ three axis CV joint internal forces (designated Fx, Fy and Fz), the applied torque, and the axial thrust force were recorded during each experiment. The components and directions of the three force components measured with the tri-axial force sensor installed inside the CV joint are depicted in Figure 3. Force component Fx represents the normal force P and is directly related to the applied torque. Force components Fy and Fz represent the axial and vertical friction forces respectively, which are the source of the total combined friction force Q. is the CV joint articulation angle. In order to calculate the internal friction coefficient that is present under all CV joint conditions, one needs to find a universal equation to cover all articulation angles and rotational phase angles. This is accomplished by introducing a coordinate transformation matrix based on Euler angles representing articulation angle , and rotational phase angle . Using the defined coordinates of the tri-axial forces shown in Figure 3, one can obtain the combined rotational transformation matrix. By multiplying each individual transformation matrix, as depicted in Figure 4, in sequence, the following equation that relates the measured internal forces to the global forces in accordance with the housing coordinate is given by: Fz Fy Fx zF yF xF )90(sincos)90(cossin)90cos(]cos1[)90sin()90cos( sin)90cos(cossin)90sin( ]cos1[)90sin()90cos(sin)90sin()90(sincos)90(cos 22 22 (1) Where, the CV joint rotational phase angle is defined as shown Figure 5, from the housing end view", "5 sec it reaches the end position), and then moving all the way towards the outward position of the housing, passing through the center position (at t = 22.5 sec it reaches the furthest outward position) and then moves to the center position (at t = 30 sec one full cycle is completed). The exact same cycle is repeated, thus a total of 60 sec data were recorded. The force components inside the CV joint as measured with the tri-axial force transducer (Fx, Fy and Fz) show some variation with the CV joint stroke. The positive directions for these forces are also clearly indicated in Figure 3. Fx is the normal force P in calculating the friction coefficient, and shows fluctuation according to the direction of stroke motion. The Fy force shown in Figure 8 (b), which is the main source of friction, is nominally constant at the articulation angle =0 \u00b0, and fluctuating with an amplitude of about 200 N at =15 \u00b0. Interestingly, the force Fz shown in Figure 8 (c), which acts perpendicular to the CV joint shaft (and in the case of =0 \u00b0, perpendicular to the housing) is negative during the inward stroke, which means that the force is acting downwards, or in compression in regards to the spider assembly, while Fz is positive during the outward stroke" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001554_eej.20585-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001554_eej.20585-Figure3-1.png", "caption": "Fig. 3. Phasor diagrams at power generation (s12 < 0).", "texts": [], "surrounding_texts": [ "If r2 g , xl2 g , xm12, rc, and E . g12p under operating conditions can be identified, then I . 1p and I . 1n can be found from Figs. 2(a) and 2(b), and the electrical characteristics of the PMIG can be calculated as shown below. \u2022 Phase current \u2022 Output \u2022 Input \u2022 Efficiency" ] }, { "image_filename": "designv11_20_0003288_amr.308-310.1513-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003288_amr.308-310.1513-Figure5-1.png", "caption": "Fig. 5. Views of assembly position of guide rail and the brake block.", "texts": [ " The stress and deflection distribution over the brake block cannot be exactly predicted nor calculated. It can be examined experimentally. Guide rail and brake block. The car transfers its potential and kinetic energy to friction energy that cause to high impact forces and accelerations. T70 guide rail profile was used which is most commonly used in the car supporting of elevator systems. To avoid contact between guide rail and brake block, a gap about 2 mm was left on both left side and inner base (fig. 5). Using Abaqus software, element type C3D8R (an 8-node linear brick, reduced integration, hourglass control) were selected. The stresses obtained were in terms of Von-Misses. Brake block material. St42 steel brake block material used in the current system. The obtained nominal stress-strain graph of St42 steel block material was used to get elastic and plastic behaviour during the analysis. Essential material properties are: Modulus of elasticity E = 201x10 5 N/mm 2 , Poisson ratio v = 0.3 Density p = 7850 kg/m 3 Roller Material" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000367_pesc.2005.1581876-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000367_pesc.2005.1581876-Figure1-1.png", "caption": "Fig. 1. Schematic diagram of a three-phase synchronous machine with permanent magnet on the rotor.", "texts": [ " Simulation results of the electromagnetic torque produced by a brushless DC drive system, using conventional six step (without current control loop), are presented. These results are compared with those produced by the same system using the proposed model into a current open loop system and into a current closed loop system. Implementation of such control and practical results are being presented in another paper of this conference (PESC05). The model formulation considers the following hypothesis for the SM-PMSM viewed in figure 1 [6] [7]: \u2022 the machine is a symmetrical three phase machine; \u2022 negligible rotor reluctance variation; \u2022 there are no rotor and stator saturation over the operation range of the machine; \u2022 the iron cores of the stator and rotor are non-saturated over the operation range of the machine; The phase voltages of the machine are described by (1). 18070-7803-9033-4/05/$20.00 \u00a92005 IEEE. va vb vc = Ls Ms Ms Ms Ls Ms Ms Ms Ls d dt ia ib ic + Rs ia ib ic + ea eb ec + vn vn vn (1) where: va, vb, vc: stator phase voltages; vn: neutral connection voltage (normally not used in commercially available machines); ia, ib, ic: stator phase currents; ea, eb, ec: stator back-EMF waveforms, induced by kinetics of rotor magnets; Rs: stator phase resistance; Ls: self-inductance of the stator; Ms: mutual inductance between stator coils" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002984_iros.2010.5650918-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002984_iros.2010.5650918-Figure1-1.png", "caption": "Fig. 1. Spherical ultrasonic motor used in this study", "texts": [ " Fukaya is with Tokyo Metropolitan College of Industrial Technology, Tokyo, Japan. depicts the principle of SUSM and the classification of the torque control strategy of SUSM. Section III describes three kinds of position control methods based on the torque control strategy. Section IV demonstrates the experimental system and the experimental results. Finally, Section V provides the conclusions and future works. The SUSM used in this study consists of one spherical rotor and three ring-shaped stators. Figure 1 shows an overview of the SUSM. The geometric schemes are illustrated in Fig. 2. The stator includes a metallic elastic body and piezoelectric elements. When an AC voltage is applied to the piezoelectric vibrator, a standing wave is generated on the elastic body. By applying two AC voltages with a phase difference to the positive and negative sections of the piezoelectric elements, a traveling wave is generated due to combination 978-1-4244-6676-4/10/$25.00 \u00a92010 IEEE 3061 of the two standing waves[7]", " ci(fi) = 2 \u2016md\u2016 3 c\u03c6 | sin (\u03c8\u2032 \u2212 \u03b8i) | sin \u03c1i = sgn {sin (\u03c8\u2032 \u2212 \u03b8i)} (11) This type of torque control is referred as FR method. 3) Torque Control Based on Phase Difference and Frequency (HB): We control both the phase difference and the frequency with Eq. (8). That is, we control fi and \u03c1i using a following condition. ci(fi) = 2 \u2016md\u2016 3 c\u03c6 sin \u03c1i = sin(\u03c8\u2032 \u2212 \u03b8i) (12) This may be called a hybrid control[2] (in this paper, refer to HB method). A picture of the SUSM used in this study was shown in Fig. 1. Potentiometers shown in the figure measure the angle of the lever (\u03b8x, \u03b8y) via guide rails (see Fig. 4(b) and Fig. 5). \u03b8x and \u03b8y are the angles between the axis Z and the lever projected on the X-Z plane and the Y-Z plane, respectively. On the other hand, we express the position of lever as the posture of vector. The original position vector of the lever is set to k = [0, 0, 1]T when it corresponds to Z-axis. The position vector of the lever p is expressed as follows: p = Ry(\u03b8x)Rx(\u03b8\u2032y)k , (13) where, Rx(\u03b8) means the rotation matrix around the axix x through an angle of \u03b8" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000308_0094-114x(87)90058-9-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000308_0094-114x(87)90058-9-Figure4-1.png", "caption": "Fig. 4. The two solutions with ~2 ffi 30\u00b0 for the five-bar pair given in Fig. 2. (Dashed lines show the five-bar pair given in Fig. 2.)", "texts": [ " The admissable rotating range of the input link of a five-bar linkage is described in Fig. 3. For an arbitrarily selected//2 value, there exist 0, 2, 4, or 6 geometric inversions of the paired five bars which possess the above mentioned relationships. The angular displacements of links A3 and A~ is//3, of links A4 and A~ is fi4, of link A5 in the five-bar AI-A2-A3-A4-A5 is/~s, and of link A5 in the five-bar A~-A~-A~-A~-A5 is /~6. And the values of pj, for j = 3 to 6 are the solutions of the compatibility equations. Figure 4 shows the two solution pairs of the two five-bars given in Fig. 2 with ~2 = 30\u00b0. A more compact geometric expression of the compatibility equations is to connect the kinematic inversions of the two five-bar linkages to form a one-degree-of-freedom ten-bar linkage. The two kinematic inversions are taken with respect to links A 3 and A~; in other words A3 and A~ are considered now as fixed links. In order to connect the two five-bars of Fig. 2 to form a ten-bar as shown in Fig. 5, it is necessary to rotate linkage Ai-A~-A~-A~-As by an angle of/_(A3-A~)", " The ends of finks A4 and A~, which rotate identically, can be connected to form a parallel motion four-bar by adding an extra link EF as shown in Fig. 5. Similarly links Al and A~ can also be connected to form another parallel motion fourbar. The range of the freely chosen value of/~2 of the compatibility equation depends on the admissable moving range of link AKJ of the ten-bar linkage. Figure 6 shows the configurations of the ten-bar linkages at/~2 = 30\u00b0, and they correspond to the two five-bar pairs in Fig. 4. The number of the existed geometric inversions of the ten-bar linkage at each/~2 position is related to the number of branches of the ten-bar linkage. For each selected /~2 value, the relative positions between links AKJ and JHI are the same in all kinematic inversions of the ten-bar linkage. In other words the distances of HG, GA, and AH are constants in all kinematic inversions of the ten-bar linkage given in Fig. 6. Pivots H, G, and A of the ten-bar can then be considered as the three pivot points of a ternary link" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000590_1.1844991-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000590_1.1844991-Figure3-1.png", "caption": "FIG. 3. Example of a hysteresis trajectory. With a \u201cvirgin\u201d contact, the motion starts at point s0,0d along ysxd sor \u2212ys\u2212xdd. If the motion reverses at point 1, it will follow s1-2-5d. If underway it reverses at point 2, it follows s2-3-3\u2019d; if now there is a reversal at point 3, it follows s3-4d and follows again the s1-2-5d line to arrive at 5, points 2 and 3 can now be removed from the memory.", "texts": [ "77 On: Sat, 20 Dec 2014 10:50:57 last example will be called a mass on a hysteretic spring in the remainder of this text. The description here concerns the hysteresis element in the block diagram only, i.e., not the complete system, which will be treated in the next section. We recall ssee the Introductiond that hysteretic friction is function of the displacement only. The hysteretic friction is moreover not a unique function of the displacement but depends on the previous history of the movement. Thus referring to Fig. 3, if there has never been a relative movement before, the friction force follows a virgin curve fsxd. Rule I: HFfric = fsxd with fsxd = Hysxd if x \u00f9 0, \u2212 ys\u2212 xd if x \u00f8 0. Usually ysxd has the following properties: ys0d=0; y8s0d\u00f90 and y9sxd\u00f80 swhere the primes indicate derivation of the function with respect to its argumentd. The first two properties are universal, while the third implies that the hysteretic function is \u201cdissipative\u201d and softening for all displacements. If the motion direction changes at x=xm ssee Fig. 3d the friction force becomes Rule II: 5Ffric = Fm + 2fS x \u2212 xm 2 D , Fm = Ffricsxmd calculated with the formula for Ffric before the reversal. If the movement reverses again, before x=\u2212xm, the same reversal Rule II is used for the friction force. If, on the contrary, the direction of movement at x=\u2212xm has not reversed, the friction force follows Rule I again. Thus, when uxu becomes larger than the maximum of the absolute value of any reversal point, Rule I describes the friction force. If uxu is smaller than this value, the friction force is calculated from Rule II using for sFm, xmd the values of the last reversal point. When \u201can internal loop is closed\u201d ssee, for example, the hatched part in Fig. 3d the two last reversal points are not needed anymore, and are thus \u201cforgotten,\u201d and the curve based on the third last reversal point is followed. This is called the wiping out effect of hysteretic behavior.18 Based on these rules, the resulting friction force can be calculated for any input motion trajectory of the body. This kind of hysteresis is called \u201chysteresis with nonlocal memory,\u201d since every velocity reversal has to be remembered until an internal loop is closed. The rules sI and IId describing this kind of hysteresis are called \u201cmasing rules" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001864_s10778-007-0131-6-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001864_s10778-007-0131-6-Figure1-1.png", "caption": "Fig. 1", "texts": [ " Control of systems with nonholonomic constraints [8] is an important issue in developing mobile robots (see [11\u201315, etc.] and references therein). Of considerable interest are robots with several articulated wheeled platforms [1, 3\u20136]. An algorithm to set up equations of motion for such systems was proposed in [3]. In what follows, we will address, as in [1], the problem of synthesizing an algorithm to stabilize the motion of a wheeled vehicle with two components (driving and driven) along the OX-axis (Fig. 1). The problem will be solved in kinematic approximation. Deviations of the kinematic parameters of motion from their program values are not assumed small, i.e., the stabilization algorithm to be synthesized will be nonlinear. Let us outline a procedure to synthesize a robust controller. In contrast to [1], where the approach of [16] was used for this purpose, we will use the procedure from [2]. The efficiency of the procedure will be demonstrated with the example from [1]. 1. Description of the Model [1]. Consider a wheeled transport vehicle moving rectilinearly (along the OX-axis). It is schematized in Fig. 1. The vehicle has a driving component whose position is determined by the segment AB of length L1 and a driven component whose position is determined by the segment BD of length L2 . The velocity of the point B (VB ) is along AB, while the velocity of the point A (VA ) is determined by the angle > (the angle of the steerable wheel). The velocity of the point D (VD ) is along BD. The segments AB and BD make angles &1 and &2 , respectively, with the OX-axis. Denoting by x and \u00f3 the coordinates of the point B and considering thatV VD B cos( )& &1 2 , we write the following equations describing (in kinematic approximation) the motion of the vehicle: cosx VB &1, siny VB &1, & > 1 1 V L B tan , sin& & &2 2 1 2 V L B , (1" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.1-1.png", "caption": "Fig. 14.1 Derrick used in ancient Greece and Rome", "texts": [ " The available historical data indicate that cranes form the oldest class of construction machines invented by humans. Cranes were known in Greece as early as the Part B 1 4 .1 5th century BC. They were used for transporting structural elements vertically and horizontally and putting them in specified places, among other tasks, in the construction of magnificent temples. Greek cranes incorporated structural components such as toothed and worm gears, pulley blocks, rope drums, supports, and power units in the form of levers mounted directly on the hoisting winch\u2019s shaft, or various treadmills. Figure 14.1 shows a derrick commonly used in both ancient Greece and Rome, with a mast in the form of the letter \u201cA\u201d. The hoisted block is grabbed by a crampon. The winch is turned by means of two levers and the mast\u2019s inclination is adjusted by guy-ropes. As late as the second half of the 19th century, i. e., until the industrial revolution, cranes were driven by people or, less commonly, by animals. Treadmills, including treadwheels with steps to be climbed, or pole windlasses were used as the driving devices [14" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003271_s00170-011-3475-3-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003271_s00170-011-3475-3-Figure10-1.png", "caption": "Fig. 10 Cutting error of the tooth profile curve", "texts": [ " Therefore, a n angle which describes coordinates in plane XY of points on the involute curve is considered (see Figs. 2 and 6). Points G and H show casually two points on the involute curve (see Fig. 8), where angles =G and =H show the involute functions corresponding to these points. In the radial cutting method, the end mill linearly cuts interval of these casual points on the tooth profile curve according to increment value (am) in angle an (see Fig. 5). In this case, a cutting error occurs on the involute tooth profile (see Fig. 9). In this study, Fig. 10 illustrates how to determine the quantity of this cutting error. Point L in Fig. 10 is also a point on the involute tooth profile, but this point is assumed as the midpoint of points G and H. A perpendicular distance (LM) drawn to distance GH from this point is named as the quantity of the cutting error of the tooth profile curve. The following process steps are applied to obtain a mathematical expression about this distance (see Figs. 2 and 10): OH \u00bc rH \u00bc r b Cos an \u00fe am\u00f0 \u00de \u00f011\u00de OG \u00bc rG \u00bc rb Cosan \u00f012\u00de Fig. 7 A spur gear manufactured by the end mill Where, rG and rH are radius of points G and H on the involute curve", " 4), where interval 0\u2264an\u2264amax is described for angle an corresponding to each point on the involute curve. \u201cAlpham\u201d and \u201calphan\u201d in the figures are used instead of angle am and an, respectively (i.e. alpham=am and alphan=an). It is seen that the cutting errors on the tooth profile curve increase depending on increments of am in the figures above. Quantities of these errors further increase when an is equal to \u03b1max (see Figs. 12, 13 and 14). Because distance GH increases depending on the increment in the radius rG, ra and rH, quantity of the cutting errors increase (see Fig. 10). Besides, it is seen that the cutting errors also increase depending on increments in values N, f and m of the gear (see Figs. 13 and 14). For the increment in these values raises the dimensions such as rr, rb, rp, ro of the gear (see Fig. 4). These rise in the dimensions cause to increase interval GH, and this case increases the cutting errors of the tooth profile. As a result, it is seen that the increment or decrement in the length GH according to parameters such asN, f, m of the gear affects the cutting errors of the tooth profile" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000513_2004-01-1058-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000513_2004-01-1058-Figure1-1.png", "caption": "Figure 1: General Driveline Configurations", "texts": [ " As these systems have limitations, a tendency to use additional systems which help to overcome these limitations and to add additional features can be observed. Torque Management Devices like limited-slip differentials and on-demand couplings offer the required improvements. In particular, actively controllable TMD's complement brake based traction control and stability control systems and allow for further improvements. The following results from vehicle tests with various driveline configurations shall give an overview about the improvements TMD's can offer to vehicle traction, stability, safety and comfort. Figure 1 shows the general driveline configurations on the market. TMD's can be used in FWD, RWD as well as in AWD vehicles. In permanent 4WD vehicles TMD's are applicable as control units parallel to the differentials, i.e. as limited-slip differentials. In an on-demand AWD vehicle a TMD can also be used as \"hang-on\" coupling. Figure 2 shows principle sketches of these two 4WD configurations. In a permanent 4WD vehicle the gearbox output torque is fed into a transfer case with a center differential which distributes the torque with a static ratio (50:50 or unequal torque distribution) to the front and rear axles" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003320_gt2010-22058-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003320_gt2010-22058-Figure4-1.png", "caption": "FIG. 4 SCHEMATIC OF DYNAMIC TEST RIG", "texts": [ " The reaction forces are calculated by integrating the measured circumferential pressure distribution in two cavities: F r=F r 1F r 2 , F r 1,2=\u2212R t\u222b0 2 p1,2cos d F t=F t 1F t 2, F t 1,2=\u2212Rt\u222b0 2 p1,2sin d (5) The static method is based on the strong assumption of constant circumferential pressure distribution along the cavity length. Another disadvantage of this method is that the pressure distribution is measured on the stator. But pressure distributions on stator and rotor surface can differ significantly at certain operating conditions (high rotational speed or high swirl) [27]. Dynamic Test Rig Figure 4 shows a schematic of the dynamic test rig. A flexible shaft with diameter of 25 mm is supported by elliptical fluid-film bearings. The first natural frequency of the dynamic test rig rotor system is about 29 Hz. The low frequency value is necessary for the identification of rotordynamic coefficients by the dynamic method. The test rotor is fixed in the middle of the shaft with precise clamping bushings. The same testing assembly as in the static test rig is located at the midspan position. The approach to identify the stiffness and damping coefficients on the dynamic test rig is described firstly in [6] and called in this work the dynamic method" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003607_robot.2010.5509766-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003607_robot.2010.5509766-Figure6-1.png", "caption": "Fig. 6. Sketch of an area S on the Riemannian metric sphere (shaded area) and several subsets of it, as used in the proof in this section. Here S is depicted after a reflection across v, hence S+ v = S\u2212v , but S+ u \u2265 S\u2212u . S0 is the dark shaded area between the planes defined by u and v.", "texts": [ " Let us assume we have done this several times, and that Gv was the last transformation we applied. The reachable set at this point is, as we have just shown, symmetric with respect to v. This also means that S+ v = S\u2212v = Sv = 1 2 S. However, we can also write S in terms of its u-defined subsets, S = S+ u + S\u2212u . If either of the two parts is larger than half the sphere, then applying Gu will end up covering the whole sphere. If not, we have S+ u = Sv +S0, S\u2212u = Sv\u2212S0, where, without loss of generality, we assumed S+ u \u2265 S\u2212u . Figure 6 should help with seeing these relations. Applying Gu, as before, we get that the new S is such that Snew = 2max{S+ u ,S\u2212u }= 2Sv +2S0 = S +2S0, which shows an increase of the reachable area by 2S0 on every iteration. This, of course, does not apply to the last iteration which adds just enough area to finish covering the sphere, and to the first iteration: no iteration can add more than the area it starts with, and since the first iterations starts with S0 it can only add S0; after this, the reachable area is at least as big as 2S0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000779_j.jelechem.2006.07.007-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000779_j.jelechem.2006.07.007-Figure1-1.png", "caption": "Fig. 1. Schematic representation (A) and perpendicular cut (B) of the TETLC. Micrographs of the perpendicular cut of the TETLC (C) fabricated with a different number of toner masks: one (left) and two (right) toner masks.", "texts": [ " Reference and auxiliary electrodes were positioned in the reservoir of the outlet side. The twin electrodes were positioned next to the outlet side (1 mm) in order to minimize the uncompensated resistance. The final dimensions of the device are the following ones: micro channel height and consequently gap between the electrodes (12 lm); micro channel width (1000 lm) and, electrodes width (500 lm). With these dimensions, the volume between the electrodes was calculated to be 6 nL (12 \u00b7 1000 \u00b7 500 lm). Fig. 1A and B shows schematic representations of the structure fabricated by the proposed pro- cedure. Microscopic examination of the gap between both gold electrodes in the twin-electrode thin-layer cell was performed with a Cambridge scanning electron microscope (Stereoscan 440) (Fig. 1C). A silver layer was sputtered on the device prior to the microscopic analysis. In this model, we assume that the concentration distributions in the solution layer between the electrodes depend only on the distance from them. Given the small height of the solution layer if compared to the electrode area, one can conclude that this is a good approximation when there is no convective flow of solution from the reservoirs. We assume that the only mechanism available for the transport of electroactive species in the solution layer is diffusion, thus, the time evolution of the concentration distributions is given by oci ot \u00bc Di o 2ci ox2 \u00f01\u00de where ci is the concentration distribution of the ith chemical species being considered and Di is its diffusion coefficient", " The simulation domain was discretized in an irregular mesh of triangles and the Galerkin method was used to obtain the linear system of algebraic equations to be solved for the concentrations of all species at the nodes of the mesh. A complete description of this standard method of numerical analysis is out of the scope of this paper, thus the reader is referred to the excellent material available about this topic [25]. The sparse linear system obtained was solved by the generalized minimum residual (GMRES) method using the incomplete LU decomposition of the matrix as preconditioner. The simulation program was written in C++ and compiled with the GNU C compiler. Its source coded can be obtained by contacting us. Fig. 1C shows micrographs of structures fabricated with different gaps separating both gold electrodes (named W1 and W2), the ability to control the channel height being possible by increasing the thickness of the toner layer. The distance between both electrodes (d) (around 6 and 12 lm by using one and two toner masks, respectively) affects the collection efficiency value when the device operates in the generator\u2013collector mode, as it will be shown later in this article. To clearly understand the behavior of the TETLC when both electrodes are polarized, it is instructive to begin by studying a well-known system such as the ferri/ferrocyanide couple, which has relatively fast electron-transfer kinetics and is stable in both redox forms in aqueous medium" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001564_tie.2006.885468-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001564_tie.2006.885468-Figure1-1.png", "caption": "Fig. 1. Schematic of FD subsystem.", "texts": [ " The unique aspects of FDC, FDMM, and FDMet are the materials, the fabrication process, and the post processing. An FD machine basically consists of a material-deposition subsystem, which consists of one or multiple driver\u2013liquefier(s) (called head thereafter) attached to a carriage and their controllers, and a positioning subsystem, which consists of an X\u2013Y \u2013Z table and its controller that enables the carriage to move in the horizontal X\u2013Y plane and a fabrication platform to move in the Z direction (see Fig. 1) [5], [10]. The function of the head is to pump the spooled, solid (ceramic, metallic, fugitive) filament into the heated liquefier via a pair of dc servomotor-driven rollers and extrude the semisolid material through the nozzle onto the build platform following the designed tool path. Multiple types of materials to be deposited within a single layer will need multiple sets of heads with each head for each type of filament material. For a layer comprising multiple types of materials, different types of materials are deposited one after another", " Furthermore, the shearing effect of a nozzle may push the overfilled material to the outer and the inner directions of the part boundaries and cause large errors in the part\u2019s dimensions [9]. Hence, it is imperative to keep a uniform road width everywhere in a layer, which will remove overfills if the design parameters such as road width and offset are appropriately selected. In the next section, a coordination-control scheme that can ensure uniform road geometry and deposition accuracy in an FD-fabrication process is presented. Deposition subsystem consists of a filament-driving component and a liquefier-extrusion component (see Fig. 1). The former is a first-order electromechanical subsystem of a small time constant, while the latter is a first-order extrusion subsystem of a big time constant with a pure time delay [3]. The process of the liquefier-extrusion component can be further divided into a filament-extrusion process and a road-forming process. The reason to make this distinction is that the former is influenced by the filament-driving process but (almost) not by the dynamics of X\u2013Y table, while the latter is influenced by both" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001097_elps.200600081-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001097_elps.200600081-Figure2-1.png", "caption": "Figure 2. Schematic drawing of the microchip-CE with PAD. Channel width: 50 mm, channel depth: 50 mm, double-T arms: 4 mm long, double-T volume: 1.25 nL, separation channel length: 52 mm, solution reservoirs: 6 mm diameter, detection electrode: 25 mm diameter.", "texts": [ " BGEs were prepared by weighing the appropriate amount of sodium borate and adjusting the pH with 2 M NaOH. For pH measurements, a glass electrode and a digital pH meter (Orion 420A1, Thermo) were used. All chemicals were used as received without any further purification. A 99.95% gold wire (25 mm diameter, Alfa Aesar) was used as the detection electrode. All experiments were performed at room temperature (22 6 27C) and unless noted otherwise, the points and error bars on the plots represent the averages and SDs obtained for at least three measurements. A previously described design [53] was used (Fig. 2) for the experiments. Briefly, a master mold was produced using a clean 100 mm silicon wafer (Silicon) and SU-8 2035 negative photoresist [54]. For that reason, a layer of photoresist (50 mm) was spun using a spin coater (Laurell Technologies) by dispensing approximately 3 mL of photoresist onto the wafer. A spread cycle of 500 rpm for 10 s followed by 3000 rpm for 30 s was performed followed by two preexposure baking steps at 65 and 957C for 5 and 10 min, respectively. A digitally produced mask (3600 dpi) containing the channel pattern was then placed on the \u00a9 2006 WILEY-VCH Verlag GmbH & Co", " The detection electrode (25 mm) was placed in the interface between the separation channel using a perpendicular channel (30 mm), and according to a previously presented arrangement [53, 57\u201363]. The tight sealing between the Au wire and the PDMS chip avoids the solution from filling the electrode channel. The present design allows the isolation of the detector from the separation current through the end-column configuration. During the sample injection, potentials of 1500, 250, and 1500 V were applied to reservoirs S (sample), SW (sample waste), and B (buffer), respectively (see Fig. 2). In order to avoid Joule heating in the S\u2013B channel during the injection, a resistor (1 MO) was included in series with the chip. During the separation step, the potential applied to reservoir B was raised to 11200 V (or to the corresponding separation potential), while the potential applied to reservoir SW was changed to 1500 V. The waste reservoir (W) was always grounded. Electrical connections were made to the microfluidic devices with platinum electrodes placed into the reservoirs at the ends of each channel", " Therefore, 1200 V was selected as the optimum separation voltage at which the separation of the six phenolic acids was achieved in less than 3 min with a baseline noise of 0.7 6 0.2 nA (peak-to-peak). The injection time affects the sample volume and, therefore, also affects the peak current and peak shape. The effect of injection time was studied by varying the injection time from 1 to 20 s using the optimum analysis conditions (data not shown). A significant increase in the peak current was observed in the 1\u201310 s range. This can be explained considering that (using 5 mM borate, pH 10.0) the double-T injector (see Fig. 2) is filled in approximately 7 s. Further increments in injection time (.10 s) only increased the response slightly. For that reason, 10 s was selected as the optimum injection time. Using the optimum conditions for separation (5 mM borate, pH 10.0 as BGE, 1200 V as the separation potential, and 10 s as the injection time) and the optimized PAD waveform (see Table 1), the six phenolic acids were separated and detected within 200 s (Fig. 6). Under these conditions, linear relationships between concentration and the peak current were obtained for HPA, SA, FA, CA, \u00a9 2006 WILEY-VCH Verlag GmbH & Co" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003103_tim.2008.2009410-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003103_tim.2008.2009410-Figure6-1.png", "caption": "Fig. 6. (a) Diagram of the proposed measuring system and (b) the aluminum cuvette holder realized. The cuvette and the housing for the LED and monitor photodetector are visible.", "texts": [ " Due to the polycrystalline nature of the thin-film semiconductor, the cost only weakly scales with the chip\u2019s size (available larger than 15 mm2). The polycrystalline nature is also responsible for the superior UV-radiation hardness. The drawbacks in terms of slow response time (larger than 100 \u03bcs) and quantum efficiency (less than 20%) are not critical in sensing dialysis adequacy. Taking into account the design consideration reported in the previous section, a simple experimental setup was developed to validate the proposed approach. Fig. 6(a) shows a diagram of the proposed system. It consists of the optical source, the cuvette holder, and the monitor and measuring photodetectors. As discussed in the design considerations, to achieve the highest sensitivity, the optical source was realized by using the UVLED L255. At the wavelength emitted by this device, we can obtain the maximum sensitivity to the most interesting retained substances, i.e., urea, creatinine, potassium, and phosphate. Fig. 6(b) shows the realized aluminum cuvette holder. In the figure, the cuvette and the housing for the LED and monitor photodetector are visible. Thanks to the excellent thermal conductivity of the aluminum, the heat generated by the LED can easily be dissipated, thus reducing the operation temperature of the device and possibly improving its mean lifetime. The solution being studied was placed into a 10 \u00d7 10-mm UV cuvette (101QS, Hellma, U.S.). Monitor (PD0) and measuring (PD) photodetectors consist of two TiO2 Schottky-type thinfilm detectors (TW30SY, Roithner Lasertechnik, Austria)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000853_jsen.2005.850991-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000853_jsen.2005.850991-Figure1-1.png", "caption": "Fig. 1. Schematic of the optical-fiber sensor design showing the modified cladding region.", "texts": [ " Incorporation of Cu ions in the doped polypyrrole structure is studied. The best dopant, processing technique and substrate nature are selected and investigated for sensitivity to DMMP and its cross sensitivity to other gases, i.e., water vapor, ammonia and organic vapor like acetone. 1530-437X/$20.00 \u00a9 2005 IEEE The optical-fiber sensor developed in this work is based on the modified cladding or coating design [9], [10]. The passive cladding of the optical fiber is removed from a small section and is replaced by a chemo-chromic material as shown in Fig. 1 [11]. The refractive index is a complex quantity given by where complex refractive index of a material; real part of the refractive index; complex part related to the absorbance of the material. Any change in the real refractive index or the absorbance of the material due to the presence of the analyte, changes the transmission properties of the optical fiber. There can be three different modes of operation (cases 1\u20133 next) of the sensor depending upon the refractive index of the chemo chromic material ( ) relative to the refractive index of the core ( ) and the refractive index of the original cladding ( )", " Since the reported value of the polypyrrole film refractive index is 1.817 [9] which is much higher than the core refractive index 1.45 the sensor operates in the leaky mode, i.e., case 2. Case 1: In this case, since the refractive index of the modified cladding is equal to that of the original cladding, the wave-guiding conditions do not change in the modified region and the light has the same gaussian intensity profile throughout the fiber. There is an evanescent field penetrating in the modified region shown in Fig. 1, any change in the absorbance of the material will result in evanescent absorption which will lead to intensity modulation. Case 2: Since this condition does not satisfy total internal reflection at the interface between the core and the modified cladding, guided modes will be changed to leaky modes in the modified cladding as shown in Fig. 1. The boundary between the air and the modified cladding can support total internal reflection and some light is reflected back into the core leading to transmission of light through the modified region. Any change in the complex refractive index of the modified cladding due to the analyte can change the waveguide transmission conditions and can result in an intensity change. Case 3: In this case, the sensor operates in the partial leaky mode, i.e., the critical angle in the modified region is higher than the critical angle for the fiber, so some higher order modes will leak through the modified cladding and the lower order modes will continue as guided modes" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000919_robot.2006.1642075-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000919_robot.2006.1642075-Figure6-1.png", "caption": "Fig. 6. Example of extended uncrossing operation", "texts": [ " (11) Operation EUOIV is applicable to subsequences represented as follows: El- Kx-C u/l i - \u00b7 \u00b7 \u00b7 -Cl/u i - \u00b7 \u00b7 \u00b7, (12) \u00b7 \u00b7 \u00b7 -Cu/l i - \u00b7 \u00b7 \u00b7 -Cl/u i - Kx-Er. (13) Operation EAOIII is applicable to a subsequence represented as permutation of the following three subsequences: \u03b1 \u2032, \u03b2\u2032, and \u03b3 \u2032: \u03b1\u2032 : \u00b7 \u00b7 \u00b7 -Cu i/j -Kx-Cu j/i- \u00b7 \u00b7 \u00b7, (14) \u03b2\u2032 : \u00b7 \u00b7 \u00b7 -Cl/u j/k-Ky-C u/l k/j - \u00b7 \u00b7 \u00b7, (15) \u03b3\u2032 : \u00b7 \u00b7 \u00b7 -Cl i/k-Kz-Cl k/i- \u00b7 \u00b7 \u00b7. (16) Note that operation EUOI counterchanges only the location of all crossing points included in a subknot, while operations EUOII, EUOIV, and EAOIII do not change the crossing state of the subknot. Fig.6-(a) shows an example of a knot including a subknot. Its initial crossing state is described as El-C u\u2212 1 -Cl\u2212 2 -C l+ 3 -C u+ 4 -Cu\u2212 5 -Cl\u2212 1 -C u\u2212 2 -Cl\u2212 5 - Cu+ 6 -Cl+ 7 -C u+ 8 -Cl+ 6 -C u+ 7 -Cl+ 8 -C l+ 4 -C u+ 3 -Er. Subknot K1 cor- responds to Cu+ 6 -Cl+ 7 -C u+ 8 -Cl+ 6 -C u+ 7 -Cl+ 8 . If only uncrossing operations are considered, two operations UO IV can be applied to the initial state. Including extended uncrossing operations, operation EUOII also can be applied. Fig.6-(b) shows the crossing state after applying operation EUO II. B. Introduction of Arranging Operation V It is difficult to move a target segment including a subknot without changing the crossing state of the subknot. So, it is prefer to grasp and move segments which does not include any subknot to perform extended uncrossing operations. In this paper, the fifth operation is introduced, which is referred to as arranging operation V, denoted by AOV. Operation AOV permutes crossing point Cu/l i and subknot Kx as follows: \u00b7 \u00b7 \u00b7 -Cu/l i -Kx- \u00b7 \u00b7 \u00b7 \u21d2 \u00b7 \u00b7 \u00b7 -Kx-C u/l i - \u00b7 \u00b7 \u00b7", " (18) This subsequence is equivalent to that for operation UO I described by (1) because subknot Kx is negligible. This implies that operation EUOI corresponds to the combination of operation AOV and operation UOI. Moreover, in the latter case, the crossing state of subknot Kx is not changed. Operations EUOII, EUOIV, and EAOIII are also equivalent to the combination of operation AOV and operations UOII, UOIV, and AOIII, respectively. Fig.8 shows an example of combination of operations AOV and UOII. Its initial state is equivalent to that shown in Fig.6-(a). First, segments adjacent to crossing point Cu\u2212 5 are passed over subknot K1 by operation AOV. Then, a region to which operation UO II is applicable appears as shown in Fig.8-(b). After applying UOII, the crossing state becomes that after applying EUO II shown Fig.6-(b). Thus, we can unravel knots without moving subknots by applying operations I through V instead of extended operations I through IV. By introducing the subknot and operation AOV, the number of possible state transitions can be decreased. For example, the crossing state graph for untying a bowknot has 932 states and 4282 operations as mentioned before. The bowknot includes subknot K1, which corresponds to an overhand knot. Regarding it as a segment, a graph with 346 states and 1505 operations was derived", " When a vision system recognizes the current crossing state of a linear object by tracing segments of its binarized and thinned image, there may be some parts which can not be identified. If such part has two boundary segments, we can regard it as a subknot. Then, using five basic operations proposed in this paper, the object can be unraveled until it includes subknots alone even if the state of those subknots is unidentified. For example, let us assume that subknot K1 included in a knot shown in Fig.6-(a) can not be identified. Then, this knot can be incompletely unraveled as follows: El-C u\u2212 1 -Cl\u2212 2 -C l+ 3 -C u+ 4 -Cu\u2212 5 - -Cl\u2212 1 -C u\u2212 2 -Cl\u2212 5 -K1-C l+ 4 -C u+ 3 -Er \u21d2 El-C u\u2212 1 -Cl\u2212 2 -C l+ 3 -C u+ 4 -Cu\u2212 5 - -Cl\u2212 1 -C u\u2212 2 -K1-C l\u2212 5 -C l+ 4 -C u+ 3 -Er \u21d2 El-C u\u2212 1 -Cl\u2212 2 -C l+ 3 -C l\u2212 1 -C u\u2212 2 -K1-C u+ 3 -Er (19) \u21d2 El-C u\u2212 1 -Cl\u2212 2 -C l+ 3 -C l\u2212 1 -C u\u2212 2 -Cu+ 3 -K1-Er \u21d2 El-C u\u2212 1 -Cl\u2212 1 -K1-Er \u21d2 El-K1-Er, where operations AOV, UOII, AOV, UOII, and UOI are applied in this order. Let us define a subknot which can be unknotted completely by applying operations UOI, UOII, and/or AOIII as an untightenable subknot [5]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002213_13506501jet575-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002213_13506501jet575-Figure3-1.png", "caption": "Fig. 3 Physical and computational planes including the bearing", "texts": [ " 2, because of the complex flow geometry in the (x, y) plane, the governing equations are transformed into a simple computational domain, (\u03be , \u03b7), such that, the physical domain which is the region between two eccentric circles, is conformally mapped into a rectangular computational domain. The transformation between physical and computational planes can be performed in two steps. These transformations along with their transformation functions are shown in Fig. 2. For the bearing, a separate transformation function has to be used to map this region into the rectangle in the computational plane. Figure 3 shows this Proc. IMechE Vol. 223 Part J: J. Engineering Tribology JET575 \u00a9 IMechE 2009 at LAKEHEAD UNIV on March 12, 2015pij.sagepub.comDownloaded from transformation with its function. The detailed relationships of the transformation functions are given in the Appendix 2. From these transformation functions, the relations between physical and computational planes and thereby the transformed forms of the governing equations in the computational domain can be obtained. For example, U and V , which are the \u03be- and \u03b7-velocity components in the computational domain, can be written as U = (y\u03b7u + y\u03be v) J (16) V = (\u2212y\u03be u + y\u03b7v) J (17) In these equations, J is the Jacobian of transformation, which is calculated from the relation J = y2 \u03be + y2 \u03b7 (18) such that x\u03be = y\u03b7 and x\u03b7 = \u2212y\u03be , based on the Cauchy\u2013 Reimann relations that hold for analytic functions" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000645_j.dental.2005.04.025-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000645_j.dental.2005.04.025-Figure2-1.png", "caption": "Figure 2 The dimensions of 5 mm-curvature metal abutment.", "texts": [ " The prepared tooth was reproduced in a gold-silver-palladium alloy (12% CASTWELL M.C., GC Corporation, Tokyo, Japan), and 1 mm-curvature metal abutment was made (Fig. 1, left). The additional 1 mm-curvature metal abutments were produced, then 3, 5 mm-curvature metal abutments were respectively fabricated by further apical reducing buccolingualmargins of these 1 mm-curvature metal abutments and keeping the proximal margins untouched (Fig. 1, middle, right). The dimensions of 5 mm-curvature metal abutments are shown in Fig. 2. Wax patterns for metal ceramic crowns were fabricated directly on metal abutments with blue inlay wax (CROWN WAX, GC Corporation, Tokyo, Japan). Thirty wax patterns were fabricated (ten for each type of abutment). Wax patterns were examined under a stereoscopic zoom microscope (SMZ-1, Nikon Corporation, Tokyo, Japan) to ensure that there was no gap between wax pattern and metal abutment margins. After storage at room temperature for 24 h, the wax patterns were vacuum invested in phosphate-bonded investment (CERAVEST G, GC Corporation, Tokyo, Japan) according to the manufacturer\u2019s recommendation for the liquid: power ratio" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000724_physreve.69.011705-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000724_physreve.69.011705-Figure1-1.png", "caption": "FIG. 1. A grain boundary made of a network of parallel edge dislocations is present in a SmA liquid crystal confined in a Grandjean-Cano wedge. In the vicinity of the SmA-SmC transition, the shift of the transition temperature due to the local stress makes the edge dislocations visible under the optical microscope.", "texts": [ " For example, recent experiments @9,10# on the dynamics of oscillatory plastic flow @11,12# in smectic liquid crystals observed in layer-normal stress experiments have been interpreted as the helical instability of screw dislocations giving rise to edge dislocations @13,14#, but up to now no such instabilities have been directly shown. Edge dislocations in the smectics have, however, been directly observed under the microscope in various experimental situations: in lyotropic lamellar phases @15# or in ferrosmectics @16# between a spherical lens and a plane, in the neighborhood of the smectic A(SmA) to smectic C(SmC) transition @17# in a Grandjean-Cano wedge geometry. In the latter case, sketched in Fig. 1, the contrast of the dislocation is enhanced by the fact that the molecules are normal to the layers in the SmA phase, whereas they are tilted in the SmC phase. The local compression of the layers \u00a92004 The American Physical Society05-1 in the vicinity of a dislocation yields a local shift of the SmA-SmC transition. Under a polarizing microscope the SmC domains ~and thus the dislocations! appear clearly within the dark background of the homeotropic SmA domains. Note that all the reported observations of dislocations have concerned only motionless dislocations in samples at rest, except for some observations in free-standing smectic-A films @18,19#" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003679_0022-2569(69)90012-3-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003679_0022-2569(69)90012-3-Figure4-1.png", "caption": "Figure 4.", "texts": [ " Figure 3c shows a fractionated chain used as a mechanism, the familiar parallel action of the drafting machine. A S S E M B L Y L I N K W O R K PARENT LINKAGE KINEMATIC INDEX There are assemblies with different kinematic type properties during different parts of the cycle. Those which alter their characteristics when 2 revolute pairs come into coincidence [20] are classified without regard to this feature, which is a matter of performance rather than of type as considered here. Those in which the kinematic type itself alters are considered separately. An example of such a mechanism is shown in Fig. 4. It consists of a 4-bar linkage with a jointed coupler, the joint being locked until the input angle ~b reaches 30 \u00b0. The joint is then freed, and gearing between the 2 links hinged to the frame is engaged. The assembly is now a geared 5-bar linkage, and remains so until q~ = 60 \u00b0, when the first condition is restored. The freeing of the joint and the engagement of the gear-train are effected by, say, electro-mechanical devices not shown in the figure. In this case the assembly takes the kinematic index which first has a higher number in the sequence" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002017_iscas.2007.378433-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002017_iscas.2007.378433-Figure3-1.png", "caption": "Fig. 3 Microchip-type sensor Fabrication steps are shown in Figure 4 [6]. On a Pyrex glass wafer, polyimide was spin-coated and then cured. Platinum (Pt) was deposited and patterned by the liftoff technique and used as the auxiliary and working electrodes. Deposition and patterning of silver (Ag) was repeated for the reference electrode. The last step, the glucose enzyme was immobilized on the electrode by another way.", "texts": [ " The work electrode and the reference electrode are printed to the base of materials though printing technology. The working operations are very complicated, and having some processing property. These performances will influence directly to oxidation-reduction reaction. The silver plasma is selected to the work electrode, and the carbon plasma is selected to the reference electrode. The base of materials is high polymer, as follows Fig.2 [1]. Fig.2. The structure of electrode fabrication A microchip-type sensor has been proposed and designed as shown in Figure 3 [6]. III. USING FOLDED-CASCODE AMPLIFIER TO CONSIST TRANSDUCER A completely amperometric sensor test circuit diagram for the glucose detection is shown in Fig. 5. The glucose sensor detect the ionic within surface of the work electrode corresponding to the concentration of glucose and then the production current input to the instrument amplifier convert to the voltage signal in the test system. The proposed folded-cascode operational amplifier with high-Gain and high-impedance at input node is used in the transducer [7]-[9]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002478_jjap.46.4698-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002478_jjap.46.4698-Figure5-1.png", "caption": "Fig. 5. Operation principle of motor using upper mode vibration.", "texts": [ " 4698\u20134703 #2007 The Japan Society of Applied Physics mode-coupled vibrator can be driven to excite concurrently the longitudinal and flexural vibrational components even by undivided simple electrodes and a single-phase electric source. This is feature advantageous for constructing a small motor. The contact point of the vibrator is pressed on the rotary shaft by an external preload, then a friction force is created between them, and the surface of the shaft is pushed in the central and circumferential directions by the vertical and horizontal displacements uV and uH of the vibrator, respectively, at the contact point, as shown in Fig. 5. In the half cycle of rotation, the positive direction of uV increases the friction force; simultaneously, the shaft is entrained to the direction of uH. In another half cycle, uV and uH displace in the inverse direction, and the friction force is decreased and the contact point is slipping. Therefore, the shaft can be rotated only in a given direction. The direction of the horizontal displacement uH when the vibrator is stretched inversely between the upper and lower modes. Thus, if we switch the driving mode from the upper mode to the lower mode, the rotational direction will be inverted" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003587_1.3622200-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003587_1.3622200-Figure3-1.png", "caption": "FIG. 3. A profiled substrate was premachined using a CNC lathe.", "texts": [ " Impellers Laser consolidation is a material addition process that can directly build functional features on an existing component to form integrated structure without the need of welding or brazing. For example, laser consolidation can be used to build net-shape functional components on premachined substrates. An impeller shape (Fig. 2) was selected to demonstrate the capability of laser consolidation process using 5-axis motion. The impeller has a diameter of about 77 mm and height of about 26 mm. There are nine long blades and nine short blades uniformly distributed. A profiled substrate was premachined using a CNC lathe (Fig. 3) and was mounted on a 5-axis motion system using a simple designed fixture. Laser consolidation of IN-718 alloy was performed to build blades on the premachined base using the 5-axis motion system. Laser consolidation can be performed to complete building up a blade and then another one. However, the completed blade(s) will interfere with the laser beam and powder flow when building other blades. In order to avoid the interference issue, laser consolidation was conducted to build one layer at a time for each blade" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002844_jsfa.2740220209-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002844_jsfa.2740220209-Figure3-1.png", "caption": "FIG. 3. Effect o / p H on the hydrolysis of DL-BAPA by soyabean", "texts": [ " Freeze-drying caused 15-20% reduction of the activity. Storage a t 4\"c in the liquid form also caused rapid loss of activity. The enzymes were most stable in the liquid form at room temperature under strictly sterile conditions. Table I1 summarises the effect of certain conditions on the stability of soyabean protease 1. Similar results were obtained with proteases I1 and Il l . Effect of pH on the relative activity of protease I Maximum activity of protease I, using BAPA as substrate, was 'obtained in the pH 8 . 2 - 8 . 5 region (Fig. 3). Other proteolytic enzymes isolated from soyabeans exhibit their optimum pH values at pH 5.0.2-4 Thus, there is a distinct difference not only in substrate specificity, but also in optimum pH of activity between the enzymes described in this report and the previously isolated proteases. Effect of metal ions on protease I The change in relative activity of purified protease I in the presence of metal ions is shown in Table 111. Mg2+ and Ca2- caused approximately 25:/, increase in activity a t 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002261_j.mechmachtheory.2009.09.006-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002261_j.mechmachtheory.2009.09.006-Figure6-1.png", "caption": "Fig. 6. Boundary conditions.", "texts": [ " The calculations are performed both for spiral bevel gear and pinion. Their geometrical characteristics are presented in Table 1. A software has been developed in order to generate Gleason spiral bevel gears in CAD software. Not only the teeth are created, but also the support and axis. The FEM computations take into account all the surroundings by the use of parabolic tetrahedral elements. The boundary conditions are assumed to be close to the reality. For example, one side of the pinion\u2019s shaft is fixed, while the other side can rotate (Fig. 6). In order to calculate the displacements with the functions, it is necessary to establish the bending definition matrix \u00bd~A . To obtain this matrix, K points have to be fixed. These points must be regularly spaced on a tooth flank (Fig. 7). Seven points along the tooth width and five points along the tooth height are sufficient to obtain correct results. Finally, a total of thirty five points is necessarily to determine the bending definition matrix. Table 4 Maximal and mean errors for the calculation of displacements for the 2nd comparison" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002333_s11249-009-9512-9-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002333_s11249-009-9512-9-Figure3-1.png", "caption": "Fig. 3 Schematic diagram of a rolling contact fatigue test rig", "texts": [ " The thrust ball bearing, which has Grade 25 balls of 3.69 mm diameter, was inserted between two specimens. The upper specimen was rotated at 1840 rpm, and the lower specimen was fixed in the test rig. An axial load that produces a maximum Hertzian stress of 2720 MPa was imposed on the upper specimen. The bearing and the rolling contact specimens were immersed in SAE-30 lubrication oil, which was circulated through a 0.25 lm filtered-pump feed system at a rate of 56.8 cm3/min, while a test was run. Figure 3 shows a schematic diagram of the rolling contact fatigue test rig. It was reported that the experimental lives by this test rig show less than 10\u201317% differences with those by the Falex multi-specimen rolling fatigue tester [15]. The rolling contact fatigue test was monitored by an accelerometer connected to a vibration meter. The vibration signal, which is influenced by both the upper and the lower specimens, was acquired by the data acquisition system (Fig. 4). When the vibration level reached a threshold level, the test rig and the timer were stopped by a PLC reset" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000718_1.1829068-Figure12-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000718_1.1829068-Figure12-1.png", "caption": "Fig. 12 Kinematic similarity of the wobble gear to a planar gear train with a large planet \u201eS1 \u2026", "texts": [ " Using Table 1 with link 2 fixed and links 3 and 4 the input and output, respectively, the gear ratio is vo /v i512N2 /N4 (4) Assigning N4 to be N221 as is true with the wobble gear, the ratio reduces to 1/(12N2). This is the same input-output relation as the wobble gear, as pointed out near the beginning of this paper, so the wobble gear and the sun-ring-planet train are similar mechanisms, and Eq. ~4! represents the speed ratio of each. Eliminating link 1 of the sun-ring-planet train gives the more intuitive similarity relation illustrated in Fig. 12. The planar trains in Figs. 11 and 12 differ only in numbers of teeth and are topologically the same. It follows that the wobble gear is also related to the harmonic drive through their respective reduction ratio formulas, giving a similarity relation as shown in Fig. 13. Once again labeling the input and output as links 3 and 4, respectively, the two formulas for v4 /v3 are identical ~see Eq. ~3!!. Here link 2 functions as a planet link. For the wobble gear mechanism, the input is analogous to a wave generator in that it initiates contact between the two meshing links" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003905_ecc.2013.6669571-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003905_ecc.2013.6669571-Figure2-1.png", "caption": "Fig. 2. Axial piston hydraulic pump (left, swashplate design) as well as hydraulic motor (right, bent axis design).", "texts": [], "surrounding_texts": [ "Dynamic system modelling plays a dominant role in modern control technology. An accurate system model is the key to improving the overall system performance. The considered mechatronic system is divided into a hydraulic subsystem and a mechanical subsystem, which are coupled by the torque generated by the hydraulic motor [11]." ] }, { "image_filename": "designv11_20_0002509_icelmach.2008.4800231-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002509_icelmach.2008.4800231-Figure4-1.png", "caption": "Fig. 4. Chart of the magnetic flux density for rated-load operation", "texts": [], "surrounding_texts": [ "The boundary conditions concerns: \u2022 null value of the local magnetic flux density on the outer and inner contours of the stator and rotor magnetic cores, where the lines of the magnetic field are tangent; \u2022 anti-cyclic periodicity related to the state variable in correspondent points placed on the radial lines that close the computation domain, Fig. 1, because the computation domain contains one of the two machine poles. The initial values of the rotor speed and of the electromagnetic torque are the values of steady state no-load operation for the study of motor dynamics after the rated load apply and the values of steady state rated-load operation, for the study of DC braking dynamics." ] }, { "image_filename": "designv11_20_0003587_1.3622200-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003587_1.3622200-Figure5-1.png", "caption": "FIG. 5. Cross-section showing laser-consolidated IN-625 shell-based molds embedded with cooling channels for demonstration.", "texts": [ " A unique backfill material was developed, which has good compression strength and thermal conductivity, and its thermal expansion is comparable to that of the laser-consolidated shell.20 The backfilled shell structure can operate at relatively high working temperatures (above 350 C). By using the backfill method, complex cooling channels and heating elements can be embedded into a mold at desired locations to improve the functionality and productivity of the mold. The possibility of using the laser-consolidated shell structure to make mold inserts was also demonstrated. Figure 5 shows a cross-section of a laser-consolidated IN-625 shell-based mold showing the internal cooling channels, while Fig. 6 shows a multipiece mold inserts with embedded cooling channels. The surface finish of the as-consolidated IN-625 test-piece is about Ra\u00bc 1.5\u20131.8 lm. The LC IN-625 alloy shows good tensile properties.9 Along the horizontal direction (perpendicular to the build direction), the yield and tensile strengths of the LC IN-625 material are 518 and 797 MPa, respectively, while the elongation is about 31%" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000980_1-4020-3559-4-Figure28-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000980_1-4020-3559-4-Figure28-1.png", "caption": "Figure 28. First phase of the unfolding of the SAR antenna (rigid and flexible models).", "texts": [ " The central panel is attached to the satellite, defined as body 1, which has much higher mass and inertia. The data for this antenna is reported in the work of Anantharamann [24]. In the first phase of the unfolding antenna, the rotational springdamped-actuator is applied in the revolute joint R3. For the second phase, the revolute joint R3 is blocked and the system is moved to the next equilibrium position by a spring-actuator-damped positioned in joint R1. The unfolding processes for rigid and flexible models are shown in Fig. 28, only for its first phase. The different behavior between the rigid and the flexible models is noticeable in Fig. 28. Though not shown here, the rotational actuator moment responsible for the start of the unfolding is not correctly predicted by the rigid multibody model. Being a very light and flexible structure, the discrepancies, if not identified during the design stage, would lead to the failure of the unfolding process. For flexible multibody systems experiencing nonlinear geometric and material deformations, the equations of motion for a flexible body are given by Eq. (29). However, due to the time variance of all its coefficients, Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000586_j.ijmachtools.2005.01.003-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000586_j.ijmachtools.2005.01.003-Figure2-1.png", "caption": "Fig. 2. Consequence of eccentricity on estimated trajectory, caused by using angles defined in {M}.", "texts": [ " The angle, at which each reading is taken, is estimated from the total number of samples taken from beginning (feed in) to end (feed out). This angle is used to estimate the coordinates of the actual path. Unfortunately a data processing challenge occurs because the ball bar length readings are taken from the center of the workpiece ball, whereas the trajectory angles are defined relative to the actual trajectory center. In the presence of an eccentricity the workpiece ball and the trajectory center do not coincide which causes the fictitious ovalization. To visualize this issue, Fig. 2 shows a test conducted on a perfect machine in the presence of a setup eccentricity. The eccentricity is grossly exaggerated for visualization purposes only. It shows the position of the DBB at the ith reading. The DBB length reading ri is the distance between the workpiece ball center at OB and the tool ball center at position i, Pi. This reading can also be described as the instantaneous trajectory radius measured from the workpiece ball center reference frame which has its x-axis, i\u0302B passing through the center of the initial tool ball position on the test circle. This radius and its associated estimated travel angle are used to estimate the coordinates of the point on the path. fBgP0 Z ri\u00bdcos\u00f0fMgqPi \u00desin\u00f0fMgqPi \u00de T (1) Note that the pre-exponent {B} indicates the coordinates of Pi as observed from reference frame {B}. As can be seen in Fig. 2, an error occurs because fMgqPi is used instead of fBgqPi since only the former is known and can be calculated by Eq. (2). fMg qPi Z i n Q (2) where i is the reading number, n is the total number of readings on the test trajectory and Q is the total angle travelled during the test arc. In the case shown in Fig. 2, the point falls inside the circular trajectory. The resulting processed trajectory can be theoretically analyzed as follows. The ball bar reading ri is calculated using Eq. (3). r2 i Z \u00f0R cos fMg qi Ke cos a\u00de2 C \u00f0R sin fMgqi Ke sin a\u00de2 Z R2 Ce2 K2eR cos\u00f0fMgqi Ka\u00de ri Z ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R2 Ce2 K2eR cos\u00f0fMg q qi Ka\u00de (3) Assuming a small eccentricity e relative to the nominal radius R, a polar plot of ri as a function of fMgqPi exhibits an oval like shape", " at tZti the tool ball is located at an unknown distance li from the trajectory center on a line passing through this center (OM) and making an angle fMgqPi with i\u0302M , so its coordinates in {M} are defined as follows: x Z li cos\u00f0fMgqPi \u00de y Z li sin\u00f0fMgqPi \u00de (6) 2. the tool ball is somewhere on the periphery of a circle of radius equal to the DBB length (ri) centered on the workpiece ball at coordinates (u, v). The equation of this circle in {M} is written as follows: \u00f0x Ku\u00de2 C \u00f0y Kv\u00de2 Z r2 i (7) The position of the tool ball on the circular path is at the intersection of the line and circle just described as shown in Fig. 2. Substituting the coordinates of the tool ball from Eq. (6) into Eq. (7) leads to a quadratic equation in li which is solved to determine the intersection of the line with the circle. l2 i K2\u00f0u cos\u00f0fMgqPi \u00deCv sin\u00f0fMgqPi \u00de\u00deli C\u00f0u2 Cv2 Kr2 i \u00deZ0 (8) This equation yields two values of li but only the positive signed one is acceptable thus li is calculated by Eq. (9) li \u00bc \u00f0u cos\u00f0fMgqPi \u00de\u00fe v sin\u00f0fMgqPi \u00de\u00de \u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r2 i K\u00f0u sin\u00f0fMg qPi \u00deKv cos\u00f0fMg qPi \u00de\u00de2 q (9) where fMgqPi and ri are known but the eccentricity vector (u, v) must be estimated iteratively since it is initially unknown" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001843_detc2007-34010-Figure12-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001843_detc2007-34010-Figure12-1.png", "caption": "Figure 12. Gear set photo after superfinishing process", "texts": [], "surrounding_texts": [ "Hypoid gears are widely used in automotive industries to transfer rotation between non-intersecting axes in rear wheel drive and 4WD vehicles. Compared to other options for gear types (such as straight and spiral bevel gears), that geometrically are capable of transferring power between \u2217 Manager, Gear Design & Development, American Axle and Manufacturing nloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/01/2016 T perpendicular axes, hypoid gears have more advantages which allows this type of bevel gear to dominate in automotive axle applications. In general two basic different cutting processes are used to generate hypoid gears namely face-milling (FM, also called single indexing) and face-hobbing (FH, also called continuous indexing), which have their own advantages and disadvantages over each other. However, face-hobbing method is dominated in automotive industry applications mostly because they need shorter cutting time compared to face-milling method [2-4]. In hypoid gears due to having nonintersecting axes, a higher sliding velocity between contact surfaces exists; as a result, sliding friction is one of the main power loss sources in addition to rolling friction. Therefore, hypoid gears have considerably more mechanical power loss during gear mesh than intersecting types of bevel gears and as a result are less efficient than other types of bevel gears. In a study on gears surface finish effects on friction [5] by comparing frictional losses of conventionally ground (Ra =0.4 um) with superfinished (Ra=0.05 um) teeth, it was shown that with the same load and speed this surface finish improvement will decrease friction around 30 percent in addition to decreasing tooth surface temperature. Moreover, based on Xu\u2019s proposed model for hypoid gear efficiency prediction [6, 7] which uses an EHL model with contact data provided by a FEA based modeling software (9), (depending on lubricant temperature at inlet) a change in surface finish from Ra=0.2 um to Ra=0.6 um may decrease hypoid gear efficiency around 0.5 percent. As a result, improving surface finish can be one way to increase efficiency. In this study a set of measurements were done to see how superfinishing and lapping will change surface finish of hypoid gear sets. The aim of this study is to investigate the effects of superfinishing and lapping on surface finish of hypoid gears for to have an insight of effects of these processes on surface finish. Moreover, it will be experimentally shown how superfinishing and lapping may change transmission errors (up to first two harmonics for lapping and first harmonic for superfinishing). 1 Copyright \u00a9 2007 by ASME Inc., Email: massetj@aam.com. erms of Use: http://www.asme.org/about-asme/terms-of-use First, surface finish measurement procedure will be explained and then the results of measurements with more details of measuring procedure will be provided. In this study, superfinishing effects on surface finish and transmission errors will be explained as a complement of previous study by the authors of this paper. This study will not cover theoretical issues related to this phenomenon (effects of superfinishing and lapping on surface finish and transmission errors) at this step; the goal here is to discuss the issue experimentally and more experiments and theoretical studies in future will help investigating the superfinishing and lapping effects in more details. 2 SURFACE FINISH MEASUREMENT PROCEDURE To experimentally see the effects of lapping or superfinishing on surface finish of hypoid gears; surface finish measurements were performed on nine hypoid gear sets before and after lapping. All gear sets were the same and had 11.5 in outer diameter and their geometric parameters are as mentioned in table1. To measure surface finish a CNC form-measuring machine1 (figure 1), equipped with software called FormTracePack was used to analyze measured data to extract surface finish. Table 2 shows an example of data sheet of surface finish measurement with several surface finish parameters (i.e. Ra, Ry, RzDIN, etc) and settings. There are several measuring parameters which need to be set before beginning measurement that are mentioned in table 2. Machine is equipped with both pinion and gear fixtures (holders) in order to keep parts securely in place while measurements is being performed. The software on the machine is capable to remove surface curvature from data and calculate pure surface finish for curved surfaces. It should be mentioned here that all measurements were done with 0.8 mm sample length (length of taking data \u201ccut-off\u201d). Measuring surface finish quality in different location on gear and pinion shows that surface finish considerably varies in both lengthwise and profile directions. Therefore in order to have consistent surface finish data to compare results before and after lapping process; data should be taken from same location on flank meaning that lengthwise (from toe to heel) and profile (from top to root) location of measuring spot should be consistent for all measurements. Table 1. Hypoid gear set geometric parameters Geometric parameters Pinion Gear Number of teeth 11 41 Diametral Pitch --- 3.57\u201d Face width 2.13\u201d 1.78\u201d Pinion offset 2.00\u201d Shaft Angle 90 \u0652 Outer cone distance 5.36\u201d 6.46\u201d Pitch diameter --- 11.50\u201d Pitch angle 25D 23M 62D 50M Mean spiral angle 49D 59M 27D 38M Hand of spiral LH RH Generation type Generated Non-Generated Depthwise tooth taper FH 1 - Mitutoyo SV3000 Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/01/2016 Te In order to check surface finish variation on pinion flank a pinion surface was divided into 9 regions (three divisions from toe to heel and three divisions from top to root) and surface finish was measured in the middle of each region. The results shows that surface finish improves from top to middle and then get worse continuing further to root in profile direction. In addition, surface finish will improve from toe and heel toward center (in lengthwise direction). Although it may not be a general rule, it is consistent result for most of measured pinions. 3 LAPPING EFFECTS ON SURFACE FINISH AND TRANSMISSION ERRORS Lapping is one of the processes used for gear finishing. While for many types of gears grinding may also be (sometimes) economical, for bevel and hypoid gears still lapping is the 2 Copyright \u00a9 2007 by ASME rms of Use: http://www.asme.org/about-asme/terms-of-use D most applicable process and is used as an economically applicable procedure (except for some case in aerospace applications). It is also the aim of lapping to make surface smoother through increasing conjugacy between pinion and gear and hence reduction in level of noise [3, 4]. As for hypoid gears in automotive industries, due to large production volume, grinding with currently available technology of machines and procedures is hard to be used instead of lapping. The main advantage of lapping over grinding in large volume production lines is that lapping needs cheaper machines and shorter processing time [8]. Depending on hypoid gear geometry (specially the amount of offset), the sliding velocity and contact pressure will be changed during mesh cycle. As a result, sliding distance caused by the combination of sliding velocity and contact pressure on every contact point (or spot) results in surface wear. Therefore the complex physical quantity of sliding distance on each surface point forms a surface wear distribution over the gear flank. To experimentally measure how much lapping will effect on surface finish, some sets of experiments have been performed. In order to check lapping effects on surface finish, a set of measurements were performed to evaluate surface finish (namely Ra and RzDIN) on both gears and pinions. All measurements were done on same location in all gears and pinions (at the center of lengthwise and profile directions). For all gear sets the same lapping settings for lapping machine were used and all were lapped with a same abrasive (silicon carbide) lapping compound. Also lapping procedure was conducted under light brake load with about 10 N.m torque on gear shaft and pinion speed was kept at 2300 RPM. For all gears and pinions, measurements were performed on both drive and coast sides before and after lapping. The measurements results in each of these four sets i.e. PinionDrive, Pinion-Coast, Gear-Drive and Gear-Coast sides are shown in tables 3 and 4 respectively (all measurements were performed on the same tooth). Moreover, average and variation of each column of data is shown at the end of table 3 and 4. To have a graphical view of surface finish changes by lapping, the measurement results for gear drive and coast sides before and after lapping is drawn in figures 2 and 3 respectively. As it can be seen surface finish of all gears is higher (rougher) after lapping; compared to what they were before lapping. In addition, surface finish changes for drive and coast sides of pinion before and after lapping are showed in figures 4 and 5. In this graphs Ra was used however RzDIN also was measured (as they are mentioned in tables 3 and 4) and same trend has been observed. As for pinion, there are no consistent changes and lapping effects on surface finish varies from part to part. In addition, to see how other hypoid gear characteristics may be affected by lapping; transmission errors of the gear sets were measured (by Gleason SFT machine) and the measurement results are as table 5 for both drive and coast sides respectively. The results of table 5 graphically are shown in figures 6 to 9 for the first two harmonics for both drive and coast sides. As it can be seen in graphs, lapping decreased both harmonics for both drive and coast sides. ownloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/01/201 3 Copyright \u00a9 2007 by ASME 6 Terms of Use: http://www.asme.org/about-asme/terms-of-use Also, to see the effects of rolling2 of lapped pinions and gears on surface finish the roughness of five gear sets before and after rolling were measured and the results for Ra on both drive and coast sides of pinions and gears are as figure 10 part a and b respectively. Rolling was performed by SFT machine under 17 N.m brake load on gear shaft with 100 RPM for pinion speed and light weight oil (SAE 30 W) was used for lubrication for full hunting tooth cycle time. As it can be seen rolling gear sets together after lapping will improve surface finish slightly. Although these figures (10 (a) and (b)) are for Ra, surface finish improvement with the same trend were observed for RzDIN as well. 2 - This rolling here is due to single flank test. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/01/201 4 Copyright \u00a9 2007 by ASME 3 - All units for transmission errors are micro radians 6 Terms of Use: http://www.asme.org/about-asme/terms-of-use Do wnloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/01/2016 T 4 SUPERFINISHING EFFECTS ON SURFACE FINISH AND TRANSMISSION ERRORS Isotropic superfinishing (ISF) is an abrasive type of finishing process. It is a chemically accelerated vibratory finishing that has capability to smoothen surfaces with (Ra) < 3 u-inch. A smooth work surface is produced by simultaneously loading an abrasive stone against a rotating workpiece surface and oscillating (reciprocating) the stone [12]. To see how superfinishing will effect on surface finish of hypoid gears a set of hypoid gears after lapping was superfinished and the same surface finish measurements as it was explained in previous sections was performed on both pinion and gears. Figures 11 and 12 show the gear sets before and after superfinishing respectively. The results of surface finish measurements for the pinion and gear are as mentioned in table 6. In this table first row (marked by 1) shows surface finish before superfinishing, second row (marked by 2) is surface finish after superfinishing and third column (marked by 3) shows surface finish after rolling. This table one sample result for one gear set while these measurements were done on 8 gear sets and surface finish changes were completely consistent among all parts. As it can be seen, superfinishing significantly improved surface finishing quality and the surfaces after superfinishing is much smoother. However, after rolling (with the same rolling condition mentioned for rolling after lapping) smoothness of this superfinished gear set has been decreased. The measurement results after rolling the gear set together are in third row (marked by 3) of table 6 for pinion and gear (all measurements were performed on the same tooth). Moreover to graphically see surface finish changes before superfinishing, after superfinishing and after rolling; the results are shown in figures 13 and 14 (for both Ra and RzDIN) for pinion and gear for both drive and coast sides. Also, to see the effects of superfinishing process on transmission errors, single flank tests were done on those 8 gear sets and the results for drive and coast sides are as figure 15 part a and b respectively. As it can be seen, although superfinishing improves surface finish drastically it doesn\u2019t have effect on 1st harmonic transmission error. Moreover the results for SFT shows that superfinishing also do not have any considerable and consistent effect on 2nd and 3rd harmonics as well. To see an example of surface quality before and after superfinishing and after rolling, figure 16 shows all the steps in the same graph for pinion drive side. 5 Copyright \u00a9 2007 by ASME erms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/01/2016 6 Copyright \u00a9 2007 by ASME Terms of Use: http://www.asme.org/about-asme/terms-of-use" ] }, { "image_filename": "designv11_20_0002569_aim.2009.5230020-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002569_aim.2009.5230020-Figure5-1.png", "caption": "Fig. 5. Impact force reduction considering the leg\u2019s mechanism.", "texts": [ " On the other hand, the leg mechanism of WL-16RV consists of a parallel linkage mechanism called the Stewart Platform (see Fig. 4). Because it has a higher stiffness compared with a serial linkage mechanism, it is not sufficient to obtain a high compliance of the landing-foot only by changing a position gain value as mentioned above. Therefore, we realized a larger position following error by raising the foot\u2019s edge of the traveling direction and concentrating a landing-impact force to an actuator nearest to a contact area as shown in Fig. 5. As a result, we could obtain a higher compliance against ground reaction forces. C. Foot Speed Control after Detecting a Foot-Landing To detect a foot landing on a ground, we focus attention on the force data measured by a 6-axis force/torque sensor mounted on a foot. A foot-landing is detected by the differential value of the force data along the z axis. The advantage to use the differential value is an ignorable sensor drift and a fast landing detection. When detecting a foot-landing, the foot speed is changed to zero under the law of conservation of momentum" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002032_neco.2007.19.3.730-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002032_neco.2007.19.3.730-Figure10-1.png", "caption": "Figure 10: Biped model: mH = 10 kg, m = 5 kg, a = b = 0.5 m, \u03c6 = 0.03 rad.", "texts": [ " Figure 9A shows an acquired pendulum dynamics. A sinusoidal shape that represents the pendulum dynamics (see equation 4.5) was successfully acquired. Figure 9B illustrates the learning performance of swing-up policy with estimating pendulum dynamics and state variables. After 692 trials (average over 10 simulation runs), the swing-up policies and the state estimation policies were acquired with an initially unknown model. 4.2 Application to a Biped Model. Finally, the RLSE is applied to a biped model (see Figure 10). The task in this case for the biped robot is walking down a slope without falling over. A biped dynamics model introduced in Goswami, Thuilot, and Espiau (1996) was employed here. Although it is well known that this two-linked biped model can walk down the slope without any actuation, it is necessary to set proper initial angular velocities at each degree of freedom to generate this passive walking pattern. In this study, we try to acquire a controller that can make the biped robot walk even when the initial angular velocities are not suitable for passive walking" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002982_10402000802105448-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002982_10402000802105448-Figure1-1.png", "caption": "Fig. 1\u2014Schematic representation of Tribogyr.", "texts": [ " A new device has been built to study the tribological behavior of flange\u2013roller end contacts under conditions as close as possible to the actual ones. It means that high normal loads, entrainment speeds, and spin motions are to be imposed. This article will first describe the test rig (architecture and specimen geometry), the sensor\u2019s implementation, and a specific parameter used here to quantify the spin influence. In a second part, friction and temperature results will be shown. Shearing effects induced by spin are discussed in a final section. The test rig is a three-part apparatus, as shown in Fig. 1: two assemblies including the upper and the lower specimens, plus the frame (not drawn in Fig. 1). The frame supports the two assemblies through hydrostatic thrust and hydrodynamic bearings. The upper one includes the disc specimen and the oil supply system. The lower assembly includes the spherical end specimen, the loading system, and a tilting facility with respect to the x-axis. Note that in the present work, the contact takes place in the x-z plane. The main entrainment direction is the x-direction and the mean entrainment speed Ue can reach 10 m/s. The lower body rotation (\u03bb can vary between 1\u25e6 and 10\u25e6) allows a spinning motion superposed on the linear velocity (parallel to the x-direction), as shown in Figs" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002738_9780470549148.ch6-Figure6.19-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002738_9780470549148.ch6-Figure6.19-1.png", "caption": "Figure 6.19 (a) Biomechanical variables describing the instantaneous state of a given segment in which passive energy transfers may occur at the proximal and distal joint centers and active transfers through the muscles at the proximal and distal ends. (b) Power balance as calculated using the variables shown in (a). The passive power flow at the proximal end Pjp , and the distal end Pjd , combined with the active (muscle)", "texts": [ " Therefore, \u03c9j is replaced by (\u03c91 \u2212 \u03c92), Pm = Mj (\u03c91 \u2212 \u03c92) W (6.27) Thus, if \u03c91 and \u03c92 have the same polarity, the rate of transfer will be the lesser of the two power components. Examples are presented in Section 6.4.2 to demonstrate the calculation and to reinforce the sign convention used. Energy can enter or leave a segment at muscles and across joints at the proximal and distal ends. Passive transfer across the joint [Equation (6.9)] and active transfer plus absorption or generation [Equation (6.27)] must be calculated. Consider Figure 6.19a as the state of a given segment at any given point in time. The reaction forces and the velocities at the joint centers at the proximal and distal ends are shown plus the moments of force acting at the proximal and distal ends along with the segment angular velocity. The total energy of the segment Es as calculated by Equation (6.17) must also TABLE 6.2 Power Generation, Transfer, and Absorption Functions Directions of Segmental Description of Type of Angular Muscle Amount, Type, and Movement Contraction Velocities Function Direction of Power Both segments rotating in opposite directions Concentric Mechanical energy generation M\u03c91 generated to segment 1 M\u03c92 generated to segment 2(a) joint angle decreasing (b) joint angle increasing Eccentric Mechanical energy absorption M\u03c9 absorbed from segment 1 M\u03c92 absorbed from segment 2 Both segments rotating in some direction Concentric Mechanical energy generation and transfer M (\u03c91 \u2212 \u03c92) generated to segment 1 (a) joint angle decreasing (e", " \u03c92 > \u03c91) Eccentric Mechanical energy absorption and transfer M (\u03c92 \u2212 \u03c91) absorbed from segment 2 M\u03c91 transferred to segment 1 from 2 (c) joint angle constant (\u03c91 = \u03c92) Isometric (dynamic) Mechanical energy transfer M\u03c92 transferred from segment 2 to 1 One segment fixed (e g segment 1) Concentric Mechanical energy generation M\u03c92 generated to segment 2 (a) joint angle decreasing (\u03c91 = O1 \u03c92 > O) (b) joint angle increasing (\u03c91 = O1 \u03c92 > O) Eccentric Mechanical energy absorption M\u03c92 absorbed from segment 2 (c) joint angle constant (\u03c91 = \u03c92 = O) Isometric (static) No mechanical energy function Zero From Roberston and Winter (1980). (Reproduced by permission from J. Biomechanics .) power at the proximal end Pmp and the distal end Pmd , must equal the rate of change of energy of the segment dEs/ dt. be known. Figure 6.19b is the power balance for that segment, the arrows showing the directions where the powers are positive (energy entering the segment across the joint or through the tendons of the dominant muscles). If the force\u2014velocity or moment\u2014\u03c9 product turns out to be negative, this means that energy flow is leaving the segment. According to the law of conservation of energy, the rate of change of energy of the segment should equal the four power terms, dEs dt = Pjp + Pmp + Pjd + Pmd (6.28) A sample calculation for two adjacent segments is necessary to demonstrate the use of such power balances and also to demonstrate the importance of passive transfers across joints and across muscles as major mechanisms in the energetics of human movement", " (iii) What hip muscle group generates energy to assist the swinging of the lower limb? When is this energy generated? 5. Using Equation (6.10) (see Figure 6.8), calculate the passive rate of energy transfer across the following joints, and check your answers with Table A.7. From what segment to what segment is the energy flowing? (i) Ankle for frame 20. (ii) Ankle for frame 33. (iii) Ankle for frame 65. (iv) Knee for frame 2. (v) Knee for frame 20. (vi) Knee for frame 65. (vii) Hip for frame 2. (viii) Hip for frame 20. (ix) Hip for frame 67. 6. .(a) Using equations in Figure 6.19b, carry out a power balance for the foot segment for frame 20. (b) Repeat Problem 6(a) for the leg segment for frame 20. (c) Repeat Problem 6(a) for the leg segment for frame 65. (d) Repeat Problem 6(a) for the thigh segment for frame 63. 7. Muscles can transfer energy between adjacent segments when they are rotating in the same direction in space. Calculate the power transfer between the following segments, and indicate the direction of energy flow. Compare your answers with those listed in Table A" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002122_978-3-540-36253-1_5-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002122_978-3-540-36253-1_5-Figure2-1.png", "caption": "Fig. 2 General scheme of a surface plasmon resonance sensor", "texts": [ " Surface plasmon resonance biosensors exploit special electromagnetic waves\u2014 surface plasmon polaritons\u2014to probe interactions between an analyte in solution and a biomolecular recognition element immobilized on the SPR sensor surface. A surface plasmon wave can be described as a light-induced collective oscillation in electron density at the interface between a metal and a dielectric. At SPR, most incident photons are either absorbed or scattered at the metal/dielectric interface and, consequently, reflected light is greatly attenuated. The resonance wavelength and angle of incidence depend upon the permittivity of the metal and dielectric (Fig. 2 ). The general advantages of optical techniques involve the speed and reproducibility of the measurement, and the main drawback is the high cost of the apparatus. Most of the works based on SPR have been focused on the rapid identification and quantification of different groups of toxicant pollutants. Electrochemical transduction transforms the effect of the electrochemical interaction between an analyte and the electrode into a primary signal. Such effects may be stimulated electrically or may result in a spontaneous interaction at the zero-current condition" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001893_s00170-008-1850-5-Figure14-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001893_s00170-008-1850-5-Figure14-1.png", "caption": "Fig. 14 Central axis representation (a) and of the colinearity between vector sum R and minimum moment MA on central axis (b)", "texts": [ "1 Central axis determination It is well-known that, with any torsor, it is possible to associate a central axis (except the torsor of pure moment), which is the single object calculated starting from the six torsor components [8]. A torsor [ A ] O in a point O is composed of resultant forces R and the resulting moment MO: [A]O = { R, MO. (1) The central axis is the line defined classically by: OA = R \u2227 MO\u2223 \u2223R2 \u2223 \u2223 + \u03bbR, (2) where O is the point where the mechanical action torsor was moved (here, the tool tip) and A is the current point describing the central axis. OA is, thus, the vector associated with the bipoint [O, A] (Fig. 14). This line (Fig. 14a) corresponds to geometric points where the mechanical actions moment torsor is minimal. The central axis calculation consists in determining the points assembly (a line) where the torsor can be expressed according to a slide block (straight line direction) and the pure moment (or torque) [8]. The central axis is also the point where the resultant cutting force is colinear with the minimum mechanical moment (pure torque). The test results enable us to check for each point of measurement where the colinearity between the resultant cutting force R and moment MA calculated is related to the central axis (Fig. 14b). The meticulous examination of the six mechanical action torsor components shows that the forces and the moment average values are not null. For each measure point, the central axis is calculated in the stable (Fig. 15a) and unstable modes (Fig. 15b). In any rigour, the case ap = 2 mm should be described as quasistable movement because the vibrations exist but their amplitudes are very low\u2014of the order of micrometers\u2014 thus, quasi null compared to the other studied cases. Considering the cutting depth value ap = 3 mm, the recorded amplitude was 10 times more important" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002791_fskd.2010.5569106-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002791_fskd.2010.5569106-Figure1-1.png", "caption": "Fig. 1. Forces diagram of Linear Rolling Guide at level use", "texts": [ " Therefore, the guide of the CNC machine tool must have high guiding accuracy, high stiffness, high wearing resistance, all of which could assure machine tool not vibrating at high-speed feed or crawling at low-speed [3]. The linear rolling guide was a rolling guide emerged for many years, highlighted in non-gap and could be preloaded, was commonly used for high-end CNC machine tool especially for importing ones. The forces on the rolling guide related to the using condition. When used at level it main undertook the cutting force in the vertical contribute, the gravity of the table and work-piece. It showed in Fig. 1. Meanwhile, it needed to consider the number and the distribution form of the orbits and sliders, which affected the force on the guide. In addition, the guide trip length, running speed, the impacts from acceleration and deceleration also had greater effect to the life of the guide when it reversing. This work is sponsored by the National S&T Major Project (20092x04014-102-03), the Fundamental Research Funds for the Central Universities (SWJTU09CX019), and the School Fund of Southwest Jiaotong University (2008B13)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001284_1.2359472-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001284_1.2359472-Figure5-1.png", "caption": "Fig. 5 Double pinion arrangement\u2014orientation and coordinate frames", "texts": [ " But this theoretical calcuation is still extremely useful in highlighting the interactions beween the gear meshes and bearings and comparing the single inion arrangement to the double pinion arrangement. Double Pinion Arrangement. In the double pinion arrangeent, there are two pinions in the path from the sun gear to the ing gear. The pinion that meshes with the sun gear will be desgnated as SP and the pinion that meshes with the ring gear will be esignated RP. The included angles between the centers of the sun ear, SP, and RP are , , and as shown in Fig. 5. Two reference rames will be defined\u2014one for SP and one for RP. The XSP axis s aligned along the line joining the centers of the sun gear and SP 8 / Vol. 129, JANUARY 2007 om: http://mechanicaldesign.asmedigitalcollection.asme.org/pdfaccess.ash and the XRP axis is aligned along the line joining the centers of the ring gear and RP. The ZSP and ZRP axes project out from the plane of the paper. The planes of action at the three meshes are also shown in the figure. As before, consider the mesh contact forces to be concentrated loads at the center of the facewidth" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003451_jmes_jour_1967_009_028_02-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003451_jmes_jour_1967_009_028_02-Figure2-1.png", "caption": "Fig. 2", "texts": [ " If urn, w, are the mean pressure velocities, u,, the mean shear velocity and if the surface velocities of the bearing and the journal are given by V2 and V, respectively, the principle of continuity may be expressed as h2 ap - aP h2 a ~ p ( 1 2 + 0 * 2 9 6 R e ~ ' ~ ~ ) - -2z J - - - f - - f (Fig. 1) a a -\"h(urn+umc)l+~ (hwm)+(Vzy+ v 2 x tan 8- Vly) = 0 ax \"c For the case of 'almost' parallel surfaces tan 6 = 8. After substituting for u, and w,, the 'macroscopic' equation of continuity assumes the form Employing the notation of Fig. 2 one has cos TJ 2 1, sin TJ E 0 and h z c ( ~ + E cos 0). Since the tangential velocity of the bearing VZx = 0, JOURNAL MECHANICAL E N G I N E E R I N G SCIENCE Neglecting quantities of the order of C/R, one obtains -(hum,) = -+CW sin 0 , . . (8) a ax If the bearing is stationary, the second term on the right of equation (7) reduces to - Vly, and When equations (8) and (9) are substituted into the equation of continuity, the differential equation (10) in pressure is obtained: Vly = ICOS 8+e$sin 6 " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003654_urai.2011.6145931-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003654_urai.2011.6145931-Figure7-1.png", "caption": "Fig. 7. The arena", "texts": [ " Experiment data will be saved on NXT brick as text file and it will be transferred to PC after all experiments are finished. Robot used in this research has two ultrasonic sensors (to detect the obstacles), two light sensors (to detect the target) and two servo motors. NXT Brick behaves as \u201cbrain\u201d or controller for this robot. Figure 6. shows the robot. Arena that will be used in experiments have 3 different home positions and 1 target location (by using candle as light source). The general arena is shown in Fig. 7. Beside this arena, some simple structure of some obstacles and target will also be used in order to know characteristics of learning mechanism clearly. 5. Result and Discussion First experiment that will be done is to test robot\u2019s ability in solving autonomous navigation task. Given three different home positions, robot should avoid the obstacles and find the target. The result is shown on Fig. 8. From Fig. 8. it is obvious that robot with subsumption architecture can avoid the obstacle well. Robot also succeed to find the light source as target from three different home positions" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001806_10402000801918056-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001806_10402000801918056-Figure2-1.png", "caption": "Fig. 2\u2014A schematic drawing showing the position of camlobes on a camshaft.", "texts": [ " The tappet shim is free to rotate on a groove on top of the tappet. The centerline of the cam lobe is offset from the centerline of the tappet and the tappet shim to facilitate the rotation of both the tappet and the tappet shim, which reduces friction. The shims are made out of AISI 52100 steel having a hardness of 62 Rc and have 0.1\u2013 0.2 \u03bcm (4\u20138 \u03bcin) centerline average roughness. The camshafts were made out of induction-hardened chilled cast iron having a surface hardness of Rc 50. The camshaft consisted of three sets of camlobes at 120\u25e6 apart as shown in Fig. 2. Each set consisted of D ow nl oa de d by [ FU B er lin ] at 1 6: 10 0 8 M ay 2 01 5 two camlobes oriented the same way with respect to the camshaft. Figure 3 shows three different types of patterns produced by a diamond tool. They are (a) parallel lines whose cross section could be either V-groove or square groove, (b) concentric circles with V-grooves, and (c) spiral V-grooves. The groove details are shown in Table 1. The width of the contact patch between the cam lobe and the shim is about 80 \u03bcm at the nose of the camlobe where the spring force is the highest" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000754_j.aca.2004.06.062-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000754_j.aca.2004.06.062-Figure1-1.png", "caption": "Fig. 1. A photo and cartoon of osmium-horseradish peroxidase redox polymer electrode coated on a gold radial flow ring disk. Bar on the ring indicates the cut site of cartoon.", "texts": [ "25 ml/min and sampling intervals of 20 min were employed. After preperfusion for 2 h, four stable cortical dialysate samples were collected as basal levels, and the animal was subsequently perfused with 100 mM NMDA dissolved in Ringer\u2019s solution and adjusted at pH 7.0 with a small amount of sodium hydroxide solution. Amperometry was performed by using an amperometric controller LC-4C (BAS, Indiana, USA) and an analog/ digital converter DA-5 (BAS, Indiana). The surface of a 6 mm gold radial flow ring disk (Fig. 1) was cleaned by an ultrasonicator, and 2 ml of Os-HRP redox polymer solution was dropped on the surface. After drying at room temperature, Os-HRP-RDE was kept at 4 8C until use. The Os-HRPRDE was positioned in a PEEK radial flow cell holder with a 25 mm gasket. Ach and Ch were separated on a microbore column, and reacted with post-column immobilized enzyme reactor. The enzyme-generated hydrogen peroxide was detected by the Os-HRP-RDE, which was operated at 0 mV versus Ag/AgCl. A 3 mm platinum electrode (MF1000, BAS, Indiana) was positioned in the cell holder, and applied at 450 mV versus Ag/AgCl, because the applied potential at 450 mV instead of 650 mV reduces the back ground of the platinum electrode", " In our recent Ach detection method using Os-HRPcoated glassy carbon electrode attached with ChO/HRP precolumn to eliminate choline [4], the peak of Ach on HPLC-ECD was broad, then the detection limit was 10 fmol [4]. Another disadvantage of the Os-HRP-coated glassy carbon electrode is that about one half of the coated OsHRP gel was removed from the surface within a few days and the Os-HRP-coated glassy carbon electrode was unstable to detect a small amount of Ach in vivo microdialysate. In the present study, the HRP precolumn was not attached on the microbore column, and a thin Au-coated (100 nm in thickness) ring disk electrode (Fig. 1) was utilized as RDE instead of glassy carbon, because ring disk electrode is used as a monitor of redox reaction. In the present study it was utilized with proliferation of working electrode as alternated electrode instead of disk electrode (Fig. 1). Since Au binds to cysteine and cystine [15], the HRP containing them might be retained on the electrode surface for a few weeks. As shown in Fig. 2A, by the conventional platinum electrode, the current of 100 fmol/ 10 ml Ach decreased 13 8% (mean S.D.) from the first to 10th injection in a day. With this method, therefore, the measurement of Ach required to add an internal standard such as ethylhomocholine. The signal to noise (S/N) ratio of 100 fmol Ach was 20 5 folds indicating that the detection limit (over 2 of S/N ratio) is 10\u201320 fmol and similar data has been reported by Carter and Kehr [10]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002073_epe.2007.4417591-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002073_epe.2007.4417591-Figure1-1.png", "caption": "Fig. 1b: The simplified system studied", "texts": [ " To control the flux and the DC voltage at the rectifier output, two loops operate in parallel. The first one allows us to control the flux and leads to the d-axis current reference. The second one enables us to control the DC voltage at the rectifier output and yields the q-axis current reference. The proposed strategies are then simulated using MATLAB\u00ae-SIMULINK\u00ae package. The obtained results are presented and discussed. The system studied is constituted of a wind turbine, an induction generator, a rectifier and an inverter as shown in the Figure 1a. Control of the inverter idc io isabc\u2126 Autonomous load ichabc vchabc Induction generatorWind turbine Gear box Rectifier Inverter Control of the rectifier Vsabc Vdc Fig. 1a: The system studied The goal of the device is to provide a voltage of constant magnitude and frequency to the load connected to the rectifier even if the speed or the value of the load varies. This can be achieved mainly by the control of the DC bus voltage at a constant value as long as the wind power is sufficient. Then, the system studied can be simplified into the following device: The linear model of the induction machine is widely known and used. It yields results relatively accurate when the operating point studied is not so far from the conditions of the model parameter identification" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002449_j.mechatronics.2009.06.012-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002449_j.mechatronics.2009.06.012-Figure8-1.png", "caption": "Fig. 8. Spindle with displacement meter.", "texts": [ " This proposed displacement measuring system uses Foucault current to measure the actual displacement of the spindle thermal growth instead of the traditional method of a PT 100 thermal coupler. This system applying Foucault current allows the accurate measurement of the spindle displacement by using a proper amplifier to create a stable output voltage. The output voltage can be fed back to the PLC which compensates the axis position of the spindle housing. The sensor will be placed in front of the e vs. spindle speed. spindle as shown in Fig. 8. The reading from the sensor is the actual spindle growth whether it is equipped with cooling system. Fig. 9 shows a typical motorized spindle assembly drawing, the measuring sensor is placed in front of the spindle which is 110 mm away from the position where rotor and stator placed. The magnetic field will not interfere with the sensor. To get a perfect performance of cutting result, the compensation logic has been set to allow the spindle axis compensation of every 1 lm when the output voltage changes per \u00b10" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003641_eej.21132-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003641_eej.21132-Figure10-1.png", "caption": "Fig. 10. Cross sections of experimental machines.", "texts": [ " 7, the values of \u03b2 under the rated load conditions of the conventional SG and the PMaSG are 34\u00b0 and 25\u00b0, respectively. The load angle of the PMa-SG is smaller than that of the conventional SG. The output power of the PMa-SG is 23% larger than that of the conventional SG without increasing the field current. The results show that the PMs increase the rated terminal voltage and rated output power at the same field currents. In order to verify the reduction of magnetic saturation, a small-scale PMa-SG and a conventional SG of the same dimensions were built and used for measurements [13]. Figure 10 shows the cross sections of the experimental machines. Table 4 shows the specifications common to the two machines. The winding specifications are also common to them. The shapes of the pole tips of the PMa-SG are changed in order to add the PMs. 4.1 No-load characteristics Figure 11 shows the measured no-load saturation curves of the conventional SG and the PMa-SG. The unsaturated region for the PMa-SG is enlarged compared to the conventional SG. The terminal voltage of the PMa-SG has increased" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000738_cae.20083-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000738_cae.20083-Figure10-1.png", "caption": "Figure 10 A diagram of the mechanical structure of the prototype solar tracker system. [Color figure can be viewed in the online issue, which is available at www.interscience. wiley.com.]", "texts": [ " This system had two-axes for tracking the vertical and horizontal planes of where the light source was positioned. LabVIEW was utilized for the motion control and tracking of the light source. The basic components of this system are shown in Figures 10 and 11. This system has included two VEXTA stepper motors and drivers for horizontal and vertical planes tracking. These VEXTA stepper motors in use were identical to that described in the first project. Two solar panels were mounted on a plane parallel to a perspex sheet as shown in Figure 10. The photodiodes, which were used as light sensors were positioned in a special pyramid shape mounted onto the surface, adjacent to the two solar panels of the same perspex sheet. The sensors were arranged in such a way that the opposite pairs of the photodiodes were used for tracking both the horizontal and vertical position of the light source. The light sensing circuit was basically a current-to-voltage (I-to-V) converter using a low power LM158 operational amplifier. Since there were two planes (i" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002821_j.sna.2010.07.002-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002821_j.sna.2010.07.002-Figure4-1.png", "caption": "Fig. 4. Deformation of stator an", "texts": [ " The motion of the n U eutral plane of the stator with amplitude A then becomes: z(x\u0303, t) = A cos ( 2 x\u0303 \u2212 \u03c9t ) (1) where is the wave length and \u03c9 is the angular frequency. Eq. (1) uses the spatial fixed frame of reference (x\u0303, z\u0303). The problem can be formulated in a moving frame of reference by using the transformation x\u0303 = x + vwt with wave propagation velocity vw = (\u03c9/2 ). In this frame which moves with the wave crest, the normal displacement of each point on the neutral plane of stator Eq. (1) becomes time independent as follows [13]: Uz(x) = A cos ( 2 x ) (2) Fig. 4 shows the deformation of stator and frictional layer of roller. In RIUSM the rollers are coated with the frictional layer. The frictional layer contacts with both stator and rotor. The properties of this frictional layer are given in Table 1. d frictional layer of roller. frictional layer. w p e w U c here \u201ca\u201d is the distance between surface of stator and neutral lane of stator (Fig. 5). The frictional layer is described by a viscolastic foundation model (Fig. 4). Normal and tangential stiffness hich result from shear deformation Uk x and normal deformation k z of the frictional layer is obtained as follows: N = Pb Uk z = Eb h , cT = b Uk x = Eb h2(1 + ) (4) e spri w t w l m w a t i i f m i T n here E, , , P, h and b are modulus of elasticity, Poisson ratio, angential stress, normal stress, thickness of the frictional layer and idth of frictional layer, respectively. The equivalent mass per unite ength of frictional layer is obtained as follows: = 0.5 bh (5) here is the density of frictional layer" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001838_0094-114x(75)90072-5-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001838_0094-114x(75)90072-5-Figure3-1.png", "caption": "Figure 3. Poncelet's theorem.", "texts": [ " of p, whose sides are all tangents of the required curve g,, and an equal reflection polygon PP~P~.. . of p2 is circumscribed to the corresponding curve g2. Now a well-known theorem, due to J. V. Poncelet (1822), states that any reflection polygon of a conic is circumscribed to a confocal conic (which may degenerate to the pair of focal points). This theorem, usually derived by means of the projective theory of confocal conic systems[l, section 30], will be proved by elementary methods. To find the tangents from a point P to a given ellipse g, the following construction is commonly used (Fig. 3): Draw an auxiliary circle h with center P and passing through one of the foci E, F of g, say E, and cut it by a second circle q with center F, having the major axis 2a of g as radius; the two intersection points Ql, Q2 then determine the required tangents t,, t: as the bisectors of the angles EPQI and EPQ2, respectively. If the points Q, and Q:, situated symmetrically to the line FP, coincide in Q on FP, the bisector of the angle EPQ gives the tangent t of the confocal ellipse p which passes through P" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001213_icar.2005.1507428-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001213_icar.2005.1507428-Figure7-1.png", "caption": "Figure 7: Thrust-vectoring is achieved by servoing the baffle angle (top). Theoretical and experimental results of dynamic thrust versus baffle angle (below).", "texts": [ " A mold was created and the mass of the resulting foam nacelle was 178 grams. As this exceeded the allotted weight budget a carbon \u00afber nacelle was fabricated. Although this was a more involved process, the resulting nacelle was durable, rigid, then and an acceptable mass of 74 grams. As discussed in Section 2, counter-rotating tandem rotors will conserve angular momentum. However, any differences between the two motors will yield unwanted vehicle yawing. Control surfaces, called baffles can compensate for such differences (see Figure 7 top). Figure 6: Given two 7 \u00a3 4 airfoils, the nacelle should measure 17 \u00a3 7 \u00a3 6 inches long, wide and high respectively, to suitable encase the airfoils, avionics and actuators. Changes in baffle angle enable thrust to be vectored and prevent yawing. The baffles also allow the vehicle to remain hovering despite sudden gusts. The baffles and motor-propeller combination were mounted on the test rig and dynamic thrust data was collected. Figure 7 (bottom) shows that theoretical calculations (box data points) yield conservative values of thrust. Wind tunnel and test rig tests (diamond and triangle data points respectively) yield actual dynamic thrust values. Rotorcraft control is often a challenging problem because the longitudinal and lateral flight dynamics are tightly coupled. While a tandem rotor con\u00afguration does not completely eliminate this coupling, control is greatly simpli\u00afed. The vehicle was mounted on the test rig and a voltage step input was applied" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000727_j.matlet.2004.12.019-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000727_j.matlet.2004.12.019-Figure1-1.png", "caption": "Fig. 1. Modified e-beam evaporator comprising the co-evaporation system.", "texts": [ " In depositing a NiTi film onto a silicon wafer, the titanium is evaporated using an electron beam whereas the nickel is thermally evaporated. In a typical e-beam system utilizing a single source, the crystal that monitors film thickness is used as the feedback signal for the control system. This signal controls the power output to the e-beam gun, which in turn changes the material temperature in the crucible and therefore the deposition rate. The co-evaporation setup described here is created by modifying such a typical e-beam system with the addition of a thermal evaporation heater. The modified e-beam system is shown in Fig. 1 [7]. As can be seen from Fig. 1, expected deposition rates and profiles will differ from the two sources. Deposition rates sensed at the crystal monitor would require different ratio corrections for the two crucibles. In our case, the crystal monitor has been calibrated to reflect the average rate at the substrate from the e-beam deposition system. The e-beam current is controlled through an Inficon crystal monitor, maintaining a given rate at the crystal. 100 mm (111) silicon wafers were cleaned baked and prepared for deposition", " The co-evaporation process to produce a NiTi SMA film proceeds as follows: with the shutter of the system in the closed position, the thermal evaporation of the nickel is started first and allowed to reach steady state. At this point, the crystal monitor measures the deposition rate of the thermally evaporated nickel only. Once steady state has been reached, the e-beam system is turned on for the deposition of titanium. In order for the deposited NiTi film to exhibit SMA behavior, a mass composition of roughly 45:55 wt.% titanium to nickel is required [8]. Though the crystal monitor of the e-beam system is used to acquire a deposition rate, usually in 2/s, the crystal itself actually senses mass. As seen in Fig. 1, the evaporation crucible is farther from the wafer in comparison to the e-beam crucible. The set point for the deposition rate of the ebeam system is adjusted to be twice that measured for thermal evaporation of the nickel. As discussed earlier, due to the geometry of the deposition system, equal mass of nickel and titanium will result in a range of concentrations of nickel to titanium across the substrate starting on the left side slightly nickel poor and increasing in nickel content across the wafer" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001213_icar.2005.1507428-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001213_icar.2005.1507428-Figure2-1.png", "caption": "Figure 2: Counter-rotating rotors conserve angular momentum (top). By shrouding the rotors in a nacelle, the vehicle can sustain slight collisions. While using two rotors does increase vehicle size, design trades in body material and propulsion can still result in a backpackable unit that measures 17-inches long, 8-inches wide and 4-inches tall (bottom).", "texts": [ " Section 2 describes the design thresholds, constraints and footprint justifying a tandem rotor con\u00afguration; Section 3 details the construction of a test rig to collect dynamic thrust data; Section 4 highlights the fabrication of the vehicle\u2019s carbon-\u00afber body; Section 5 presents initial flight control and stability tests; Flight tests and conclusions are given in Section 6. The Defense Advanced Research Projects Agency (DARPA) has been a long time advocate of MAVs. Their MAV Industry Briefs provide design thresholds and objectives that capture desired attributes. Examples for a Class 1 aircraft include \u00aftting in a backpack, weighing one pound (450-grams) or less, airlifting a half-pound or more, and sustaining light bumps into obstacles like walls. With this as a guideline, design trades were performed resulting in a shrouded tandem-wing con\u00afguration (see Figure 2). A preliminary design review for a vehicle weighing less than 1- pound (450-gram) yielded a weight budget as shown in Table 1. Rotorcraft, like conventional helicopters, spin an airfoil to achieve lift. The angular momentum that is generated results in a counter-rotation. Often a tail rotor is used to oppose such rotation. Other methods include using two rotors that rotate in opposite direc- tions. This tandem rotor con\u00afguration is often used in heavy-lift aircraft like the Chinook helicopter where rotors are mounted at the ends of the fuselage" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002828_1.3601557-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002828_1.3601557-Figure1-1.png", "caption": "Fig. 1 Spinning-frict ion appara tus", "texts": [ " For further increases in stress the coefficient of friction tended to increase. Reichenbach attributed the higher values of friction at the lower stresses to inaccuracies in the test apparatus. An experimental apparatus has been designed and constructed at the NASA Lewis Research Center that simulates the spinning in a ball-race contact and is capable of measuring spinning torque to stress levels greater than 400,000-psi maximum Hertz stress, spinning speeds to 3500 rpm and under varying contact configurations and conditions [2]. Fig. 1 is an isometric view of this apparatus. Tests were conducted with upper test specimens of ' / i-in-dia SAE 52100 steel, grade 25 balls of Rockwell C hardness 63, rotated against a lower test specimen having a cylindrical groove of radius 0.255 in. simulating a 51 percent ball-race conformity. (The lower specimen is supported in a frictionless air bearing housing.) Test conditions were at maximum Hertz stresses to 136,000 psi, a spinning speed of 950 rpm, and ambient temperature. Three lubricants were used" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000426_elan.200302850-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000426_elan.200302850-Figure3-1.png", "caption": "Fig. 3. Schematic diagram of the flow system used for evaluation of the carbon paste electrode modified with spinel-type manganese oxide for lithium determination. P: peristaltic pump; I: manual injector; S: sample or reference solutions; L: sample volume; C: carrier solution; EFC: electrochemical flow cell; R: voltammeter and potentiometer (recorder); W: waste.", "texts": [ " 2B), consisting of a plastic cylindrical tube (o.d. 7 mm, i.d. 4 mm) equipped with a stainless steel staff serving as an external electric contact. Appropriate packing was achieved by pressing the electrode surface (surface area of 12.6 mm2) against a filter paper. Before the use, the electrode was activated in a 0.01 mol L 1 lithium ions in 0.1 mol L 1 Tris buffer solution (pH 8.3) by cyclic voltammetry. Theelectrochemical cellwas inserted into anone-channel flow injection system schematically shown in Figure 3. The system was assembled with a peristaltic pump (Ismatec, model 7618- 40, Switzerland) and a manual injector made of Perspex with two fixed sidebars and a sliding central bar [26]. The manifold was constructed with polyethylene tubing (0.76 mm i.d.). The 0.1 mol L 1 Tris buffer solutionwas used as the carrier (C) at a flow rate of 5.0 mL min 1. The lithium(I) reference in 0.1 mol L 1 Tris buffer solution contained in the sample volume (L, 408.6 L) was injected and transported by the carrier stream after the baseline had reached a steady-state value" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001969_s00216-008-1847-9-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001969_s00216-008-1847-9-Figure1-1.png", "caption": "Fig. 1 The enzyme sensor system. 1 carrier reservoir, 2 air, 3 sample injection port, 4 enzyme sensor, 5 microtube pump, 6 excitation lightemitting diode, 7 spectrometer, 8 personal computer, A fiberoptic probe tip composed of a ruthenium complex, a oxygen optical fiber probe, b Parafilm, c ruthenium complex, d immobilized enzyme membrane, e flow cell, f buffer solution", "texts": [ " The LyOx and AWP mixture prepared was evenly applied and rubbed into the dialysis membrane. After drying for 1 h in the dark, the membrane was placed under a fluorescent lamp for 15 min to prepare an immobilized enzyme membrane. AWP is a photosensitive polymer that hardens when exposed to light, immobilizing the enzyme. The resulting membrane was cut into 4 mm \u00d7 4 mm strips, and then closely attached onto the detection portion of a fiberoptic oxygen probe (Foxy-R-Flat, Ocean Optics, Dunedin, FL, USA) using Parafilm. The fiberoptic probe tip was composed of a ruthenium complex (Fig. 1). As a reference, commercially prepared LyOx, purified from Trichoderma viride (0.6 U), was also immobilized with AWP using the same method. Apparatus and assay procedure The enzyme sensor system was used in a flow-injection measurement system (Fig. 1). An excitation light-emitting diode (LED) light source (USB-LS-450, Ocean Optics), a spectrometer (UBS2000, Ocean Optics), and a personal computer (Dimension 4000, Dell, Round Rock, TX, USA) were connected to the enzyme sensor. Then, 0.1 M PB (pH 5.9\u20138.0) from the carrier reservoir was transferred to the flow cell equipped with the enzyme sensor. When light (475 nm) from the LED excites the ruthenium complex at the probe tip, the excited ruthenium complex fluoresces, emitting energy at 600 nm. If the excited ruthenium complex encounters an oxygen molecule, excess energy is transferred to the oxygen molecule in a nonradioactive transfer, decreasing or quenching the fluorescence signal" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000759_imece2004-60714-Figure11-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000759_imece2004-60714-Figure11-1.png", "caption": "Figure 11. Mesh of a finite element model for bending actuator.", "texts": [ "2, 149.2, 118 kJ/m3 respectively. Case (B) has the highest energy density. High elastic modulus will have advantage for high energy density. It normally requires higher osmotic pressure to achieve the same amount of deformation. However the upper limit of osmotic pressure is reported to be around 10-15 MPa. Thus, reducing the shell thickness might achieve larger stroke for the same osmotic pressure and thus achieve higher energy density. The bending actuator consists of multiple rows of vesicles. Figure 11 shows a simple configuration of two rows of vesicles. In order to achieve bending deformation, osmotic pressure is only applied to one row of vesicles. For example, when pressure is applied to the lower row, the lower part of the actuator try to expand and make the whole bulk material to bend upward. A typical bending deformation is simulated in Figure 12. In the simulation, all bottom row of vesicles are subjected an osmotic pressure of 10MPa. 7 Downloaded From: http://proceedings.asmedigitalcollection", " The bending actuator is still interesting because the de- Copyright c\u00a9 2004 by ASME url=/data/conferences/imece2004/71571/ on 05/09/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use formation will be amplified significantly when a long beam of actuator is considered. Figure 13 shows deformation of two rows of 30 vesicles under different pressure configurations. In the simulation, a two dimensional model is used and the actuator is modeled as a plain stress problem in order to reduce the computation effort for the complex 3-D model. First the 2-D model and the 3-D model are both applied to the configuration in Figure 11 with five columns of vesicles in order to verify the accuracy of the 2-D model. The results shows both two models consists well and the error of deformation is less than 2%. Then we apply the 2-D model to the long beam which is stacked with six of the five column blocks for a total of thirty columns of vesicles. In all the simulations in Figure 13, the five columns of vesicles in each single block are applied the same pressure pattern . One block can be chosen to either have pressure in the upper row or in the lower row" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003015_oceanssyd.2010.5603906-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003015_oceanssyd.2010.5603906-Figure4-1.png", "caption": "Fig. 4. The main forces acting on the moving float. The float is assumed to be ascending and therefore the buoyant force is greater than the weight of the float. As the float moves, it is also subject to a quadratic drag force. Additionally, an added mass term is considered, which accounts for the additional inertia provided by parcels of water working against the accelerating float.", "texts": [ " Velocity input tests produced much smoother acceleration profiles and simpler linear control restrictions. Although velocity control is not as intuitive, its advantages in terms of power consumption and linear control are significant. Therefore, the control system design utilizes velocity input control. The system dynamics provide a foundation for specifying all system components in order to ensure the Lagrangian float\u2019s physics are matched with the onboard Auto-Ballasting System (ABS). The main forces acting on the float include buoyancy, gravity, and drag as shown in Fig. 4. Added mass, which affects accelerating objects underwater, is also considered. This term is not actually derived, but is found empirically. The equation of motion considers these factors, and the influence of the ABS, which uses a piston to pull water into or out of the system, as illustrated in Fig. 2. By changing the volume of water contained within the float, the float\u2019s volume actually remains the same. Realistically, the float\u2019s mass is the only variable changing because the water pulled into the float must move with the float as well" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure3-1.png", "caption": "Fig. 3. Curvature characters.", "texts": [ " That is to say, the tangent direction of generating circle and its perpendicular direction are two principal direction of R\u00f01\u00deb , and their corresponding principal curvatures k\u00f01\u00deb1 and k\u00f01\u00deb2 become important invariant describing the shape of the surface. Prescribing that the direction, from the generating circle to its center, is positive direction of the surface normal. Obviously, principal curvature k\u00f01\u00deb2 \u00bc 1=r is constant, and k\u00f01\u00deb1 is a primary invariant determining the shape of R\u00f01\u00deb . In order to obtain a visual understanding, the curvature axis in differential geometry is applied here. Fig. 3 is normal section of C\u00f01\u00dep at the point P. For the directrix C\u00f01\u00dep , let geodesic curvature radius q\u00f01\u00deg \u00bc 1=k\u00f01\u00deg and normal curvature radius q\u00f01\u00den \u00bc 1=k\u00f01\u00den , thereinto, k\u00f01\u00den ; k\u00f01\u00deg are normal curvature and geodesic curvature of C\u00f01\u00dep in the frame of Eq. (14), respectively. Intercepting q\u00f01\u00deg and q\u00f01\u00den at e \u00f01\u00de 2 -axis and e \u00f01\u00de 3 -axis, respectively, geodesic curvature center Og and normal curvature center On can be acquired, and OgOn is the curvature axis of C\u00f01\u00dep , i.e. the instantaneous rotational axis when the normal plane moves along the directrix" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001899_s1063785008110187-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001899_s1063785008110187-Figure2-1.png", "caption": "Fig. 2. Time variation of the contact force P during the dynamic contact of a projectile with (solid curves) steel and (dashed curves) aluminum shells of various thicknesses (indicated in millimeters by figures at the curves).", "texts": [ " ---- U k 1\u2013( ) \u03d5cot X k 1\u2013( )-+ k 0= 4 \u2211= + \u03c3 U k( )G 1\u2013\u2013 \u03b4Uk 1\u2013 \u03b4t -------------- G 1\u2013 X k 1\u2013( )+ +\u239d \u23a0 \u239b \u239e , \u03b1 \u03b10 \u03b11t \u03b12t2 \u03b13t3 \u03b14t4 \u03b15t5,+ + + + += TECHNICAL PHYSICS LETTERS Vol. 34 No. 11 2008 TRANSVERSE IMPACT OF A BALL ON A SPHERE WITH ALLOWANCE 963 the terms containing time t with same powers, we obtain systems of equations with respect to unknown quantities cj (j = 0\u20133) and \u03b1i (i = 0\u20134). Once these values are determined, it is possible to write expressions for the contact force P(t) and inflection w(t) in the form of segments of power series with the known coefficients: (20) (21) The results are illustrated by the time variation of the contact force (Fig. 2) and dynamic inflection (Fig. 3) calculated for steel (\u03c1 = 7850 kg/m3, \u03c3 = 0.3) and aluminum (\u03c1 = 2700 kg/m3, \u03c3 = 0.27) shells with various thicknesses (indicated in millimeters by figures at the curves). The other parameters of impact were ass follows: m = 1 kg; V = 10 m/s; E1 = 252.53 kN/m; r0 = 100 mm; \u03d50 = \u03c0/2; R1 = 1 m. As can be seen from Figs. 2 and 3, the character of variation of the contact force and dynamic inflection in the first half of the process is different, which is explained by the dominating contributions of the linear and cubic terms, respectively, in the initial stage of impact" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000454_b138835-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000454_b138835-Figure9-1.png", "caption": "Fig. 9. The Gibbs free energy as a function of the phase fraction for various values of the stress and the phase fraction of their maximum as a function of the stress. A=\u02db 2 d = 0:25", "texts": [ " The Energy penalty, energy barrier and hysteresis in martensitic transformations 97 phase equilibrium conditions (13\u201315) are independent of the boundary conditions, so they remain unchanged. However, the proper potential for a body under loadcontrolled experiments is the Gibbs free energy G = E d with the stress of the bar. For the two-parabola bulk energy density (16), the equilibrium conditions are still (13), (17), (18). Under the assumption that the mechanical equilibrium (15) and the equilibrium for microstructures (13) are attained, we see from (32) and (15) that G = G(z; ) = E(d; z) d = 2\u02db + d(1 2z) + A(z(1 z))2=3: (38) Figure 9a shows the Gibbs free energy above as a function of the phase fraction for various values of the stress. It is obvious thatG is a concave function of z and has a maximum at ze which satisfies the phase equilibrium condition (18) by e(ze) = as shown in Fig. 9b. Thus, there is an energy barrier between the two pure phases, z = 0 and 1. Substituting (18) into (38), we find that the differences in G between the phase equilibrium state and the pure phases are G0( ) = G(ze( ); ) G(z = 0; ) = A(1+ze) 3 z2=3e (1 ze)1=3 > 0; G1( ) = G(ze( ); ) G(z = 1; ) = A(2 ze) 3 (1 ze)2=3 z1=3e > 0: (39) From (18), we know that 0 < ze < 1 for any finite values of the stress. Thus, the above differences are always positive and the phase equilibrium states always have a larger Gibbs free energy than the two pure phases" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003675_gt2011-45228-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003675_gt2011-45228-Figure3-1.png", "caption": "Figure 3. Schematics of tested bearing configurations.", "texts": [ " Potential problem areas lie in the usage of phase angle sign conventions, bearing oil film temperature profiles in shaft reference coordinates, assumed lubricant temperature profiles at the oil inlets, etc. For the calculations presented here, the complete calculation process which blends rotordynamic modeling, iterative bearing thermohydrodynamic simulations, and simple structural mechanics was performed by a Microsoft Excel Visual Basic routine. The first prototype machine was designed, built, and tested with partial arc bearings (Fig. 3). One end of the machine also contained an integral thrust bearing (TE). On the non-thrust end (NTE) there is a significant overhung mass. The pair of original bearings were essentially identical. They were fixed-arc bearings with a single oil-inlet port just above the horizontal split-line. The machine was designed as a variable speed unit with a maximum continuous operating speed of 4200 rpm. Multiple units were built and tested. The discovery of the divergent spiral vibration problem occurred during factory testing of the second unit built", "org/pdfaccess.ashx?url=/data/conferences/gt2011/70379/ on 04/25/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use modeling and simulation process were held constant while the details of the bearing geometry were changed. A host of different bearing types and geometries were considered (e.g. tilt-pad, fixed-pad, 3 lobe, 4 lobe, 5 lobe, preload, zero preload, etc.). Using the analysis method, a four lobe fixed pad bearing design (with zero preload) was identified as a potential solution (Fig. 3). Morton Effect stability plots in the complex plane and over a speed range of 3500 to 4500 rpm are presented in Fig. 18 for the original partial arc (PA) bearing design and the 4 lobe bearing (4L). The original PA bearing is predicted to exhibit unstable synchronous vibration behavior in the range of 3850 to 4200 rpm and this prediction agrees well with the test data. The figure shows that the instability with the PA bearing is predicted to desist at speeds above 4200 rpm. No testing was ever attempted above 4200 rpm, and the existence of an upper speed limit of the instability was not demonstrated", " The calculated maximum temperatures in the load zone are similar and this is consistent with test data where bearing RTD\u2019s are located near the high temperature load zone. However, the four lobe bearing has not only a smaller calculated shaft delta T value but also significantly lower average temperatures for any point on the journal believed due to the introduction of fresh lubricant at multiple locations. Compete specifications of the bearings cannot be published due to proprietary concerns. The PA bearing has a 320 degree arc with its supply groove near horizontal (Fig. 3). The 4L bearing has a load on pad configuration with zero preload. For both bearings the nominal clearance ratio is 0.25%. 10 Copyright \u00a9 2011 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/gt2011/70379/ on 04/25/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use An undamped synchronous critical speed map is shown in Fig. 20 with the various bearing stiffness values plotted along with the frequencies of the first 3 rotor modes" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001456_s11460-007-0068-x-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001456_s11460-007-0068-x-Figure3-1.png", "caption": "Fig. 3 The equivalent circuit of the separate excited DC motor", "texts": [ " The maximum output power and the optimal torque of the wind turbine gained from Eqs. (1) and (2) are given by P k T k w w w w max = = v v 3 2 opt \u23a7 \u23a8 \u23aa \u23a9\u23aa (3) where k S D Cw w w=0 25 3. ( / ) maxr lopt p . From the equation above, it is obvious that the maximum output power is related to the cube of the turbine\u2019s angular velocity, and the optimal torque of the wind turbine is related to the square of the turbine\u2019s angular velocity. The equivalent circuit of a separately excited DC motor is shown in Fig. 3, where Ra, Rf, La and Lf represent the resistance of armature winding, resistance of exciting coil, the inductance of armature winding, and the inductance of exciting coil, respectively. udcm and idcm are the terminal voltage of armature winding and the armature current, respectively. uf and if are the terminal voltage of exciting coil and the exciting current, respectively. ea is electromotive force of armature winding. The steady-state mathematic model of the DC motor is given by Refs. [13\u201315] u i R e e C T C i a a a m dcm dcm e de t dcm = + = wv w= \u23a7 \u23a8 \u23aa \u23a9 \u23aa (4) Where Ce and Ct are the electromotive force constant and torque constant, respectively, and Ce = Ct" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001427_3.60495-Figure20-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001427_3.60495-Figure20-1.png", "caption": "Fig. 20 The optimal transfers for K/ / region RFoo reaches the straight line A = 0.", "texts": [], "surrounding_texts": [ "Fig. 5 The Foo transfer. Transfers of Type Foo If the impulse at infinity is suppressed, an optimal F-type transfer will be obtained, but a particular one, in which the primer vector of Lawden7 has final length equal to one. This leads to an additional relation between E and A: E=/(A) =arc tan 4 + t2 (4 + 4t2 (9) with / = tan A (A increases from 0\u00b0-28.99\u00b0 when E increases from 0\u00b0-45\u00b0 ) . The velocity V20 just before the second impulse is given by V20= Vj/15[6 si -3 sin2A + P sin^A) I/2] (10) The eccentricity e of the orbit lying between the two impulses is given by (e2-!) 1/2 = (sin 2\u00a3-sin 2A)/(2 cos2E-cos2A) (11) Thus, limits on e are \\2) with the forward horizontal (Fig. 5). V20 is vertical, and V2 is always smaller than V20. sin A cos5A cos <& ? = sin 0o = sin (E-A) cos2A 2 cos E sin(E-A) [cos E \u2014sin E sin A cos A cos E (2 cos2E-cos2A)] (12) (13) From Eq. (12) and the limits on E and A, 2 is limited to 0\u00b0 <2< 33.74\u00b0 and ir-4L=A0>A. For an optimal F-type transfer, the final length of the primer vector must be less than or equal to one, hence we must have E>/(A) instead of E=/(A), and thus the equality in Eq. (10) is replaced by a \"greater than or equal\" sign. RF Transfer This transfer is similar to the F transfer except the vehicle grazes the forbidden sphere. This has the effect of dividing the trajectory into the two parts, before and after graze. The primer vector is analyzed separately on the two sections. A Fig. 6 The RF transfer. Fig. 7 The PNP transfer. general result is that the graze must occur before the impulse (for the convention that V2 < V } ) . An analysis similar to that for the F transfer applies, and the illustration for the optimal configuration of the single impulse is given in Fig. 6 in which the points with the subscript 0 correspond to the unconstrained optimal F transfer. The direction of the impulse to change from the grazing hyperbolic branch to the final hyperbolic branch of the transfer must be equal to A for the unconstrained F transfer as shown in Fig. 6. The optimality of the transfer requires the point R l to be at a greater distance from M than the point R w. Transfers of Type RF oo This transfer is a combination of the two previous cases: E and A must satisfy Eq. (9), V] and V2 must verify the construction of Fig. 6 and the second impulse (at infinity) has a direction ( 2 ) given by Fig. 5 andEqs. (12) and (13). PNP Transfer The PNP transfer, or biparabolic transfer, is the adaptation to the case of a planet with finite radius of the six infinitesimal impulse transfer8 of the unconstrained transfer problem. The PNP transfer is shown in the accompanying figure (Fig. 7). The vehicle is first put on a grazing trajectory by the customary infinitesimal impulse at infinity. At periapse, at the first graze, a tangential braking impulse of magnitude (U l \u2014L) leads to the parabolic trajectory shown. At infinity, an infinitesimal impulse i2 leads to a circular orbit which is traversed until the proper orientation is reached. Then a second infinitesimal impulse i3 places the vehicle on the grazing parabolic trajectory which has a common periapse with the grazing hyperbola of the desired asymptotic velocity. At this common periapse, a tangential impulse I4 of magnitude (U2\u2014L) accelerates the vehicle and leaves it on the desired departure trajectory. D ow nl oa de d by U N IV E R SI T Y O F O K L A H O M A o n Fe br ua ry 4 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /3 .6 04 95 AUGUST 1975 OPTIMAL TRANSFERS BETWEEN HYPERBOLIC ASYMPTOTES 983 Fig. 8 The Roo transfer. The characteristic velocity of the PNP transfer is (U 1 + U2 \u2014 2L), and is independent of turn angle A. It can also be demonstrated that where a] + a2Vj. These transfers are of six different types which are always grazing: 1) Type PNP: identical to that of the less than optimal deviation case and always optimal for A = 180\u00b0; 2) Type Roo: transfer with one impulse at infinity on the side of the greater velocity; 3) Type ooRoo: transfer with two impulses at infinity, one on either side of the grazing passage; 4) Type RF: transfer with one impulse at a finite distance after the grazing passage; 5) Type RFoo: transfer with two impulses after the grazing passage, one at a finite distance and the other at infinity; and 6) Type ooRF: transfer with two impulses, one at infinity before the grazing passage and the other at a finite distance after the passage. If V2 < V} the types Roo, RF, RFoo, and ooRF become ooR, FR, ooRF and RFoo. Transfer of Type Roo This transfer begins as a grazing hyperbola H and ends at infinity with the impulse7; = V2\u2014 V20\\ Fig. 8. The impulse is easy to calculate. The direction of 11 is located by the angle 4>2 with the forward horizontal and the optimality requires arctan <2 7 ) and 2, with the local forward horizontal (measured clockwise if Hl is the counterclockwise direction and conversely). The optimality requires [2V 10L2/UW(L2+ 2V10 2) (16) Transfers of Type RF Let us use Fig. 10, similar to Figs. 4 and 6. As in the \"Less Than Optimal Deviation Angle Case\" the theory of optimization of Pontryagin4 leads to a very simple result: 1) the grazing passage must occur before the impulse; 2) the velocity Vj =IR j of arrival at the impulse, corresponds to a point R ] nearer to M than R 2 with the velocity of departure given by V2 = IR2. In this case it is possible to have \u00a3= (180-.40/4) Fig. 9 The ooRoo transfer. Fig. 10 The RF transfer. Fig. 11 The RFoo transfer. Fig. 12 The ooRF transfer. ,2 - V 2 ,2 J I O \" V I O + greater than 45\u00b0 but (2 cos 2L \u2014cos2 A) remains positive and the relation \u00a3>/( A) is still valid, hence 0\u00b02 with the local forward horizontal (Fig. 11), 2<35.264\u00b0 (18) D ow nl oa de d by U N IV E R SI T Y O F O K L A H O M A o n Fe br ua ry 4 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /3 .6 04 95 AUGUST 1975 OPTIMAL TRANSFERS BETWEEN HYPERBOLIC ASYMPTOTES 985 I.I55< Fig. 18 The optimal transfers for F7/L = 1.45, A / ) the angle which locates the direction of the first impulse, measured as usual from the local forward horizontal and positive in the clockwise direction if the transfer is in the counterclockwise direction and conversely, (Fig. 12), and let us call X and Y the radial and circumferential components of the velocity of arrival V'} at the second impulse I2 (X and Y are both positive). By the theory of Lawden7 the transfer must satisfy the following two equations: K/0cos ,= V2 cos5A sin A 2 cos \u00a3 sin (E-A) (19) Y cos A sin A I (1- Y2 2L2Ycos/ U10 2(YU10+VW 2~XV]0) (20) Thus any RF transfers which satisfy these equations satisfy the optimality requirements of the ooRF transfer. The optimality also requires that 0\u00b0 < < / > / < arcsin V3/3 = 35.264\u00b0. D ow nl oa de d by U N IV E R SI T Y O F O K L A H O M A o n Fe br ua ry 4 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /3 .6 04 95" ] }, { "image_filename": "designv11_20_0003132_jmes_jour_1969_011_008_02-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003132_jmes_jour_1969_011_008_02-Figure4-1.png", "caption": "Fig. 4", "texts": [ " (25)-\u2019 Within the limitation of small (I equations (25) are identical with directly derived equations of motion for this case (I) (2) (3). The solution for parametric instabilities by the method of Hsu has been discussed elsewhere (7). Simply-supported shafi asymvnetiic over part lengih A shaft with uniform mean cross-sectional rigidity and mass distribution EI and pA is considered, simply supported at x = 0 and x = L. The section from x = 0 to x = 7jL only is uniformly asymmetric with factor u (Fig. 4). This is a case of multiple natural frequencies of the generating system, equations (8), (9) and (11) being L\u2019ol I 1 No 1 1969 at DEAKIN UNIV LIBRARY on August 12, 2015jms.sagepub.comDownloaded from PARAMETRICALLY EXCITED LATERAL VIBRATIONS OF A N ASYMMETRIC SLENDER SHAFT A -- S,+h A ' Sy+h 63 Including the first three modes of the generating system, equations (29) yield 1 0 . 1 9 6 ~ ~ ' 1 . 1 0 ~ ~ ' 1 . 8 6 ~ ~ ' u ' I u 1 - [v. .] = - 0 . 0 6 9 ~ 2 ~ 0.402~2' 0.743~2' (30) T 0 . 0 2 3 ~ ~ ~ 0 " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002515_00423110701810596-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002515_00423110701810596-Figure1-1.png", "caption": "Figure 1. Meridian section.", "texts": [ " For the tyre description in interaction absence, the following simplifying assumptions are made: \u2022 the rim is simulated by a rigid and cylindrical structure; and \u2022 the tyre is assimilated to a structure constituted by two different parts: \u2013 the sidewalls, characterised by a behaviour of an inextensible membrane structure; and \u2013 the tread, characterised by a behaviour of a structure in \u2018belt\u2019 condition, i.e. in an inex- tensible membrane condition in the parallel planes and in a rigid body condition in the meridian ones. The tyre, modelled in the described way, constitutes an inextensible \u2018envelope\u2019 with the thin walls characterised by a constant thickness, whose geometry is defined only if the same structure results connected with the rim and pressurised. In Figure 1, the tyre meridian section is represented. In particular, the sidewalls can be assimilated to circumference arcs and the tread width is equal to the rim one. D ow nl oa de d by [ Fl or id a St at e U ni ve rs ity ] at 0 0: 16 0 8 O ct ob er 2 01 4 Vehicle System Dynamics 17 Apart from the equatorial circumference radius re, the parameters characterising the tyre meridian section geometry are: c = sidewall length; w = tread width; \u03c1 = sidewall radius; = sidewall half-angle; v = distance between the sidewall centres; t = radial distance of the tread\u2013rim" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001378_ijmmm.2007.015474-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001378_ijmmm.2007.015474-Figure7-1.png", "caption": "Figure 7 Schematic representation of the cutting forces in turning", "texts": [ " No cutting fluid was used in the machining tests. The physical properties of the AISI 1045 and AISI 316L steels and the thermal conductivities of the uncoated and coated tool inserts are temperature (T) dependent as shown in Tables 2 and 3, respectively (Davis, 1998; Jawahir and Van Luttervelt, 1993; Lacombe et al., 1990). During the turning tests the cutting forces and temperatures were measured with a designed experimental set-up. The cutting forces were measured in the directions as shown in Figure 7 using a Kistler type 9255 B three-component piezoelectric dynamometer. In order to measure the temperature distribution in the deformation zone in three-dimensional cutting, an IR thermal imaging equipment was developed and applied. A detailed description of this equipment and its calibration are shown in Outeiro et al. (2004). This equipment was installed on the lathe to allow the measurement of the temperature fields in the axial and circumferential directions as shown in Figure 8. Figure 9 shows the influence of RTS on the thermal energy distribution in the deformation zone during machining AISI 1045 steel using K68 tool", " As shown in Figure 13, a constant difference exists between calculated and measured temperatures over the range of the cutting speeds used in this study, reinforcing the hypothesis that this difference is maybe due to the ploughing phenomenon. Figures 14 and 15 show both calculated and measured three orthogonal components of the Resultant force, R, resulting from machining AISI 316L and AISI 1045 steels using H13A and KC850 cutting tools. The calculated three orthogonal components of the resultant force were determined by projecting the Fc and Ft components, calculated using Equations (21) and (23), in the directions where acting of the forces measured experimentally (see Figure 7). Figure 14 shows the three orthogonal components of the resultant force versus the cutting speed, feed and depth of cut, in turning of the 316L steel using the H13A tools. The difference between measured and predicted forces is less than 10%. However, taking into account the error associated to each experimental determination of the forces, which strongly depends on the dynamometer\u2019s calibration, it can be considered that a good agreement exists between the calculated and measured forces. In the previous case, the influence of the cutting tool edge radius in the cutting process is negligible since the value of the cutting tool edge radius (0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001877_j.ijsolstr.2008.04.006-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001877_j.ijsolstr.2008.04.006-Figure1-1.png", "caption": "Fig. 1. Geometry of the problem solved; (a) elastic punches of internal angles: (i) 90 , (ii) 120 , (iii) 60 , of contact length 2a, subject to a normal load P and a shear load Q, and (b) semi-infinite wedge geometry. (c) Loading history.", "texts": [ " Further, one property of the geometry which does have a first order effect on the solution is the angle which the end faces of the contact make, as this materially affects the properties of the asymptotic solutions. To avoid the problems of having two local length scales which would occur if the contacts were chamfered, we will, here, concentrate on the cases where the contact ends are plane, and hence the contacting body is trapezium in form.1 The trapezium shaped bodies under consideration here, shown in Fig. 1(a), were analysed using the commercial finite element programme ABAQUS. First, the monotonically increasing shear problem is addressed followed by the cyclic shear problem. 1 If the contact is chamfered the local contact edge chamfer angle can be expected to control the state of stress over a length less than the length of the chamfer, whilst the angle of the end faces beyond will have a correspondingly longer range influence. The basic asymptotic forms needed for the analysis will be described: first, assume that the whole of the contact interface is adhered. It follows that, adjacent to the contact edges, the state of stress may be represented by William\u2019s solution for a wedge of total internal angle p\u00feu radians, where u is the contact internal angle (Fig. 1(b)). The direct (p\u00f0x\u00de) and shear (q\u00f0x\u00de) tractions in this neighbourhood may be written down as p\u00f0x\u00de \u00bc rhh r; p u 2 \u00bc Ko I xkI 1 \u00fe Ko IIx kII 1; \u00f01a\u00de q\u00f0x\u00de \u00bc rrh r; p u 2 \u00bc Ko I xkI 1gI rh \u00fe Ko IIx kII 1gII rh; \u00f01b\u00de where kI; kII are solutions of the following characteristic equations: kI sin\u00bdp\u00feu \u00fe sin\u00bdkI\u00f0p\u00feu\u00de \u00bc 0; \u00f02a\u00de kII sin\u00bdp\u00feu sin\u00bdkII\u00f0p\u00feu\u00de \u00bc 0; \u00f02b\u00de and the functions gI rh; g II rh; which effectively define the eigenvectors, are given in closed form by gI rh \u00bc sin\u00bd\u00f0kI \u00fe 1\u00de\u00f0p u\u00de=2 Cs I sin\u00bd\u00f0kI 1\u00de\u00f0p u\u00de=2 cos\u00bd\u00f0kI \u00fe 1\u00de\u00f0p u\u00de=2 Cc I cos\u00bd\u00f0kI 1\u00de\u00f0p u\u00de=2 ; \u00f03a\u00de gII rh \u00bc cos\u00bd\u00f0kII \u00fe 1\u00de\u00f0p u\u00de=2 \u00fe Cc II cos\u00bd\u00f0kII 1\u00de\u00f0p u\u00de=2 sin\u00bd\u00f0kII \u00fe 1\u00de\u00f0p u\u00de=2 Cs II sin\u00bd\u00f0kII 1\u00de\u00f0p u\u00de=2 ; \u00f03b\u00de where Cs i \u00bc sin\u00bd\u00f0ki \u00fe 1\u00de\u00f0p\u00feu\u00de=2 sin\u00bd\u00f0ki 1\u00de\u00f0p\u00feu\u00de=2 ; Cc i \u00bc cos\u00bd\u00f0ki \u00fe 1\u00de\u00f0p\u00feu\u00de=2 cos\u00bd\u00f0ki 1\u00de\u00f0p\u00feu\u00de=2 : \u00f04\u00de Ko I and Ko II are the generalised stress intensity factors given by Ko I \u00bc K I cos\u00bd\u00f0kI \u00fe 1\u00de\u00f0p u\u00de=2 cos\u00bd\u00f0kI 1\u00dea cos\u00bd\u00f0kI \u00fe 1\u00dea cos\u00bd\u00f0kI 1\u00de\u00f0p u\u00de=2 cos\u00bd\u00f0kI 1\u00dea cos\u00bd\u00f0kI \u00fe 1\u00dea ; \u00f05a\u00de Ko II \u00bc K II sin\u00bd\u00f0kII \u00fe 1\u00de\u00f0p u\u00de=2 sin\u00bd\u00f0kII 1\u00dea sin\u00bd\u00f0kII \u00fe 1\u00dea sin\u00bd\u00f0kII 1\u00de\u00f0p u\u00de=2 sin\u00bd\u00f0kII 1\u00dea \u00fe \u00f0kII 1\u00de \u00f0kII\u00fe1\u00de sin\u00bd\u00f0kII \u00fe 1\u00dea ; \u00f05b\u00de where K I and K II are the more usual mode I and mode II stress intensity factors found by collocating the solution to whatever finite problem is being studied along the \u2018notch\u2019 bisector, and are defined as K I \u00bc rhh\u00f0r;0\u00de rkI 1 Lt r ", " (7c) and xo denotes the position from the contact edge where the argument of the logarithm passes through unity. The value of k for the problem under consideration here (where the wedge and the half-plane have the same elastic constants) is given by a root of the following characteristic equation: r\u00f0u; f ; k\u00de \u00bc cos kp\u00f0sin2 ku k2 sin2 u\u00de \u00fe 1 2 sin kp\u00f0sin 2ku\u00fe k sin 2u\u00de \u00fe f sin kp\u00bdk\u00f01\u00fe k\u00de sin2 u \u00bc 0: \u00f08\u00de Also note that the coefficient of friction is given a sign in the slipping asymptotic solution: a positive value of f suggests the kind of slip which is experienced at a trailing edge (\u2018inward\u2019 slip), Fig. 1(b), whilst a negative value of f implies the kind of slip seen at the leading edge of a contact (\u2018outward\u2019 slip). We aim to investigate in detail the response of trapezium-shaped punches with internal angles of 60 , 90 and 120 (as shown in Fig. 1(a)), but also to look at the general characteristics of the solution for all values of punch internal angle in a realistic range. Under a range of conditions the interface will be mainly or entirely adhered, so it is appropriate to consider, first, conditions for complete adhesion. The two basic requirements for adhesion at the contact edges are Ko I < 0; \u00f09\u00de f > gI rh; \u00f010\u00de and the first of these will certainly be fulfilled under normal loading alone, so that the second is sufficient, and follows directly from the characteristic equation, i", " The procedure to restore the separation condition within the asymptote is described in detail by Karuppanan et al. (2008). Also highlighted on Fig. 2(a) are the three angles studied in detail, showing that; (a) the 120 and 90 punches exhibit region A and B behaviour, and (b) the 60 punch exhibits region D, C, B and A behaviour with increasing friction. Fig. 2(b) provides a set of schematics to illustrate the response anticipated in all regions of Fig. 2(a). As an example problem, a finite element analysis of the 60 punch problem (since it covers all four regions), shown in Fig. 1(a), iii, was carried out using ABAQUS, and the results for the normalised pressure distribution found are shown in Fig. 3(a). These results are re-plotted enlarged at the trailing edge in Fig. 3(b) for an improvement in clarity, and it displays the kind of behaviour which is as anticipated. The procedure adopted by Churchman and Hills (2006), for the problem shown in Fig. 1(a), i, will be briefly summarised in this section. In his attempt to understand the general interfacial characteristics of complete contacts in partial slip (for a monotonically increasing shearing force) a map shown in Fig. 4, was developed. For a square punch problem, i.e. when u \u00bc p=2; the eigensolution in the adhered regime is characterised by 2 The kI \u00bc 0:5445; kII \u00bc 0:9085; gI rh \u00bc 0:5431; gII rh \u00bc 0:2189: \u00f012\u00de As 0 > kI 1 < kII 1, the state of stress is singular and, furthermore, the traction ratio is given by the dominant eigensolution, so that q\u00f0x\u00de p\u00f0x\u00de \u00bc gI rh Lt x", " In region BB \u00f01=p < f < 0:392\u00de; similar behaviour to that found in region BA is expected, except that the tractions are bounded at the trailing edge and singular at the leading edge when line EF is reached. Finally in region BC \u00f00:392 < f < gI rh\u00de; the forward slip zone at the trailing edge will grow rapidly if Q=P ratio lies above the grey line, as does the separation distance. When the gross sliding condition is achieved, line FG, the tractions exhibit square root bounded behaviour at the trailing edge and singular behaviour at the leading edge. The geometry of the problem studied here is given in Fig. 1(a), ii. A practical problem having a contact edge condition very close to the one studied here is the involute spline used in split shafts within some gas turbines, and the map giving the overall response to applied loads is shown in Fig. 5(b). It is assumed that the normal load is applied first, held constant, and then a monotonically increasing shear load imposed. If f > gI rh the application of normal load will ensure full adhesion of the contact interface, line AB. This condition will hold with a finite shear load, region AA1, until line EF is reached", " Subsequent application of a shear load will increase the size of slip zone at the leading edge, whilst the trailing edge experiences instantaneous stick followed by a minute forward slip and separation region attached to the edge, due to the influence of locked-in residual tractions, region BA. As we move in region BA closer to gross sliding condition (line OC), the influence of the sliding eigenvalue becomes increasingly important and the minute amounts of separation which occurs earlier will gradually diminish. Finally a gross sliding condition is achieved when line OC is reached where the tractions exhibit singular behaviour at both edges. The geometry of the problem studied here is given in Fig. 1(a), iii. This kind of geometry also occurs, in a modified way, in a gas turbine engine. It is the overall shape of the dovetail lug fastening in a fanblade. Of course the fanblade normally has rounded edges, but the solution to be discussed shows what might be expected if that radius were reduced to a very small value. For a 60 punch problem, i.e. when u \u00bc p=3 kI \u00bc 0:6157; kII \u00bc 1:1489; gI rh \u00bc 0:9619; gII rh \u00bc 0:3224: \u00f019\u00de The new overall response map is as in Fig. 5(c) showing a more complicated response than the first two geometries", " The shakedown limit using this technique is achieved provided that j f j> ~qQ Q P max\u00bdj 1 2~pP \u00fe ~pQ Q P max\u00bdj\u00fe 1 max 8x \u00f021\u00de where ~pP \u00bc pP\u00f0x\u00de \u00f0P=2a\u00de ; ~pQ \u00bc pQ \u00f0x\u00de \u00f0Q=2a\u00de ; ~qQ \u00bc qQ \u00f0x\u00de \u00f0Q=2a\u00de ; \u00f022\u00de and pP\u00f0x\u00de is the traction along the interface under normal load only (P) whilst pQ \u00f0x\u00de; qQ \u00f0x\u00de are the tractions resulting from pure shear load only (Q). These traction values are determined from adhered finite element analysis. It is assumed that the contact is first loaded by shear up to some value, Qmax. It is then reduced to a value Q min \u00bc jQ max ( 1 < j < 1), and is subsequently cycled between these limits, Fig. 1(c). In very general terms, the closer j lies to +1 (so that the range of shear is small), the more likely the contact is to shakedown. The steady state response map for the square punch problem is shown in Fig. 6(a). The boundary between shakedown (below, to the right) and cyclic slip (above, to the left) for a fully reversing shear case (j \u00bc 1) is represented by a thick dark line. The full description of the steady state contact response will be based on this example case. If f > gI rh; in region AC1, the contact interface will be fully adhered from outset" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000809_1.328994-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000809_1.328994-Figure3-1.png", "caption": "Fig. 3. Demagnetization curves and energy products of the 7% and 10.5%Co Cr-Co-Fe isotropic alloys along with those of Vicalloy and Cunife.", "texts": [ " These values generally become lower with decreasing Co content. A 7Co-28Cr Fe alloy, for example, develops values of Br ~ 9700G, Hc - 330 Oe and (BH)max _ 1.4 MGOe, see Table 1. The heat treatment times here are longer than those for the 10.5%Co alloy, as one might expect from the de creased decomposition kinetics with decreasing Co con tent. The isotropic low Co alloys compare very favor ably with Vicalloy in all respects and with Cunife in terms of Br and (BH)max at th~ sa~rifice of coercivity. Fig. 3 shows tYP1cal demagnet1zat10n curves for these materials. ANISOTROPIC MAGNETS PREPARED BY \"DEFORMATION-AGING\" Deformation-aging is a technique developed by Jin (9) to impart magnetic anisotropy by mechanically elongating the precipitate structure rather than re sorting to the use of an aligning magnetic field. It takes advantage of the ductile quality of the Cr-Co-Fe alloys. The principle of deformation-aging is shown sche matically in Fig. 4. The alloy is first cooled through the spinodal temperature at such a rate as to develop oversized, and generally near-spherical, particles" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.54-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.54-1.png", "caption": "Fig. 14.54 Single-disk floating machine for concrete", "texts": [ " Floating is performed after the base has gained some mechanical strength. Rough and finishing floating are distinguished. In the case of older types of floating machines, rough floating was effected using a slowly rotating solid disk, while finishing floating was effected using blades set at an appropriate angle to the base. Now this practice has changed and the entire process is carried out using blades and changing the angle of their inclination, though solid disks are also used. Single-disk floating machines (Fig. 14.54) are used for smaller concrete works. The working tools are blades or solid disks that are 600\u20131200 mm in diam- eter. They are driven by combustion engines or electric motors with a power of up to 8 kW. Combustion engine floating machines feature stepless control of the floating tools\u2019 rotational speed. Electric floating machines are equipped with two-speed electric motors whereby one can select the appropriate speed of rotation for the rough and finishing floating tools. Modern single-disk floating machines have the following features: \u2022 A long handle, enabling access to floated surfaces with no need to walk on them (during transport of the floating machine the shaft is folded)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002995_j.mechatronics.2010.05.011-Figure16-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002995_j.mechatronics.2010.05.011-Figure16-1.png", "caption": "Fig. 16. When a CAMBADA robot is on, the estimated centers of the detected obstacles are compared with the known position of the team mates and tested; the left obstacle is within the CAMBADA acceptance radius, the right one is not.", "texts": [ " Also, it would be pointless to pay attention to obstacles that are outside the field of play, since the surrounding environment is completely ignorable for the game development. To be able to distinguish obstacles, identifying which of them are team mates and which are opponent robots, a fusion between the own visual information of the obstacles and the shared team mates positions is made. By creating a circle around the team mate positions with the robot radius (considered 22 cm), a matching of the estimated center of visible obstacle area is made (Fig. 16), and the obstacle is identified as the corresponding team mate in case of a positive matching (Figs. 17c and 18c). This matching consists on the existence of interception points between the team mate circle and the obstacle circle or if the obstacle center is inside the team mate circle (the obstacle circle can be smaller, and thus no interception points would exist). Since the detected obstacles can be large blobs, the above described identification algorithm cannot be applied directly to the visually detected obstacles", " for c:=1 to total_number_of_team_mates for o:=1 to total_obstacles_of_team_mate for m:=1 to total_own_obstacles if m matches o I already know this obstacle, do nothing else if previously known by another team mate obstacle confirmed and added else obstacle considered temporarily waits for confirmation by another team mate endif endif endfor endfor endfor The matching of the team mate obstacles with the own obstacles is done in a way similar to the matching of the obstacle identification with the team mate position described earlier. The CAMBADA team mate position in Fig. 16 is replaced by the current team mate obstacle for the matching test. Fig. 21 shows a situation where robot 2, in the goal area was too far to see the obstacle on the middle of the field. Thus, it considered the obstacle in question, only because it is identified by both robots 5 and 6, as visible in the figure. The techniques chosen for information and sensor fusion proved to be effective in accomplishing their objectives. The Kalman filter allows to filter the noise on the ball position and provides an important prediction feature which allows fast detection of deviations of the ball path" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure1-1.png", "caption": "Fig. 1. Involute circular-arc gear drive.", "texts": [ " When the directrix is a spherical curve, this type of gearing is only applied to intersected-axis drive, and the directrixes are conjugate curve on the same sphere whose center is axial intersection point and two datum surfaces of gear pair coincide with each other and are a sphere where a pair of conjugate directrixes is. Normal circular-arc surface based on conjugate plane curve or spherical curve is tooth flank. In engineering application, generating circle with biggish radius is generally chosen as tooth trace, and the directrix is as tooth profile. For paralleled-axis drive, when the directrix is an involute, such a gear is called as involute cylindrical gear whose tooth trace is a circular-arc (Fig. 1(a)), some researchers have explored its theories; for intersected-axis drive, when the directrix is a spherical involute, the gear becomes a special example of spiral bevel gear (Fig. 1(b)). (2) The directrix is a curve on the pitch surface. According to the theory of gearing, for paralleled-axis or intersected-axis drive with constant speed ratio, there respectively is a pair of merely rolling cylinders or cones, namely pitch cylinder or pitch cone, which is called uniformly as pitch surface here. A pair of pitch surfaces, i.e. the axode, is tangent at an instantaneous axis. But for crossed-axis drive, because there is not one point where the relative velocity v\u00f021\u00de p is equal to zero, two axodes cannot merely roll each other; we can consider that the pitch surface does not exist" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001756_1.2736705-Figure20-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001756_1.2736705-Figure20-1.png", "caption": "Fig. 20 Oscillating process", "texts": [ " However, the use of lubricant has a pronounced influence on both the friction level and the friction hysteresis loop. Without lubricant Fig. 19 , the friction level rises with increasing oscillating frequency; with lubricant Figs. 11 and 12 , the friction level drops with increasing oscillating frequency. 2.4.6.2 Function of grease. To further explore the hysteresis characteristic associated with grease, the velocity history in one oscillation cycle is divided into four stages: positive acceleration, positive deceleration, negative acceleration, and negative deceleration. This is illustrated in Fig. 20, where for clarity the shaft is shown completely separated from the bushing. The effect of grease on the friction hysteresis is represented by the variation of the lubricant flow. Referring to Fig. 20, in stage a , the shaft accelerates to the highest speed. It brings the volumetric flow forward Fig. 21 a . In stage b , the shaft decelerates to zero speed. Once the slowdown process is triggered, the surface asperities of shaft and bushing will act as barriers for the flow and tend to lower the flow rate and in some regions create backflow Fig. 21 b . Due to inertia, the flow will not be the same as that in stage a at the same shaft speed. Hence, the hysteresis feature occurs. As for the other half oscillating cycle\u2014stages c and d , a similar process is followed" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001571_s1560354707030045-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001571_s1560354707030045-Figure2-1.png", "caption": "Fig. 2. Scheme of the Snakeboard.", "texts": [], "surrounding_texts": [ "MSC2000 numbers: 70F25, 70E55, 70E60, 70E18 DOI: 10.1134/S1560354707030045\nKey words: Snakeboard, Gibbs\u2013Appell equations, dynamics, analysis of motion\n1. INTRODUCTION\nThe Snakeboard is one of the modifications of a well-known skateboard. It allows the rider to propel himself forward without having to make contact with the ground even if a motion occurs uphill. The motion of the snakeboard becomes possible due to a specific features of its construction and due to the special coordinated motions of legs and a torso of the rider. The first snakeboard has appeared in 1989 and from this moment till now it has found a lot of fans among the amateurs of extreme sports. Soon after the invention of the snakeboard the first attempts were made to give a mathematical description of the basic principles of human snakeboarding. The basic mathematical model for the snakeboard investigated by various methods in many papers [1\u201313] was proposed by Lewis et al. [1]. In our paper we give the further development of the model proposed in [1].\nThe Snakeboard (see Figs. 1\u20132) consists of two wheel-based platforms upon which the rider is to place each of his feet. These platforms are connected by a rigid crossbar with hinges at each platform to allow rotation about the vertical axis. To propel the snakeboard the rider first turns both of his feet\n*E-mail: kuleshov@mech.math.msu.su\n321", "in (see Fig. 3). By moving his torso through an angle, the snakeboard moves through an arc defined by the wheel angles. The rider then turns both feet so that they point out, and moves his torso in the opposite direction. By continuing this process the snakeboard may be propelled in the forward direction without the rider having to touch the ground.\nThe mathematical model of the snakeboard considered in this paper is represented in Fig. 4. We assume that the snakeboard moves on the xy plane, and let Oxy be a fixed coordinate system with origin at any point of this plane. Let x and y be the coordinates of the system center of mass (point G) and \u03b8 the angle between the central line of the snakeboard and the Ox-axis. In the basic model treated in [1] platforms could rotate through the same angle in opposite directions with respect to a central line of the snakeboard (by other words, for this model \u03d5f = \u2212\u03d5b = \u03d5, see Figs. 3\u20134). We suppose that platforms can rotate independently and their positions are defined by two independent variables \u03d5f and \u03d5b (Fig. 4). The motion of the rider is modeled by a rotor, represented in the form of a dumb-bell in Fig. 4. Its angle of rotation with respect to the crossbar is denoted by \u03b4.\nREGULAR AND CHAOTIC DYNAMICS Vol. 12 No. 3 2007", "Let l be the distance from the system center of mass G to the location of the wheels (points A and B). We assume that |GA| = |GB| = l, see Fig. 4.\nThe platforms of the snakeboard are assumed to move without lateral sliding. This condition is modeled by constraints which may be shown to be nonholonomic. For the front platform corresponding constraint has a form\nx\u0307 sin (\u03d5f + \u03b8) \u2212 y\u0307 cos (\u03d5f + \u03b8) \u2212 l\u03b8\u0307 cos \u03d5f = 0 (1)\nand for the rear platform it has a form\nx\u0307 sin (\u03d5b + \u03b8) \u2212 y\u0307 cos (\u03d5b + \u03b8) + l\u03b8\u0307 cos \u03d5b = 0. (2)\nWe can solve equations (1) and (2) with respect to x\u0307 and y\u0307. Then\nx\u0307 = l\u03b8\u0307\nsin (\u03d5f \u2212 \u03d5b) (cos \u03d5b cos (\u03d5f + \u03b8) + cos \u03d5f cos (\u03d5b + \u03b8)),\ny\u0307 = l\u03b8\u0307\nsin (\u03d5f \u2212 \u03d5b) (cos \u03d5b sin (\u03d5f + \u03b8) + cos \u03d5f sin (\u03d5b + \u03b8)).\n(3)\nFurther we describe a motion of platforms using new variables \u03c81 and \u03c82, connected with variables \u03d5f\nand \u03d5b by relations\n\u03c81 = \u03d5f \u2212 \u03d5b, \u03c82 = \u03d5f + \u03d5b.\nControl of the snakeboard is realized by rotations of the platforms through \u03d5f and \u03d5b and by rotation of the rotor through \u03b4. We assume that the variables \u03b4, \u03c81 and \u03c82 are known functions of time t, i.e.\n\u03b4 = \u03b4 (t), \u03c81 = \u03c81 (t), \u03c82 = \u03c82 (t).\nThese variables are the controlled variables in this problem.\n2. EQUATIONS OF MOTION\nWe derive now equations of motion of the given model of a snakeboard in the form of the Gibbs\u2013 Appell equations. For this purpose we use the method, which was applied earlier by Ispolov and Smolnikov [14] to study the skateboard dynamics. We introduce pseudovelocity V by the formula\nV = l\u03b8\u0307\nsin \u03c81 (cos \u03c81 + cos \u03c82). (4)\nFrom this formula we have\n\u03b8\u0307 = V l \u00b7 sin \u03c81 cos \u03c81 + cos \u03c82 . (5)\nUsing (4) we can rewrite expressions (3) for x\u0307 and y\u0307 as follows:\nx\u0307 = V cos \u03b8 \u2212 V sin\u03c82 sin \u03b8\ncos \u03c81 + cos \u03c82 , y\u0307 = V sin \u03b8 +\nV sin\u03c82 cos \u03b8\ncos \u03c81 + cos \u03c82 . (6)\nIn a case, when \u03d5f = \u2212\u03d5b = \u03d5 and therefore \u03c82 = 0 equations (6) have a very simple form\nx\u0307 = V cos \u03b8, y\u0307 = V sin \u03b8.\nSo, in this case pseudovelocity V can be interpreted as velocity of the system\u2019s center of mass. In general case, this pseudovelocity is velocity of a point which can be obtained if we take a projection of the system\u2019s instantaneous center onto the line passing through the central line of a snakeboard (i.e. passing through the crossbar).\nREGULAR AND CHAOTIC DYNAMICS Vol. 12 No. 3 2007" ] }, { "image_filename": "designv11_20_0002348_1.2943299-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002348_1.2943299-Figure3-1.png", "caption": "Fig. 3 A five", "texts": [ " 2 corresponds to the two branches of the linkage CC-I and CR-I or two sub-branches of a linkage branch RR-I and RR-II . 4. In each sub-branch, as the linkage configuration varies from one dead-center position to the other, the input value increases or decreases monotonously from one limit to the other. To transform the linkage from a configuration in one sub-branch to a configuration in the other sub-branch, the variation of the input must pass one of the two input limits. 3.2 Five-Bar Linkages. A five-bar linkage has two DOFs and the inputs may be given through any two joints Fig. 3 . An un- Transactions of the ASME 6 Terms of Use: http://www.asme.org/about-asme/terms-of-use c c c n a 3 i l t c k l a J J Downloaded Fr ertainty singularity occurs when the three passive joints become ollinear. The input data points corresponding to all possible unertainty singularities form the boundary of the JRS. A common practice is to have the inputs given through two eighboring joints. Let and be the input angles and the ngle between the links containing the three passive joints Fig. a ", " By eeping =0 or , a five-bar linkage is reduced to a four-bar inkage called singular linkage because the five-bar linkage is at n uncertainty singularity , which leads to one of the four types of RS boundary curves: CC-I, CR-I, RR-I, or RR-II Fig. 2 . ournal of Mechanical Design om: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/28/201 3.3 JRS Types of Five-Bar Linkages. Based on the existence and types of boundary curves, 17 possible types of JRS can be enumerated Table 1 . Samples of these 17 types of JRS are shown in Figs. 4 and 5, in which the link lengths a ,b ,c ,d ,e , as defined in Fig. 3 a , of the linkage are listed in Table 1. These 17 types of JRS are enumerated based on the principle outlined in the aforementioned rotatability laws and the following observation. 1 If is the angle at a joint connecting a short link, it can reach 0 and 27 . There will be two different singular linkages and the JRS is in the region between the heavy and light boundary curves corresponding to =0 and Figs. 4 b \u20134 d and 5 f \u20135 m . 2 For Class I linkages, we have the following. linkages: \u201ea\u2026 CC-I, \u201eb\u2026 CR-I, \u201ec\u2026 RR-I, and ar -bar chain SEPTEMBER 2008, Vol", " It is noted that the coexistence of CC-I/CC-I, CR-I/CR-I, or RR-I/RR-I may occur only in Class I linkages and the coexistence of CC-I/CR-I boundary curves is not possible in any linkage. Thus, by enumerating all possible boundary curves and their ombinations, 13 types of JRS can be found from Class II chains Table 1 . Samples of these JRS are shown in Fig. 5. The identication and classification of JRS are demonstrated in the followng example. Example 1. Let the link lengths of a five-bar chain be a ,b ,c ,d ,e = 8,4 ,10,15,12 , where the link notations are hown in Fig. 3 a . Determine the JRS type of the input angles nd . 1. Because 15+8+4 12+10 , this is a Class II chain. 2. Because 15+8 12+10+4 , Link a is a short Link and the two input joints on it are revolvable joints. The linkage is of CC type. 3. Since 15+10 12+8+4 , Link c is a long link. Because Link d is the longest link, the angle between Links c and d may reach 0 but not . The singular linkage, which is a Class II chain with the link dimensions a ,b ,d-c ,e = 8,4 ,5 ,12 , is formed by keeping Links c and d folded", " Because the five-bar linkage is a CC type, the JRS must be the region outside the RR-II boundary curve Table 1, Fig. 5 d . Due to invariant link rotatability 6 , i.e., the rotatability beween any pair of links in kinematic chains formed by the same et of N links is independent of the order of link connection, the bove JRS types and classification scheme are also valid to clasify the rotatability of any two links with respect to a common eference link, for example, the rotation range of Links c and d ith respect to Link a in Fig. 3 b . Figures 4 and 5 represent the possible JRS types or patterns. hey offer a useful visual tool to understand the rotatability of ve-bar linkages. The usefulness of such a tool will be highlighted ater with the applications in more complex linkages. The followng section explains how to read and interpret a JRS. 3.4 Sheets and Sides of JRS. A linkage is programed by ontrolling the input parameters to reach the desired positions ith the desired linkage configurations. In the JRS of a linkage, a oint generally corresponds to more than one linkage configuraion", " Linkage configurations with 0, correspond to points on one sheet, while those with ,2 correspond to points on the other sheet. The sheet identity is irrelevant to the choice of the input joints. Side (i.e., sub-branch) identification. The side identity is af- SEPTEMBER 2008, Vol. 130 / 092303-5 6 Terms of Use: http://www.asme.org/about-asme/terms-of-use f p W t Y p p e t a o s m h D g c F b L n t J i a n o c t N 0 Downloaded Fr ected by the choice of the input joints. An N-bar linkage has three assive joints, which may be denoted as X, Y, and Z Fig. 3 . hen X, Y, and Z become collinear, the linkage is at an uncerainty singularity configuration. Let be the angle measured from Z to YX. Linkage configurations with 0, correspond to oints on one side, while those with ,2 correspond to oints on the other side. =0 or corresponds to points on the dge of a JRS sheet. A sheet or branch is usually identified before he side identification. One may note that the purpose of sheets and sides for branches nd sub-branches is to provide an intuitive model to establish the ne-to-one correspondence between the linkage configuration pace and the input domain" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002862_s11740-009-0165-1-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002862_s11740-009-0165-1-Figure2-1.png", "caption": "Fig. 2 Speed conditions during the gear honing process", "texts": [ " The cutting speed in profile direction vH depends on the radius of curvature q of the gear flanks and the revolution speed n (Eq. 1). The cutting speed in flank direction vL is influenced by the helix angle b, the crossaxis angle R, the revolution speed nW and the radius rW (Eq. 2). The grit paths of a grain are formed in the direction of the cutting speed between the workpiece and honing tool. Due to the speed conditions, a height-oriented surface structure is generated. In publications [2, 4\u20136] it is said that hereby the noise behavior of the gear flank during contact is improved (Fig. 2). The improvement results from a lower noise excitation compared to the more periodic excitation with parallel surface structure orientation. During gear honing a cutting speed within the range of vc = 0.5\u201310 m/s is possible. Due to these low cutting speeds and the brief contact times, the amount of energy induced in the flank is low. This means that grinding burn is not likely to arise on the flanks of the workpiece during the honing process [2, 7]. An additional advantage of honed gears is the resulting high residual compressive stress, which leads to a longer lifetime of the workpieces [4, 5]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001027_icit.2005.1600634-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001027_icit.2005.1600634-Figure1-1.png", "caption": "Fig. 1. Elementary induction machine", "texts": [ " This approach takes advantage to define inductances of a machine by taking into account the skew of the slots, and it can be extended to the study of other types of axial asymmetries, namely, the axial eccentricities. In this work, a 2-D model of the induction machine will be approached while focusing the study on its first aim; the modelling of induction machine inductances with nonsinusoidal distribution of the stator winding, the axial and radial non-uniformity of the air gap. The study will be based on an extension of the MWFA, simulation results as well as comments will be exposed. To formulate the problem, we refer to the Fig. 1 which gathers together two cylindrical masses separated by an airgap. One of it is hollow out and represents the stator, and the other represents the rotor. An abcda arbitrary contour is defined thanks to a reference frame fixed on the stator, to an axial reference and to the mechanical position of the rotor measured by respecting a fixed stator reference. At a rotor position O, are defined the angles 9oo = 0, zo = 0 and they are located by the points a and b, and by the same way, the angle o and the length z are located by using the points c and d" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000991_s1474-6670(17)61875-5-Figure2.3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000991_s1474-6670(17)61875-5-Figure2.3-1.png", "caption": "Fig. 2.3. Successive refinements of a poly gon track by manipulations of a", "texts": [ " Rational polygon tracks for better approximation of the ideal form can be de rived from the hexagon by modifying it in the corner areas, wherby the vicinity of the midpoints of the hexagon sides remains unchanged. All rational modifications of a corner area of a hexagon consist of just re placing pieces of one of the corner's straight lines by pieces of the other one and vice versa. The perimeter of the track thereby remains unchanged. All rational polygon tracks have the same perimeter as the hexagon they are derived from perimeter = 413 'l' o (2.1) where 'l' is the half opening of the original hexagono(Fig. 2.3). The tip of the stator flux-linkage vector moves along the polygon track either with the constant linear velocity of (2.2) where U D - DC-link voltage or stands still! (pauses). Combining equa tions (2.1) and (2.2) we get the total ac tive time per revolution around the polygon 613 'l' o =--- neglecting the influence of the stator re sistance hexagons (A) corner areas. A hexa gon corner is replaced by 3 (B), 5 (C) or 7 (D) corners. ~e radial distance measured between the cen ter of the original hexagon and the polygon track has always extremum values at corners" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000576_1.2194070-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000576_1.2194070-Figure2-1.png", "caption": "Fig. 2 Simple robotic arm", "texts": [ " To eliminate chattering in the existence subspace, the width of the smoothing boundary layer has to be made greater or equal to E and, therefore defined as a function of the level of uncertainty in the model. JUNE 2006, Vol. 128 / 347 Terms of Use: http://www.asme.org/about-asme/terms-of-use angular velocity 348 / Vol. 128, JUNE 2006 Downloaded From: http://dynamicsystems.asmedigitalcollection.asme.org/ on 01/28/2016 To illustrate the approach, the extended variable structure filter is applied to a simple robotic arm as shown in Fig. 2 49 . The mathematical model of the arm can be obtained as T = J d2 dt2 + B\u0307 + MgL 2 sin 75 Where J=0.58 kg m2 is the moment of inertia of the overall system, M =6.97 kg is the mass, L=0.3 m is the length of the arm, B=3.5 Nms/rad is the coefficient of viscous friction, g =9.82 m/s2 is the acceleration due to gravity, T in Nm is the applied torque and is the deflection angle in degrees. Assuming a sampling interval =0.001 s, with x1= and x2=d /dt, the discrete state equations may be obtained as x1k+1 =x1k + x2k and x2k+1 = 1\u2212B x2k + T /J\u2212MgL/2J sin x1k " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003079_13506501jet670-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003079_13506501jet670-Figure1-1.png", "caption": "Fig. 1 Physical configuration of the rough porous partial journal bearing", "texts": [ " In this article, the Christensen\u2019s stochastic theory for rough surfaces is used to analyse the effect of surface roughness on the squeeze film characteristics of short porous partial journal bearings with micropolar fluids. Two types of one-dimensional (1D) surface roughness (longitudinal and transverse) patterns are considered. The modified stochastic Reynolds-type equation governing the mean film pressure in the presence of micropolar fluids is derived for the two types of roughness patterns. The closed-form expressions for the mean film pressure, the mean load-carrying capacity, and squeeze film time are obtained. The physical configuration of a rough porous partial journal bearing is shown in Fig. 1. A solid journal of radius R approaches the rough porous bearing of wall thickness H0 at any circumferential section \u03b8 with a velocity V . The lubricant in the film and also in the porous region is taken to be Eringen\u2019s [13] micropolar fluid. The stochastic film thickness H is represented by H = h + hs(\u03b8 , z, \u03be) = C(1 + \u03b5 cos \u03b8) + hs(\u03b8 , z, \u03be) (1) where h(= C + e cos \u03b8) denotes the nominal smooth part of the film geometry, while hs(\u03b8 , z, \u03be) is the part due to the surface asperities measured from the nominal level and is a randomly varying quantity of zero mean, \u03be is an index determining a definite roughness arrangement, and \u03b5(= e/C) is the eccentricity ratio parameter" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001713_crat.19750100903-Figure13-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001713_crat.19750100903-Figure13-1.png", "caption": "Fig. 13 . Hypothetical structurc of point singularities with s ~ 1 in smectic schlieren textures", "texts": [ " With respect to this tendency it seems more realistic to have plane layers in the B schlieren textures than crumpled layers. 4.3 . Structure around Point Singularities Now the structure around the point singularities of the smectic schlieren textures is to be considered. The considerations are based on the structures of the neighbourhood of point singularities in nematic liquids. Assumed the molecules are arranged parallel to the glass surface and the smectic layers are not plane but concentric around the point singularity. With these assumptions the structure of points with s = _+1 (Fig. 13) does not possess the optical properties of the real observed singularity points. Both types of points would show spiral type optical properties because of the fact that the molecules do not diverge radially from the centre, but change continuously the direction going from the centre to an outer part. This type of structure cannot give the really observed rectangular dark cross and the most occuring non spiral structure ot the dark brushes. Assumed the molecules are tilted (tilt angle constant) with respect to the glass surface and the layers are plane and parallel t o the surface (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001213_icar.2005.1507428-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001213_icar.2005.1507428-Figure4-1.png", "caption": "Figure 4: Free body diagram for the dynamic thrust test rig.", "texts": [ " In small backpackable aerial robots like ours, such over-engineering can lead into a vicious design cycle. To avoid this, accurate motor-propeller data is needed to narrow margins of error. Unless thrust measurements are made under real or simulated flight conditions, they may not be realistic indications of actual performance. To be meaningful, thrust measurements needs to be dynamic, with the propeller moving through airflow as in actual flight. As such, a test rig (see Figure 3) was constructed to collect dynamic thrust data. The test rig\u2019s free body diagram is given in Figure 4. The motor and propeller combination is attached to one end of the arm. The arm is then pivoted at a distance D from its center of mass. The arm angle, \u00b5, as measured by the protractor depends on the generated thrust T . Let M be the distance between the propeller and pivot point while W is the weight of the entire test bed and motor-propeller combination. At equilibrium, the torques generated by thrust and weight are balanced such that DW cos \u00b5 = MT . Consequently, the dynamic thrust is given by T = DW cos \u00b5 M (4) The net effect is that such a test rig allows one to rapidly collect dynamic thrust data and compare different motor-propeller combinations" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000865_004051758205200606-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000865_004051758205200606-Figure2-1.png", "caption": "FIGURE 2. Test fixture for biaxial loading of triaxial fabrics:", "texts": [ " In both cases all yarn sets maintain the same spatial relationship with one another and, in both cases, the interlacing of warp and filling is not due to conventional shedding, but rather to the special movement of the warps. If one warp set is removed, the interlacing ceases to exist, therefore the leno weave is felt in many ways to be the true conventional equivalent of the triaxial fabric. Testing Procedure - The first objective of the testing procedure was to attempt to use Equations 1 and 2 to predict the modulus of different triaxial fabrics during crimp removal. This was done using a device, illustrated in Figure 2, which was developed in the School of Textiles for this investigation. The fixture accepted a fabric in the form of a Maltese cross, the dimensions of which are illustrated in Figure 3. Loading was performed by two independent screw controlled compression springs, one for each pair of jaws. Each specimen was preloaded with a biaxial load of 9.0 N in each direction, and the dimensions of the center square were measured using a vernier caliper accurate to 25.4 ~.m. Alternating incremental loading of 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000935_00207720600566495-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000935_00207720600566495-Figure4-1.png", "caption": "Figure 4. Schematic diagram of the NEROV UUV.", "texts": [ " The second terms in equations (50) seem to correspond to a forward-travelling wave in the backpropagation network that provides a second-order correction to the weight tuning for F\u0302. The convergence of s to zero and the boundedness of the weights F\u0302 and G\u0302 can be shown again using Barbalat\u2019s lemma (Slotine and Li 1991). The computational study in this paper was based on the model structure of the Norwegian Experimental Remotely Operated Vehicle (NEROV) (Spangelo and Egeland 1994). A sketch of the NEROV vehicle is shown in figure 4. The vehicle is controlled in all six DOFs by six dc-motor-driven thrusters. The fluid velocity was chosen to be zero in all computations. The desired velocities and positions were generated by D ow nl oa de d by [ L ak eh ea d U ni ve rs ity ] at 0 6: 11 1 2 M ar ch 2 01 3 a reference trajectory generator. The simulation, using the developed controller, was performed at 5Hz, which implies that all the measurements and the control action occurred at a time step of 0.2 s. The tracking performance of the X and Y positions, the depth and the heading angle were considered in this study" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002504_robio.2009.4913206-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002504_robio.2009.4913206-Figure2-1.png", "caption": "Fig. 2 Different kinds of BWS (a) and Lokolift (b) [4]", "texts": [ " 1 b) and a ball joint with a range of movement limited to 2 passive DOF by the plate, so that the range of pelvic movement can be adjusted (see Fig. 1 c), taking into consideration the ease of walking. The pelvic support mechanism has been reported in detail [4]. In this paper, we focus only on the force control we have adopted for our BWS system. Many different methods of providing unloading force exist. According to previous study [5], many BWS systems which is available on the market use a winch [6], counter weight [7], or elastic spring [8], and lift up body weight from above via a harness connected to a wire (Fig. 2 a). Lokolift is the system that is composed of a passive elastic spring element to take over the main unloading force and an active closed-loop controlled electric drive to generate the exact target unloading force [5] (Fig. 2 b). The BWS system we have developed mounts only one actuator to provide unloading force and uplift body weight from below with pelvic support. This approach to support body weight has been less well studied; thus, a basic study on force control method is needed to realize the desired support for this system. 978-1-4244-2679-9/08/$25.00 \u00a92008 IEEE 1403 A precise and constant unloading force is believed to be an important prerequisite for BWS gait therapy in terms of conserving energy expenditure [9]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002239_j.euromechsol.2008.06.008-Figure19-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002239_j.euromechsol.2008.06.008-Figure19-1.png", "caption": "Fig. 19. Results obtained for Problem 5 without stability constraint (Nm = 3, Np = 7, Na = 6). (a) The end-effector\u2019s goal orientation is imposed. (b) The end-effector\u2019s goal orientation is free.", "texts": [ " It is interesting to observe that, in both versions of the problem, the calculated solution does not tend towards reducing the length l of the platform path. This result somewhat disagrees with the first step of the method proposed in Foulon et al. (1998), and which consists of forecasting the approximate final position of the platform by minimizing l. On the other hand, if stability is not included then calculations no longer seem to contradict the approach proposed in the above-cited reference. This is clearly apparent in Fig. 19 that gives, for both versions of the problem, the solutions obtained without considering stability. We notice the important stretching of the arms and the resulting set back position of the platform near the end of the motion. The purpose of this example is to illustrate the behavior of RPA when the number of free nodes is increased. The workspace is a (11 m \u00d7 10 m) flat floor with several obstacles (Fig. 20). We consider the minimum-time problem for the 2-link WMM with constraints on maximum torques and on stability" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002842_1.4001054-Figure11-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002842_1.4001054-Figure11-1.png", "caption": "Fig. 11 Model of the two-stage rotor: \u201ea\u2026 model of rotor stage no. 1; \u201eb\u2026 model of rotor stage no. 2", "texts": [ " 8 Forced response of the worst-case bladed disk when the blade mass and stiffness is mistuned: \u201ea\u2026 stiffness mistuned; \u201eb\u2026 mass mistuned as mentioned before. Transactions of the ASME 6 Terms of Use: http://www.asme.org/about-asme/terms-of-use n m s a a R q n g a f t e p c m F w t J Downloaded Fr 4.2 Three-DOF-Per-Sector Model 4.2.1 Natural Characteristic of the Tuned System. The second umerical example consists of a lumped parameter bladed disk odel with three-degrees-of-freedom per sector. The model conidered in Ref. 37 is here chosen for the numerical simulations, s shown in Fig. 11. The nominal system parameters for the model re given in Table 2 where k1i tun is the nominal value of k1i; see ef. 37 for detailed descriptions . The free vibration natural freuencies of isolated rotor stage nos. 1 and 2 are plotted versus the umber of nodal diameters in Fig. 12. The natural frequencies rouped into narrow bands of mode families is an interesting charcteristic of periodic structures; it is clear that there are three amilies of modes with well separated frequencies. The blade moion dominated modes of the first family of modes do not stiffen nough so they form lines that are approximately horizontal", " 9 Statistical distribution of the peak maximum amplitude hen the blade mass or stiffness is mistuned: \u201ea\u2026 mass mis- uned; \u201eb\u2026 stiffness mistuned istuning has been simulated by considering the variations in the ournal of Engineering for Gas Turbines and Power om: http://gasturbinespower.asmedigitalcollection.asme.org/ on 01/29/201 modal stiffness k1i only; for investigating the effects of interstage coupling, the results from the two-stage rotor are compared with those from the single-stage case. 4.2.2 Isolated Single-Stage Rotor. The excitation level applied to the models in Fig. 11 is the same as in Ref. 37 . The amplitude responses of the tuned isolated rotor are shown in Fig. 13. Only a DECEMBER 2010, Vol. 132 / 122501-7 6 Terms of Use: http://www.asme.org/about-asme/terms-of-use s h v d F 1 t t j r 1 w 1 Downloaded Fr ingle peak can be seen for the tuned system, and all the blades ave same amplitude. It indicates that identical frequency and ibration mode is associated with each individual blade. The results of the optimization search of the worst mistuning istributions of the isolated rotor stage nos", " The extremely large amplitude magnification factor of the frequency response spectra can be attributed directly to the effects of mistuning jump values on the mistuned forced response. We refer to this relation phenomenon as the mistuning jump-localization. The proposed method is again employed to predict variations in the mistuning sensitivity. Figure 17 shows the maximum amplitude magnification factor of the isolated rotor stage nos. 1 and 2 as the mistuning strength is increased from 0% to 20%. Recall that the amplitude magnification factor is normalized so that the maxi- he models in Fig. 11 i un /m k2 N/m k3 N/m kc N/m ,300 17,350,000 7,521,000 30,840,000 of t k1 t N 430 tuning pattern shown in Fig. 14 \u201eisolated rotor no. 1\u2026 Transactions of the ASME 6 Terms of Use: http://www.asme.org/about-asme/terms-of-use m o t s c T i m m F r t h = a c i d r i a m a j r F t J Downloaded Fr um tuned response corresponds to a value of 1.0. The sensitivity f the maximum amplitude magnification to mistuning is relaively high near the tuned value; note that as the mistuning trength is increased to 1%, the maximum mistuned response beomes as much as 80\u2013110% higher than that of the tuned system" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000326_1.1897410-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000326_1.1897410-Figure4-1.png", "caption": "Fig. 4 Geometry and forces related to one set of intermediate rollers", "texts": [ " It should also be noted that for the configuration shown in all figures, the input member must rotate counterclockwise to drive the output member through the intermediate roller assemblies. If the input member rotates clockwise, the output member will not rotate because the intermediate roller assemblies are inclined to selfactuate for counterclockwise rotation. If the input member rotates counterclockwise, driving the output member counterclockwise, and the output member speeds up in a counterclockwise direction, it will simply overrun the input and coast along at the higher speed. Figure 4 shows one set of dimensioned intermediate rollers. Relationships are necessary to calculate the angles , 1, and 2 Transactions of the ASME hx?url=/data/journals/jmdedb/27807/ on 03/23/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F in terms of radii of the various members. In conjunction with friction coefficients, these angles are critical to ensure the speed reducer\u2019s self-actuating characteristic. Consideration of the distances along horizontal and vertical directions produces the following two equations: r2 + r3 cos + r3 + r4 cos + = r5 \u2212 r4 cos + 1 + 2 r2 + r3 sin + r3 + r4 sin + = r5 \u2212 r4 sin + 1 + 2 3 When Eqs. 3 are squared and added together, the angles , 1, and 2 are eliminated, leading to an expression for = cos\u22121 r5 \u2212 r4 2 \u2212 r3 + r4 2 \u2212 r2 + r3 2 2 r2 + r3 r3 + r4 4 Similarly, and can be eliminated to solve for 1+ 2 . Equations 3 are first rearranged r5 \u2212 r4 cos + 1 + 2 \u2212 r2 + r3 cos = r3 + r4 cos + r5 \u2212 r4 sin + 1 + 2 \u2212 r2 + r3 sin = r3 + r4 sin + 5 These are then squared and the results added to produce 1 + 2 = cos\u22121 r2 + r3 2 + r5 \u2212 r4 2 \u2212 r3 + r4 2 2 r2 + r3 r5 \u2212 r4 6 Figure 4 can also be used to write the following horizontal and vertical component equations: r2 + r3 cos + r3 cos + = R cos + 1 r2 + r3 sin + r3 sin + = R sin + 1 7 where R is the distance from point Q to the contact point between rollers 3 and 4. When these are squared and added, R can be expressed in terms of the roller radii R = r2 + r3 2 + r3 2 + 2r3 r2 + r3 cos 1/2 8 Equations 7 are then rewritten to isolate terms involving , and the results are squared and added to develop the following expression for 1: 1 = cos\u22121 r2 + r3 2 + R2 \u2212 r3 2 2R r2 + r3 9 Then 2 is easily determined from Eq. 6 . Thus , 1, and 2 are determined once values are selected for the radii of all members. Journal of Mechanical Design rom: http://mechanicaldesign.asmedigitalcollection.asme.org/pdfaccess.as In order for the device to work properly, the rolling elements must not slip. Appropriate values for the coefficients of friction and the inclination angle of the intermediate roller sets will guarantee slip-free operation of the speed reducer. Figure 4 shows one of the intermediate roller sets including the inner roller, the outer roller, and the roller plate. The speed ratio is given by Eq. 2 . Under ideal, theoretical conditions the power loss is assumed to be zero, and the power input will equal the power output T2 2 = T5 5 10 Then substitution of Eq. 2 into Eq. 10 will produce the following relationship: T2 = r2 r5 T5 11 Furthermore, since the input and output power are transmitted through friction forces T2 = f3r2 T5 = f4r5 12 When the relationships of Eq. 12 are substituted into Eq. 11 , it becomes apparent that f3 = f4 = f 13 Summation of forces along x and y directions in Fig. 4 will produce two scalar equations, and summation of torques about point Q will add a third F3 cos 1 + f sin 1 \u2212 F4 cos 2 + f sin 2 = 0 14a \u2212 P \u2212 F3 sin 1 + f cos 1 \u2212 F4 sin 2 \u2212 f cos 2 = 0 14b fr2 \u2212 PR \u2212 fr5 = 0 14c where R is given by Eq. 8 . Equation 14c may be used to eliminate P in Eq. 14b , and the resulting pair of equations are solved to determine expressions for F3 and F5 F3 = r5 \u2212 r2 R cos 2 + cos 1 + 2 \u2212 1 sin 1 + 2 f 15a F4 = r5 \u2212 r2 R cos 1 \u2212 cos 1 + 2 + 1 sin 1 + 2 f 15b To prevent slip between the input member and the inner roller as well as between the outer roller and the output member, the following relationships must be satisfied: f 3F3 f 4F4 16 where 3 is the coefficient of friction between the input member and the inner roller and 4 is the coefficient of friction between the outer roller and the output member" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003253_ems.2011.79-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003253_ems.2011.79-Figure5-1.png", "caption": "Figure 5. The Structure of the RNNC.", "texts": [ " The model of the RNNI is as follows: The input and output layers of the jth neuron : I (t) I (t) \u2211 WR(t)O (t 1)I , j 1, , m (1) O (t) f I (t) I ( ) I ( )I ( ) I ( ) , j 1, \u2026 , m (2) The input layer and output layer of the kth neuron : ( ) ( ) \u2211 ( ) ( ) , 1, \u2026 , (3) Error function: EI(t) \u2211 (x (t) O (t))I , (4) where x (t), k 1, \u2026 , n , are the outputs of the system. The weights WR(t) and WO(t) use the steepest descent algorithm which equal are to: WO(t 1) WO(t) \u2206WO(t) WO(t) \u03b7IO EI( )WO ( ) (5) WR(t 1) WR(t) \u2206WR(t) WR(t) \u03b7IR EI( )WR( ) (6) The gradient of error EI(t)is represented as: EI( )WO ( ) x (t) O (t) O (t) (7) EI( )WR( ) \u2211 x (t) O (t) WO(t) 1IO (t) O (t 1) (8) RNNC is used as a controller. Figure 5 shows the basic structure of the applied recurrent neural network controller. The controller controls and regulates the output of the power source through a reference input voltage. The RNNC includes two layers, namely the input and output layers. The input layer has mc neurons where as the output layer has one neuron only. The input and output layers are called S (t) and T (t) respectively. There are five signals which are entered into the input layer; u(t 1), x (t 1), x (t 1), x (t 1). Each neuron of the input layer is joined with each other using recurrent weights VR(t)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000738_cae.20083-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000738_cae.20083-Figure4-1.png", "caption": "Figure 4 A picture of the hardware setup of the prototype. [Color figure can be viewed in the online issue, which is available at www.interscience.wiley.com.]", "texts": [ " In-order to place the writing instrument at the desired location, another smaller stepper motor was utilized for its Z-axis motion, to extend and retract the writing instrument. This system was easy to use, and capable of drawing various geometrical shapes. The drawing of complicated shapes of objects was possible through utilizing a low-cost machine vision system for image acquisition and processing. The apparatus of this system is described in Ref. [11]. A system overview is shown in Figure 3. The basic system setup is shown in Figure 3. A pictorial view of the system is shown in Figure 4. This system was designed, fabricated, and assembled by five MechUGs. The model of the two identical stepper motors and drivers in use was VEXTA CSK series two-phase type for the X and Y planes. The operating power supplies to these stepper motors were either \u00fe24 or \u00fe36 VDC. The stepper motor output was configured as Clockwise/Counter Clockwise mode. This stepper motor has an angular resolution of 1.88 and would require 200 steps to complete a revolution. A third Z-axis was formed using a salvaged stepper motor assembly from an old CDROM player. The purpose of this vertical Z-axis was for extension or retraction of a securely fastened marker pen, as the drawing element. With reference to Figure 4, a pulley and belt system was utilized for the X-axis motion of the drawing mechanism. The X-axis mechanism was mounted on a wooden arm structure. This wooden arm structure was attached to a lead-screw mechanism so that motion propagation was allowed along the Y-axis. LabVIEW 7.0 and a PCI-7344 motion controller were utilized in this system. The UMI-7764 breakout box has connectors for interfacing to two stepper motor drivers. A low cost machine vision system utilizing a CREATIVE USB webcam was used" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000736_j.mechmachtheory.2006.01.016-Figure2-1.png", "caption": "Fig. 2. Datum surfaces.", "texts": [ " This type of drive based on the above idea can be widely applied and will be discussed in detail. To crossed-axial normal circular-arc gear, the condition is comparatively loose, so the choice of the directrix is much, which is convenient to engineering application obviously, but also brings some difficulties to the determination of the directrix. For such a case, above-mentioned research course, from datum surface to the directrix, is necessary; here we discuss two aspects of this question. In Fig. 2, z1 and z2 are the rotational axes, their shortest distance A = jo1o2j is the center distance, a is its unit vector. X(1), X(2) are rotational angular velocity vector, R is the crossing angle between X(1) and X(2), thereinto jX(1)j = 1 and jX(2)j = I, I is speed ratio. Unit vector b = X(1) \u00b7 a, so {o1,abX(1)} is vertical right-hand coordinates system. In this coordinates system, a generatrix Cm is predesignated, and its equation is described as Cm: q \u00bc q\u00f0u\u00de The rotational surface, generated by rotating Cm about z1-axis, namely is the known datum surface R\u00f01\u00dep , so the expression of R\u00f01\u00dep can be represented as R\u00f01\u00dep : P\u00f01\u00de \u00bc B1\u00f0k1\u00deq\u00f0u\u00de \u00f01\u00de Here, B1(k1) is rotation group about z1-axis [5], k1 is a rotational angle, and k1, u are two independent parameters of datum surface", " (1) and referring to Appendix A, the hyperboloid of one sheet can be represented as R\u00f01\u00dep : P\u00f01\u00de \u00bc B1\u00f0k1\u00deq\u00f0u\u00de \u00bc r1e\u00f0k1\u00de \u00fe u\u00bd sin d1e1\u00f0k1\u00de \u00fe cos d1k1 \u00f038\u00de Hence, normal vector of datum surface is represented as N\u00f01\u00de \u00bc P \u00f01\u00de k1 P\u00f01\u00deu \u00bc r1 cos d1e\u00f0k1\u00de u sin2 d1k1 u sin d1 cos d1e1\u00f0k1\u00de \u00f039\u00de In Fig. 9, the circle in the middle of the hyperboloid of one sheet is called as the gorge circle, whose radius is r1. According to the property of hyperboloid of one sheet, the surface can be formed by two symmetrical straight generatrixes, as Cm, C\u00f0 \u00dem . Comparing Fig. 9 with Fig. 2, related vectors in Eq. (3) can be translated as a \u00bc i1 b \u00bc j1 X\u00f01\u00de \u00bc k1 X\u00f02\u00de \u00bc I j1 X\u00f021\u00de \u00bc \u00f0I j1 \u00fe k1\u00de \u00f040\u00de Substituting Eqs. (37) and (40) into Eq. (3) and noticing the orthogonal condition (i.e. R = p/2), the conditional expression of the line of action of hyperboloid of one sheet can be obtained: r1 cos k1 \u00fe u sin d1 sin k1 \u00bc A sin2 d1 \u00f041\u00de This expression can be represented as k1 = k1(u). Substituting it into Eq. (38), the vector expression of line of action is described as P(0) = P(1)(k1(u))" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001410_0301-679x(77)90021-4-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001410_0301-679x(77)90021-4-Figure1-1.png", "caption": "Fig 1 Gear test rig and material combinations", "texts": [ " By knowing the appearance of tooth flanks after running-in for safe running gears, it may be possible to predict service life of new gears by examining tooth flanks after running-in. This is of special interest for large gears. The aim of the investigation described in this article was to study the running-in process and to test some recognised measuring methods which could be of use both on gears in the laboratory and for gears in service. Gear test rig The gear test fig used is of the usual back-to-back type (Fig 1). The load is applied by a hydraulic torque device. Each gear box had a separate lubrication system and the inlet oil temperature could be kept constant. The running conditions, applied load, driving motor torque, temperature of the oil at inlet and outlet, etc, were continuously recorded. The test gears were made of quenched and tempered steels SIS 2216 and SIS 2240 (Table 1) with a Brinell hardness of about 250. The lubricant was a mineral oil (type Nyniis TD-35X) with a viscosity 50 cSt at 50\u00b0C" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000283_0020-7403(85)90082-7-Figure41-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000283_0020-7403(85)90082-7-Figure41-1.png", "caption": "FIG. 41a) Link Model.", "texts": [ " We see that the initial imperfection has little or no effect for small values of A, when the load is, for example, less than half that o f the elastic critical load. A similar analysis of the link model was carried out for unstable-symmetric buckling and the corresponding results are shown in Fig. 9. Now considering the strut o f Fig. 5 with an initial curvature described by r~x 14/o = Q o s i n - - (60) L T h e dynamics of symmetric post-buckling 243 ( b ) 14 A a~ 0.6 0.4 0.2 i i J -o.t5 -o.'i -0.05 0 0.05 O.I 0.15 Oo Io ol FzG. 8 (b) Impeffecdon sensitivity for stable link model of Fig. 41a). AOo:O.Ot75 FIG. 9 (a) Imperfect unstable equilibrium paths. ~ J ! 06 i A 04,~ 0 2 -ols -o', -o6s o 0,: 0 OI 015 \"\"',\" ci : I 1 FIG. 9 (b) Imperfection sensitivity for unstable link model. L4 I f: A 0.4 0.2 0 (c) , % w 2 ,14 DisoLacement in radians FIG. 9 (c) Frequency squared against load tot imperfect unstable model. we can write the strain energy as 1 ~ 4 L L 1 = 6 , L + \u00bdK (QZ - QQo), (61) where higher-order terms have been neglected. Following our earlier theory we finally arrive at \u2022 ~ h: /QV 3,:o/QV~ /Qo'~/z'~ and w a = 2 l A - l ) + 3 -~- ~ " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003643_aim.2011.6027000-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003643_aim.2011.6027000-Figure10-1.png", "caption": "Fig. 10. Internal structure of a unit", "texts": [ " Each unit of this pump can be contracted toward the center annularly similar to that in the intestinal tract. The peristaltic action of the pump can be carried out by contracting each unit according to a pattern of regular motion. It is possible to transport an inclusion by this peristaltic action. In addition, the closing rate of cylindrical tube was approximately100% and the volume exclusion rate was also approximately 100% [8]. In this study, we constructed a peristaltic pump having six units. The internal structure of a unit is shown in Fig. 10, and its specifications are shown in Table II. When the units are connected, the chamber contains an air tube supplying air to each unit; the air tube is arranged such that it saves space. Moreover, it is circularly arranged to prevent it from breaking when the unit contracts. This structure prevents the air channel from being cut. In this unit, flanges A and B form a pair. The cylindrical tube is fixed in place by interleaving one end of each unit between flanges A and B. By adopting this method, we can eliminate transport losses due to the influence of the thickness of the flanges" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003015_oceanssyd.2010.5603906-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003015_oceanssyd.2010.5603906-Figure1-1.png", "caption": "Fig. 1. The main components of the new mid\u2013depth Lagrangian profiling float. The electronics consist of an Ocean Sever battery management system, four lithium ion 95 W\u2013hr batteries, a NetBurner Mod 5234 microcontroller, and multiple feedback and scientific sensors. The Auto-Ballasting System consists of a motor driven piston mechanism and is driven by a controller that utilizes pressure and altitude readings. The optional CTD sensor can be attached depending on mission requirements. The top end cap contains housings for a GPS/Iridium antenna and a strobe. Additional housings are provided for other communications, such as radio and pre\u2013mission networking.", "texts": [], "surrounding_texts": [ "Index Terms\u2014float, drifter, sensor platform, control system, state\u2013space feedback, nonlinear control\nI. INTRODUCTION\nWith the exponential growth in technology, the philosophy of a network of systems and sensors is driving advancements in ocean engineering systems every day. As part of this philosophy, Lagrangian floats play an important role with their ability to operate remotely and autonomously for long time scales over large areas [1]. Ideally, Lagrangian floats follow the three dimensional motion of a fluid parcel and track its motion over time. This concept allows for insights into ocean circulation which both compliments stationary instruments and provides observations not obtainable with static sensors alone.\nWhile deep Lagrangian floats have a rich evolution in tracking global currents [2], there has been little emphasis on adapting the technology for shallow water coastal regions where the floats are subject to much lower pressures, but in many cases, more dynamic environments. These factors contribute to fundamental differences in these two classes of Lagrangian floats. While the deep Lagrangian floats must limit their buoyancy engines to lower speeds and higher mechanical advantages, mid\u2013depth and coastal floats have much different restrictions. This enables their buoyancy engines and ultimately their performance to be improved drastically. These performance advancements are important because shal-\nlow waters can present a much more variable and turbulent environment. As an example, Eric D\u2019Asaro, has deployed \u201cMixed Layer Lagrangian Floats (MLFII)\u201d [3] with much larger volumes for buoyancy control. The speed and number of times these floats can profile is far greater than that of deep Lagrangian floats. Still, these floats are not well equipped for shallow water regions where the sea floor is within the float\u2019s operating range. Additionally, the control systems, which are comparable to those in deep Lagrangian floats, are often limited to basic trajectories with long rise times [1].\nTo satisfy this niche, the mid\u2013depth profiling float has been designed to handle shallow water challenges, while advancing trajectory types, performance, and efficiency. Beyond these improvements, the float is designed to perform missions over several days and report back to the surface for retrieval. Be-\n1 978-1-4244-5222-4/10/$26.00 \u00a92010 IEEE", "cause this float is a modular design, the mission types can vary greatly. One mission could be primarily based on near\u2013bottom work, while the float maintains a constant altitude. Another mission could consist primarily of rate controlled profiling with CTD readings between two user\u2013defined depths. These new mid\u2013depth Lagrangian profiling floats are designed to be a very versatile platforms that can support many oceanography studies.\nTo control buoyancy and therefore vertical movement, the float is equipped with an Auto\u2013Ballasting System (ABS) that can drive water into or out of a piston tube within the float (Fig. 2). The volume of the piston tube with respect to the float is quite large so that the ABS can maintain performance in dynamic environments often associated with shallow waters. Essentially, by adjusting the volume of water within the piston tube, the effective volume of the float changes, thereby changing the buoyancy.\nThe total active volume (Va) for buoyancy control is a function of the piston stroke (h) and plunger diameter (dp) and area. This large volumetric change is required so that the float can handle dynamic environments and also so that the float can properly lift its iridium antenna high enough out of the water. This volume and its effect on the float dynamics is described in the following sections.\nThe Animatics SmartMotors SM2315DT and SM2316DTPLS2 (used in two new prototype Lagrangian floats) are\nDC brushless servo motors that include microcontrollers with embedded PID control systems. [4] The PID controller and an integrated encoder are used to track position, velocity, or torque inputs. For position commands, the SmartMotor follows a trapezoidal velocity profile, meaning that it will accelerate at a specified constant until it reaches the maximum selected velocity. The velocity will then decelerate at the same rate until it again reaches zero such that when the motor stops moving, the position command will be reached. This yields a trapezoidal velocity profile which is standard for many servo motors. Fig. 3 shows this behavior.\nControlling the motor in velocity mode follows a similar procedure. In both modes, a maximum velocity and constant acceleration must be defined. These values are critical to the control system performance and efficiency and are discussed later in this paper. Fig. 3 also shows that the motor is inherently nonlinear in both control modes. In order to achieve linear control, the input signals must acknowledge the maximum position, velocity, and acceleration specifications. By driving the motor in position mode, while controlling the amplitude and period of a sinusoidal position input, linear control limits were established. However, when subject to a sinusoidal position command, the velocity showed much fluctuation due to the trapezoidal operation. Constantly accelerating the piston under load would contribute to additional power consumption. Velocity input tests produced much smoother acceleration profiles and simpler linear control restrictions. Although velocity control is not as intuitive, its advantages in terms of power consumption and linear control are significant. Therefore, the control system design utilizes velocity input control.\nThe system dynamics provide a foundation for specifying all system components in order to ensure the Lagrangian float\u2019s" ] }, { "image_filename": "designv11_20_0002278_j.msec.2007.10.016-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002278_j.msec.2007.10.016-Figure7-1.png", "caption": "Fig. 7. (A) Cyclic voltammograms recorded in a solution of CH2Cl2+0.1 M TBAP o benzoic anhydre (10 mM) and 1-methylimidazole (5 mM) without O2. Scan rate: 10 presence of benzoic anhydre (10 mM) and 1-methylimidazole (5 mM) in presence o (10 mM) and 1-methylimidazole (5 mM) and O2. Scan rate: 10 mV/s, (B) Cyclic volta presence of 1-methylimidazole (5 mM) and O2. Scan rate: 10 mV/s, (a) without ben", "texts": [ " In this regard, special attention has been devoted to electropolymerized Mn porphyrin films as catalytic electrode materials for biomimetic oxidations with oxygen. Recently, Albin and Bedioui have electrochemically demonstrated the existence of an oxomanganese (V) porphyrin as catalytical active species in the reduction of oxygen and H2O2 [28]. Consequently, the electrochemical behavior of the poly MnTPFPPy electrode for the biocatalytic activation of oxygen in the presence of benzoic anhydride was examined in CH2Cl2+0.1 M TBAP. Fig. 7A shows the cyclic voltammograms of a thin poly MnTPFPPy electrode (\u0393Mn=7.810 \u221210 mol cm\u22122) recorded in the presence of benzoic anhydride (100 mM) and 1-methylimidazole (5 mM) as axial base (Fig. 7A, a). In the presence of dissolved oxygen, the reduction reaction is illustrated by the increase in the cathodic peak current intensity of the MnIII/II couple (Fig. 7A, b). In comparison with the bare glassy carbon electrode (Fig. 7A c), a strong catalytic wave is observed at less negative potential. This phenomenon was attributed to the activation of oxygen via the formation of high valent manganese oxo-porphyrin species (MnV=O). This chemically generated species are immediately reduced into Mn II porphyrin at the potential value corresponding to the MnIII reduction. To corroborate the activator role of benzoic anhydride on the electrocatalytic activity of the poly Mn porphyrin film, cyclic voltamograms were collected in aerated CH2Cl2 with and without benzoic anhydride (Fig. 7B). It clearly appears that the addition benzoic anhydride induces a marked enhancement of the cathodic wave in the presence of oxygen (Fig. 7B, b). This confirms that the observed catalytic phenomenon is not due to the mediated reduction of oxygen. Since the generation of MnV=O in aqueous media was reported in the reaction of Mn(III) porphyrin with H2O2 or oxygen in the presence of anhydride [28,29], the potential electrocatalytic properties of the poly MnTPFPPy electrode towards the detection of H2O2 in the presence of oxygen were investigated by cyclic voltammetry in phosphate buffer 0.1 M (pH 7). Under argon and without H2O2, the modified electrode exhibits a reversible MnIII/II peak system at \u22120" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001744_j.optlastec.2006.12.009-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001744_j.optlastec.2006.12.009-Figure2-1.png", "caption": "Fig. 2. Schematic of formation of a cross-sectional clad profile.", "texts": [ " The interface of solidification is a three-dimensional surface involving the curve AB. Its edge is the boundary of the liquid phase, solid phase and gas phase, of which the back part is shown in Fig. 1 as three-dimensional curve bAd. This back part will contribute to the surface of the formed clad bead. If a series of boundaries are identified, the surface of a clad bead can be plotted. If the clad bead surface is exposed, it is certain that the corresponding cross-sectional profiles are disclosed. In Fig. 2, the origin of the coordinates is at the longitudinal center line of the bottom of a clad bead, x-axis is along the orientation of the laser scanning, positive z-axis is above the bottom of a clad bead, y-axis points from the left of the positive x-axis to the right. Assumed that a section and planview of a pool. cross-sectional profile is located on plane y\u2013z, which is a cross section of the clad bead, it can be seen that all the points at the cross-sectional profile are the intersection ones between plane y\u2013z and the numerous former pool edges" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000980_1-4020-3559-4-Figure25-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000980_1-4020-3559-4-Figure25-1.png", "caption": "Figure 25. General motion of a flexible body.", "texts": [ " The outcome of the experimental test, which is rather similar to the outcome of the simulation, is further used to validate the vehicle model [21]. Let the principle of the virtual works be used to express the equilibrium of the flexible body in the current configuration t+\u2206t and let an updated Lagrangean formulation be used to obtain the equations of motion of the flexible body [23]. Let the finite element method be used to represent the equations of motion of the flexible body. Referring to Fig. 25, the assembly of all finite elements used in the discretization of a single flexible body results in its equations of motion written as [6] \u23a1\u23a3Mrr Mrf Mrf M\u03c6r M\u03c6\u03c6 M\u03c6f Mfr Mf\u03c6 Mff \u23a4\u23a6\u23a1\u23a3 r\u0308 \u03c9\u0307\u2032 u\u0308\u2032 \u23a4\u23a6 = \u23a1\u23a3 gr g\u2032 \u03c6 g\u2032 f \u23a4\u23a6\u2212 \u23a1\u23a3 sr s\u2032\u03c6 s\u2032f \u23a4\u23a6\u2212 \u23a1\u23a300 f \u23a4\u23a6 \u2212 \u23a1\u23a30 0 0 0 0 0 0 0 KL + KNL \u23a4\u23a6\u23a1\u23a30 0 u\u2032 \u23a4\u23a6 (29) where r\u0308 and \u03c9\u0307\u2032 are respectively the translational and angular accelerations of the body-fixed reference frame and u\u0308\u2032 denotes the nodal accelerations measured in body fixed coordinates. The local coordinate frame \u03be\u03b7\u03b6 attached to the flexible body, is used to represent the gross motion of the body and its deformation" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002728_jae-2010-1099-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002728_jae-2010-1099-Figure8-1.png", "caption": "Fig. 8. Measured angle.", "texts": [ " The data of the excitation waveform were sent to the D/A converter and amplified with the power amplifier and supplied to the exciting coil. The pickup voltage from the search coil was inputted the personal computer after passed through the pre-amplifier and the A/D converter, and the magnetic flux density was calculated with (2). The magnetic field strength was calculated by the excitation current method with (3). The current was measured with the shunt resistance. Then the iron loss was evaluated by (1). In the measurements, the magnetic flux density waveform was controlled to be sinusoidal by feed-back control. Figure 8 shows the measured angles (the line corresponds to the symmetrical line). The each inside yoke was moved by 20 degree step. The exciting frequency was 50Hz. Because the inside yokes vibrated due to alternating electromagnetic force, they were fixed by clamping. The magnetic flux density in the core material was changed from 0.5T to 1.4T at each angle. The measurements at the each yoke position were made three times and the average value was calculated. The comparison of the iron loss of the punched process with that of the laminated process is shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000869_itherm.2006.1645406-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000869_itherm.2006.1645406-Figure5-1.png", "caption": "Fig. 5: Thermal model with adjusted load condition for a solid block", "texts": [ " The following can be deduced: * , 2 2 0 2lq lx lI q x A \u2217 = \u23a1 \u23a4= \u2212 \u2212 \u22c5 =\u23a2 \u23a5\u23a3 \u23a6 (11a) * ,0 02 2 Losses q x PlI q x A\u2217 = \u23a1 \u23a4= \u2212 \u2212 \u22c5 = \u2212\u23a2 \u23a5\u23a3 \u23a6 (11b) * , 2 2 Losses q l x l PlI q x A\u2217 = \u23a1 \u23a4= \u2212 \u2212 \u22c5 = +\u23a2 \u23a5\u23a3 \u23a6 (11c) This means that the heat flow through the elements 1th R and 2th R resulting from internal losses changes between 0 and 2 Losses P \u00b1 . In the following, this part of the heat flow through the elements 1th R and 2th R will be set being constant and equal to the mean value (please have in mind that the direction of 1m I as shown by arrows in figure 5 determines the sign in the following equation): 1 2 1 4 Losses m th PT TI R \u2212 = \u2212 + (12a) 1 2 2 4 Losses m th PT TI R \u2212 = + + (12b) According to the above assumptions, the complete solid block can be modelled with its corresponding thermal network with two thermal resistances, 1th R and 2th R , and one heat generator source, P\u2217 , connected to the midpoint of the element, see figure 5. The value of the load applied as heat generator in the model is: 2 LossesP P\u2217 = (13) that means only one-half of the total power losses of the solid block. Applying half of the losses of the real system in the thermal network, the temperature at the midpoint of the solid block is: 2 1 1 1 1 1 1 2 8m th m Losses th T TT T R I T P R\u2212 = + \u22c5 = + + \u22c5 (14) Using the above expression, the temperature at the midpoint of the solid block for 1 2T T= is: 1 11 20 10 30m th mT T R I C C C= + \u22c5 = \u00b0 + \u00b0 = \u00b0 which is the same like for the ANSYS calculation and the exact analytical calculation" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002337_1.3147211-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002337_1.3147211-Figure2-1.png", "caption": "Fig. 2. Forces on pendulum 1 at equilibrium. F 0 is the force of the spring, and R is the tension in the rod.", "texts": [ "2 Only the first of the two bodies, 1 and 2, will be considered because the dynamics of the second can be deduced by symmetry. At equilibrium the spring applies a horizontal force F 0 =kcd 0 on 1, where d 0 is the rest position of 2 relative to 1.7 For small oscillations around 0, pendulum 1 moves along \u02c6 the direction x1. 834 Am. J. Phys. 77 9 , September 2009 http://aapt.org/ajp Downloaded 02 Oct 2012 to 136.159.235.223. Redistribution subject to AAPT The equilibrium situation for pendulum 1 is shown in Fig. 2. If S =mg +F 0, we have S = mg/cos 0. 1 The force of the spring varies according to the relative displacements of the bodies and couples the motion of the two bodies. The interaction and the unperturbed behavior are usually treated separately. In the unperturbed case pendulum 1 moves under the action of the horizontal constant force F 0 =kcd 0.8 A constant F 0 yields a constant S . For these conditions pendulum 1 makes small harmonic oscillations around 0 with frequency9 p = g L cos 0 . 2 This motion is related to the harmonic force P along x\u03021 given by P = \u2212 m p 2x1 x\u03021" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001657_bfb0119410-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001657_bfb0119410-Figure3-1.png", "caption": "Figure 3: Experimental Rover Dimensions", "texts": [], "surrounding_texts": [ "The three degree-of-freedom manipulator mounted on the front of the rover is shown in Figure 4. The manipulator's light weight (approximately 16 ounces) is achieved by using low-weight aluminum members and small, highly geared motors. The joints are driven by MicroMo DC motors with gear ratios of 2961:1, 3092:1, and 944:1 at the trunk, shoulder, and elbow joints, respectively. With the high gear reduction, the manipulator is capable of exerting large forces. In some configurations, it can exert a force equal to one-half the rover weight. This highforce capability could be useful for manipulator-aided mobility or trap recovery. The base of the arm is mounted to a six degree-of-freedom force-torque sensor, which is used for control (see Section 4.1) [8]. which is used for control (see Section 4.1) [8]." ] }, { "image_filename": "designv11_20_0001893_s00170-008-1850-5-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001893_s00170-008-1850-5-Figure6-1.png", "caption": "Fig. 6 The resultant cutting force component variable zoom in the three cutting directions; test case using parameters ap = 5 mm, f = 0.1 mm/rev, and N = 690 rpm", "texts": [ " Also, in agreement with research on the dynamic cutting process [28], we note that the self-excited vibration frequency is different from the workpiece rotational frequency, which is located around 220 Hz. 2.3 Forces decomposition The force resultant components detailed analysis highlights a plane in which evolves a variable cutting force Fv around a nominal value Fn (see further). This variable force is an oscillating action (Fig. 5) that generates u tool tip displacements and maintains the vibrations of elastic system block-tool BT [7] (Fig. 6). Thus, the cutting force variable (Fig. 5) and the self-excited vibrations of elastic system WTM are interactive, in agreement with research work [18, 28]. The cutting force variable part can be observed and compared. Not to weigh down this part, the cutting forces analysis is voluntarily below restricted at only two different situations: \u2013 Stable process using cutting depth ap = 2 mm (Fig. 7a), \u2013 Unstable process (with vibrations) using cutting depth ap = 5 mm (Fig. 7b). The vibration effects on the variable forces evolution are detailed in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.87-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.87-1.png", "caption": "Fig. 14.87 Mobile hanging scaffold with electric drive", "texts": [ " Thanks to the platform\u2019s sectional design, hanging scaffolds of different shapes (angular and arched) can be formed, as shown in Fig. 14.85. The vertical motion of the scaffolds is effected by an electric motor, cardan shafts, and winches. The possible ways of anchoring the booms are shown in Fig. 14.86, and the specifications of sectional hanging scaffolds are listed in Table 14.14. Mobile Hanging Scaffolds Mobile hanging scaffolds can be moved horizontally on an industrial car without disassembling and reassembling the booms from which they are suspended. A typical mobile hanging scaffold design is shown in Fig. 14.87. The vertically moving work platform is suspended by steel cables from booms mounted in a swinging mode on an industrial car. Drives for traveling on a track laid on the roof of a building and for hoisting the platform are installed on the industrial car. The scaffold can be steered from both the platform and the roof. Scaffolds of this type usually have a hoisting capacity of about 300 kg and are capable of an elevation of 100 m. Mobile cranes are intended for lifting and lowering loads and transferring them in the horizontal plane [14" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001689_gt2008-51204-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001689_gt2008-51204-Figure2-1.png", "caption": "Fig. 2 Deformation in balls and races", "texts": [ " ' 'C M K\u03b1 \u03b2= + With the inclusion of proportional damping in the model, the element equation (5) takes the form [ ] [ ] { } [ ]{ } [ ]{ } { } ' 1 ' (1 ) (1 ) t r e e e M M q G q k q Q \u03b1 \u03b2 \u23a1 \u23a4+ + +\u23a3 \u23a6 + + = Assuming perfect rolling of balls on the races, the angular velocity of the centre of the ball is equal to angular velocity of the cage and is given by ( ) i cage rotor o i R R R \u03c9 \u03c9 \u239b \u239e = \u239c \u239f+\u239d \u23a0 (8) The varying compliance frequency is given by vc cage bN\u03c9 \u03c9= \u00d7 , (9) where Nb is the number of balls. From equation (8), we can write vc rotor BN\u03c9 \u03c9= \u00d7 ; where I b I O RBN N R R \u239b \u239e = \u239c \u239f+\u239d \u23a0 . (10) The BN number depends on the specifications and dimensions of the bearing. The dotted circle represents the boundary of the balls in Fig. 2 and the innermost circle represents the inner race in displaced position. When centre of the inner race moves from O to O' as indicated in Fig. 2, and radial displacement in any direction \u03b8, due to displacement (O O\u2019) is more than the radial internal clearance (\u03b30); there is interference between inner (7) Copyright \u00a9 2008 by ASME rms of Use: http://www.asme.org/about-asme/terms-of-use Do race and balls. Consequently, the elastic deformation of balls and races takes place. In Fig.2, ( x , y ) denotes the displaced position of the centre of the inner race. From simple geometrical analysis, the expression for elastic deformation (\u03b4(\u03b8)) is 0( ) ( )xcos ysin\u03b4 \u03b8 \u03b8 \u03b8 \u03b3= + \u2212 (11) The expression inside the bracket is greater than zero in the angular zone \u03b8LZ., as shown in Fig.2. In this zone, the ball-race contact deformation generates a restoring force with nonlinear characteristic because of Hertzian contact. ( ( ))n bF C\u03b8 \u03b4 \u03b8= ; 3 / 2n= (12) The value of Cb (N/m3/2) and n (dimensionless) are arrived at by performing the elastic analysis of the Hertzian contact between ball and races [18]. For the dimensions of the ball bearing considered in the present modeling, the computed value of Cb is 3.52x109. Following equations (11 and 12), the restoring force generated by ith ball is, 3 2 0( ) i b i iF C xcos ysin\u03b8 \u03b8 \u03b8 \u03b3= + \u2212 In case of \u03b4(\u03b8i) > 0, ball at angular position \u03b8i is loaded giving rise to restoring force F\u03b8i and if \u03b4(\u03b8i) < 0, the ball is not in the load zone, and restoring force is set to zero" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000224_1.2114987-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000224_1.2114987-Figure1-1.png", "caption": "Fig. 1 Principle of optical triangulation.", "texts": [ " Method/instrument Gear parameter obtained Single probe runout check radial and tangential Gear runout \u2022 Pitch measuring instrument Pitch variation \u2022 Tooth space comparator Involute profile measuring instrument \u2022 Interchangeable base circle disks Tooth profile \u2022 Variable base circle \u2022 Tooth thickness caliper Tooth thickness \u2022 Addendum comparator \u2022 Span measurement \u2022 Measurement over pins Table 2 Advantages and limitations of Method Advantages Analytical discrete gear measurement \u2022 Direct meas most parame \u2022 Individualize of gear param \u2022 Higher degr Functional gear checking \u2022 Single instr \u2022 Reduced m Optical Engineering 103603-2 ed From: http://opticalengineering.spiedigitallibrary.org/ on 05/15/2015 Term Theoretical Work he optical arrangement, shown in Fig. 1, illustrates the rinciple of dimensional measurements using optical trianulation. The equation = Y f cos f sin + \u2212 Y cos + 1 ives the relationship between the dimensions Y and Y . he dimension Y is the displacement made by the measured urface from position AB to position A B , while Y is the isplacement of the spot image produced by the imaging ens. Figure 2 illustrates the principle of optical triangula- easurement and inspection methods. Limitations nt \u2022 Time-consuming \u2022 Special-purpose instrument in most cases ription \u2022 Costly time, instrumentation ccuracy \u2022 Expensive instrument g time \u2022 Requires analysis \u2022 Less accurate than discrete measurements Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003936_j.ymssp.2013.05.014-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003936_j.ymssp.2013.05.014-Figure1-1.png", "caption": "Fig. 1. The DSTO gear test rig.", "texts": [ " Nominal stresses under variable-amplitude loading are first calculated according to mechanical principles. A twostep estimation method is used to estimate parameters of the crack growth model. Finally, conclusions and concluding remarks are given in Section 6. Some experiments of gear vibration were conducted at DSTO of Australia for investigating tooth crack growth in spur gears. The test rig was run under a fixed input shaft speed of 2400 rpm (40 Hz) and a variable load from 0 kW to 45 kWwith the configuration is shown in Fig. 1. The test gearbox is driven by an electric motor through a belt drive, which supplies a full loading capacity of 45 kW. The test gear is the input pinion, which is labeled by G6 and has a spark-eroded notch (2 mm 0.1 mm 1 mm) at the root fillet across the middle of the tooth width. The gear was manufactured under the aircraft standard (AGMA Class 13), and has 27 teeth, a width of 10 mm and a rated load of 24.5 kW. The output gear shaft is coupled through another belt drive with a hydraulic dynamometer where the torque load can be generated" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002984_iros.2010.5650918-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002984_iros.2010.5650918-Figure4-1.png", "caption": "Fig. 4. Relationship of moment and direction of motion", "texts": [ " 2(a), geometry of the each stator becomes \u03b81 = \u03c0/2, \u03b82 = \u2212\u03c0/6 and \u03b83 = \u22125\u03c0/6, and the Eq. (4) can be expressed as follows: \u03c4 = 2 3 \u221a 3 c\u03c6 0 \u2212 \u221a 3 \u22123/2 \u221a 3/2 3/2 \u221a 3/2 [ mx my ] (5) = \u2212 2 3 c\u03c6 cos ( \u03c0 2 ) sin ( \u03c0 2 ) cos ( \u2212\u03c0 6 ) sin ( \u2212\u03c0 6 ) cos ( \u22125\u03c0 6 ) sin ( \u22125\u03c0 6 ) [ mx my ] (6) = 2 \u2016md\u2016 3 c\u03c6 sin (\u03c8\u2032 \u2212 \u03b81) sin (\u03c8\u2032 \u2212 \u03b82) sin (\u03c8\u2032 \u2212 \u03b83) , (7) where, \u03c8 = atan2(my, mx) is the direction of the target moment md on the X-Y plane, and \u03c8\u2032 = \u03c8 \u2212 \u03c0/2 is the direction of the motion of lever as shown in Fig. 4(a). \u2016md\u2016 = \u221a m2 x +m2 y is a norm of the moment. Generally, the single stator torque can be expressed as \u03c4i = ci(fi) sin \u03c1i, (8) where, fi is the frequency and \u03c1i is the phase difference of two AC voltages (A and B). Here, ci(fi) is a magnitude of torque, a function of the fi. When fi to be tuned to resonant frequency fres,i, ci(fi) becomes maximum value, cmax,i. Thus, we can obtain maximum value of the stator torque when the system satisfies the following condition. fi = fres,i, and \u03c1i = \u00b1\u03c0/2. (9) Eq", " ci(fi) = 2 \u2016md\u2016 3 c\u03c6 | sin (\u03c8\u2032 \u2212 \u03b8i) | sin \u03c1i = sgn {sin (\u03c8\u2032 \u2212 \u03b8i)} (11) This type of torque control is referred as FR method. 3) Torque Control Based on Phase Difference and Frequency (HB): We control both the phase difference and the frequency with Eq. (8). That is, we control fi and \u03c1i using a following condition. ci(fi) = 2 \u2016md\u2016 3 c\u03c6 sin \u03c1i = sin(\u03c8\u2032 \u2212 \u03b8i) (12) This may be called a hybrid control[2] (in this paper, refer to HB method). A picture of the SUSM used in this study was shown in Fig. 1. Potentiometers shown in the figure measure the angle of the lever (\u03b8x, \u03b8y) via guide rails (see Fig. 4(b) and Fig. 5). \u03b8x and \u03b8y are the angles between the axis Z and the lever projected on the X-Z plane and the Y-Z plane, respectively. On the other hand, we express the position of lever as the posture of vector. The original position vector of the lever is set to k = [0, 0, 1]T when it corresponds to Z-axis. The position vector of the lever p is expressed as follows: p = Ry(\u03b8x)Rx(\u03b8\u2032y)k , (13) where, Rx(\u03b8) means the rotation matrix around the axix x through an angle of \u03b8. By solving the above equations, when the angles \u03b8x and \u03b8y are obtained from the potentiometers, the position vector of the lever, p, is expressed as follows: p = sin \u03b8x cos \u03b8\u2032y sin \u03b8\u2032y cos \u03b8x cos \u03b8\u2032y , (14) where \u03b8\u2032y = tan\u22121 {cos \u03b8x tan \u03b8y}" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003354_20100901-3-it-2016.00065-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003354_20100901-3-it-2016.00065-Figure2-1.png", "caption": "Fig. 2. Schematic sketch of the pendulum system", "texts": [ " Here, our disturbance rejection scheme applies to plants with nonminimum phase and arbitrary relative degree, while the dynamic order of the adaptive law is equal to three for each harmonic, which compares favorably with known * This work was supported by the RFBR (projects 09-08- 00139-a), the ADTP (project 2.1.2/6326), and the FTP (projects NK-92P/5: P498/05.08.2009, NK-495P/1: P127/13.04.2010). results (Hou (2005), Marino et al. (2003), Xia (2002)). To demonstrate the efficiency of control law that can reject the unknown disturbance in condition of input delay we construct the special mechatronic control kit (see Fig. 1 and Fig. 2) and consider stabilization problem for mechatronic plant represented by the reaction wheel pendulum on movable platform. The disturbance is created by hand moving platform in horizontal plane. Complex of Mechatronics system Inc., disposable of Cybernetics and Control System Laboratory of SPbSU ITMO has been used to demonstrate efficiency of proposed algorithm (see Fig. 1). 978-3-902661-80-7/10/$20.00 \u00a9 2010 IFAC 481 10.3182/20100901-3-IT-2016.00065 Mechanical part of plant represents single-link pendulum fixed at the pivot pin with the reaction wheel situated at the opposite end of pendulum. Platform of a pendulum is movable (see Fig. 2). Moving of pendulum is provided by changing a direction and rate of turn the reaction wheel. Wheel rotation is controlled by regulating an input voltage in DC-motor fixed together with wheel. It is necessary to note, that platform of system possesses high inertia in compassion with a pendulum itself. We use the program buffer to imitate the input delay. Actually one can ask why authors make artificial delay. The answer is to built all conditions of the considered problem from mathematical point of view and to apply proposed control scheme to this complex system", " \u00bb(2k\u00a11)(t) 3 7 7 7 5 . (46) Since relations (13) are achieved, then control law (12) provides purpose of control (8). Thus, we get realizable regulator (5), (29) \u2013 (31), (35), (44) \u2013 (46), solving the problem of compensation of non-measurable uncertain disturbance in condition of delaying control signal. VI. EXPERIMENTAL RESULTS To demonstrate the efficiency of control law that can reject the unknown disturbance in condition of input delay we construct the special mechatronic kit (see Fig. 1 and Fig. 2) and consider stabilization problem for mechatronic plant represented by the reaction wheel pendulum on movable platform. Controlling of the pendulum angle is provided by changing a direction and rate of turn of the reaction wheel. Wheel rotation is controlled by regulating an input voltage in DCmotor fixed together with the wheel. Mathematical model describes physical character of the mechanical part of plant without movement of platform. It looks like: \u00c4\u00c3(t) + a sin\u00c3(t) = bpu(t), (47) where \u00c3 is an angle of the pendulum and the measured output, a = 60, bp = 0:5 are complex parameters of the pendulum, received from preliminary identification procedure" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001915_j.triboint.2008.09.003-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001915_j.triboint.2008.09.003-Figure9-1.png", "caption": "Fig. 9. Pressure distribution on the shaft surface for the case of rotary speed 4000 rpm.", "texts": [ " The larger deviation between the computational and measured pumping rates at higher rotary speeds is caused by the assumption of constant oil film thickness used in the computational model. It is known that the oil film thickness will increase with the rotary speed. However, for the first attempt to use CFD to predict the pumping rate and to elucidate the pumping mechanism, the result is satisfactory. The role of increasing oil film thickness with the rotary speed in predicting the accurate pumping rates will be investigated in the future study. Fig. 9 shows the pressure distribution on the shaft surface in the computational domain. It is seen clearly that the hydrodynamic effects cause the oil pressure to rise as it approaches the converging space between the rib and lip. The highest pressure occurs at the corner of the converging space. The high pressure fans out across the converging space and most of the lip band. Vacuum pressure also happens on the leeward side of the rib. Fig. 10a shows the pressure distribution on the selected plane in the inset for the typical case of rotary speed 4000 rpm, while Fig", " The negative stagnation line moves inside the lip microchannel which induces a leakage from the oil film to the air-side chamber, as seen in Fig. 13b. This leakage of very small flow rate is confined in a very small region close to the periodic boundaries in the computational domain. The pressure difference between the high pressure inside the lip micro-channel and the low pressure on the leeward side of the rib causes this minute leakage, illustrated at the junction of the rib and the lip at the lower periodic boundary in Fig. 9. Note that the leaking oil is supplied by the upstream pumping from the air-side chamber, instead of the natural leak from the oil-side chamber. The right branch of streamlines that diverges from the negative stagnation line inside the lip micro-channel still goes along the lip micro-channel and finally into the oil-side chamber, as illustrated in the Figs. 13b,c. The Go\u0308rtler\u2013Taylor vortex generated at the rotating shaft surface in the oil bulk is again clearly demonstrated in Fig. 13c. In this paper, the flow field of a radial ribbed helix lip seal around the contact region between the seal lip and the shaft surface in the environment of a pumping-rate test rig, where both air and oil sides are filled with oil of equal pressure initially, was simulated" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002930_s00604-009-0257-9-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002930_s00604-009-0257-9-Figure1-1.png", "caption": "Fig. 1 Schematic diagram of the FIA-CL for studied drugs determination. S, sample solution; C, carrier stream; R1, KMnO4; R2, HCHO in acidic medium; P1 and P2, peristaltic pump; V, valve; W, waste solution", "texts": [ "xaremex. com) controlled by a personal computer. A mixing coil was made by coiling a glass tubing (100 mm\u00d71.0 mm i.d.) into a spiral disk shape and placed close to the photomultiplier tube (PMT). The PMT was placed close to the flow cell and housed in a light-tight ferrous cylinder and was operated at \u2212800 V. A WGY-10 spectrofluorophotometer (Tianjin, China) was used for recording the chemiluminescence spectra, which was used in the luminescence mode. A schematic diagram of the set-up was shown in Fig. 1. The flow system employed in this work consisted of two peristaltic pumps. One delivered the carrier stream, formaldehyde in acid medium and potassium permanganate solution at a flow rate (per tube) of 5.0 mL\u00b7min\u22121. The other delivered the studied drugs stream or their sample stream at a flow rate of 3.0 mL\u00b7min\u22121. The glass tube only was used in the detection tube and PTFE tubing (0.8 mm i.d.) was used to connect all components in the flow system. Fifty \u00b5L sample solution was injected into the carrier stream by an eight-way injection valve" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002615_acc.2008.4586892-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002615_acc.2008.4586892-Figure2-1.png", "caption": "Fig. 2. Illustration of Proposition 4.3", "texts": [ " Thus, if A(x0) \u2227 A(xf ) \u2208 Mc, then we have \u2200\u0398 4 A(x0) \u2227 A(xf ), \u0398 4 A(x0) and \u0398 4 A(xf ). Thus the problem reduces to verifying the condition A(x0) \u2227 A(xf ) \u2208 Mc. This can be done algebraically by checking if all the elements of the corresponding SA(x0)\u2227A(xf ) defined earlier are nonzero, or studying the spectral properties of the associated Laplacian matrix [9]. This involves verifying if the second smallest eigenvalue of the Laplacian matrix corresponding to the graph represented by A(x0)\u2227A(xf ) is nonzero. 1) Example: The simulation result in Figure 2 illustrate the result in Proposition 4.3. The initial and final desired configuration are marked a) and d), respectively. The graph representing the common minimal element between initial and final configurations \u0398 is represented by bold lines. In the example considered we have \u0398 = A(x0) \u2227 A(xf ). It can be seen that this bold graph is preserved at intermediate configurations b) and c). Proposition 4.3 provides us a condition under which the optimal solution to Problem II can easily be computed" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001664_978-3-540-77608-6_20-Figure20.4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001664_978-3-540-77608-6_20-Figure20.4-1.png", "caption": "Fig. 20.4. Schematic drawings of a Tokay gecko including the overall body, one foot, a crosssectional view of the lamellae, and an individual seta. \u03c1 represents number of spatulae", "texts": [ " The front claws are also used for locomotion. The syrphid fly uses setae on the legs for attachment. In both cases, fluid is secreted at the contacting surface to increase adhesion. The explanation for the adhesive properties of gecko feet can be found in the surface morphology of the skin on the toes of the gecko. The skin is comprised of a complex hierarchical structure of lamellae, setae, branches, and spatulae [73]. Figure 20.3 shows various SEM micrographs of a gecko foot, showing the hierarchical structure down to the nanoscale. Figure 20.4 shows a schematic of the structure and Table 20.1 summarizes the surface characteristics. The gecko attachment system consists of an intricate hierarchy of structures beginning with lamellae, soft ridges, that are 1\u20132 mm in length [73] that are located on the attachment pads (toes) that compress easily so that contact can be made with rough bumpy surfaces. Tiny curved hairs known as setae extend from the lamellae with a density of approximately 14,000 per square millimeter (Schleich and K\u00e4stle, 1986)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002239_j.euromechsol.2008.06.008-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002239_j.euromechsol.2008.06.008-Figure2-1.png", "caption": "Fig. 2. Free-trajectory planning problem for a WMM. (a) GPP task, (b) MPP task.", "texts": [ " With this simplification, the stability index \u03a6 is defined as follows: \u03a6(q, q\u0307, q\u0308) = 1 \u2212 [( x0 a )2 + ( y0 b )2] (5) where a and b are the half-diameters of the inscribed ellipse. We note that \u03a6 is a dimensionless number that takes positive (negative) values if the ZMP is inside (outside) the stable region. In particular, \u03a6 = 0 if the ZMP is at the boundary of the stable region and \u03a6 = 1 if the ZMP coincides with the MSP. The WMM is required to move freely from a given initial configuration \u03a9START = [(XS p)T, (qS a) T]T to a given final configuration \u03a9GOAL = [(XG p)T, (qG a )T]T (in the case of a GPP task, Fig. 2(a)), or to a given final end-effector\u2019s state U GOAL e = [xG e , yG e , zG e , \u03b8G e , \u03b2G e ,\u03d5G e ]T (in the case of an MPP task, Fig. 2(b)). In either case, we must find the trajectory q(t), the time history of vector \u03c4 (t) of actuator efforts and the travel time T so that boundary conditions are matched, all constraints are respected and a given performance index is optimized. In addition to nonholonomic constraints (1), the set of feasible trajectories is restricted by numerous other constraints that must be satisfied during the travel from \u03a9START to either \u03a9GOAL or U GOAL e . These constraints concern the boundary conditions, the noncollision between the WMM and obstacles, the physical limitations on kinodynamic performances and the stability of the system" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001878_j.ijsolstr.2008.06.008-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001878_j.ijsolstr.2008.06.008-Figure4-1.png", "caption": "Fig. 4. A cylindrical membrane.", "texts": [ " This is typical of cylindrical and axisymmetric membranes, for which, by reasons of symmetry, the t1 and t2 directions are known a priori, and coincide with that of the principal stretches tI and tII , since a \u00bc 0. Thus, to verify the soundness of the governing system of equations (5)\u2013(9), in the following sections we will solve cases of cylindrical and axisymmetric membranes for which some close form solutions either already exist in the literature or are easily deducible. Let us consider a cylindrical membrane built by rolling up an initially flat infinite strip of constant width L. When a pressure p acts internally, the membrane moves from its initial configuration C0 to the inflated one C (Fig. 4). The problem can be analyzed in the \u00f0O; x; z\u00de plane. Now, the only meaningful curvilinear abscissa is s1 s and the Kuhn\u2013 Tucker conditions (5)\u2013(9) become d ds T1 dx ds p dz ds \u00bc 0; d ds T1 dz ds \u00fe p dx ds \u00bc 0; \u00f013a;b\u00de k1 \u00fe c2 1 1 \u00bc 0; k2 \u00fe c2 2 1 \u00bc 0; \u00f013c;d\u00de c1T1 \u00bc 0; c2T2 \u00bc 0; \u00f013e; f\u00de which are to be solved with suitable boundary conditions. Starting with Eq. (13e), we see that if c1 6\u00bc 0, then T1 \u00bc 0, and the equilibrium equations (13a,b) will be satisfied only in the meaningless case where p \u00bc 0; vice versa, if c1 \u00bc 0 (i" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002348_1.2943299-Figure13-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002348_1.2943299-Figure13-1.png", "caption": "Fig. 13 RCRCR: \u201ea\u2026 dimension notations, \u201eb\u2026 sph linkage", "texts": [ " With Linear Constraint AB, the linkage contains two branches on both sides of the JRS sheet and contains no dead position. On the other hand, with Linear Constraint CD, there is only one branch and the branch points at C and D are dead-center positions, which divide the branch into two sub-branches identifiable with the side identification of the JRS sheet. 6.2 Spatial Linkages. Each spatial linkage connected to revolute or cylindrical joints has a spherical indicatrix 13 , which is formed by bringing all joint axes to intersect at a common point Fig. 13 . The rotatability of a spherical indicatrix is depicted by the JRS and therefore the concept of sheet and side becomes applicable to spatial linkages. This new concept is illustrated with simple RCRCR mechanisms in the following example. Example 6 (simple RCRCR mechanism). Using the D-H notations, the link and joint parameters of a RCRCR mechanism can be modeled in Fig. 13 a and defined by dual angles 13 as \u0302ij = ij + aij and \u0302i = i + Sii 1 where 2=0, \u0302ij is the skewed angle between Joint Axes i and j, and \u0302i is the joint parameter at the Joint Axis i. If the dual parts of all joint and link parameters are equal to zero, an RCRCR mechanism becomes a five-bar spherical linkage, which is the spherical indicatrix 13,29 of the RCRCR mechanism. Any constraint ex- and \u201ec\u2026\u2026 gear constraints in JRS al indicatrix, and \u201ec\u2026 analogy \u201eS33=0\u2026 to a six-bar eric SEPTEMBER 2008, Vol", "asme.org/about-asme/terms-of-use h t t r t a l f l s d a m e c w i r E o w t p E v I m a lin 0 Downloaded Fr ibited in the five-bar spherical indicatrix is also a constraint for he original RCRCR mechanism. One may observe that the exisence of the dual parts, i.e., aij and Sii, in the joint and link paameters, leads to an additional constraint equation 13 to the wo-DOF spherical indicatrix and therefore creates the effect of an dditional loop to the spherical indicatrix to form a single DOF inkage Fig. 13 c . Thus, since the concept of JRS is applicable or spherical multiloop linkages, it may be applicable to spatial inkages. Due to the striking mobility similarity exhibited in imple RCRCR mechanisms and the Stephenson six-bar linkages iscussed above, a simple RCRCR mechanism is used as an exmple to highlight the potential applicability of JRS in spatial echanisms. In a simple RCRCR mechanism Fig. 13 a , the joint paramter S33=0. Since the branch condition is not affected by the hoice of the input joint, for simplicity without losing generality, 5 is assumed as the input. With S33=0, the relationship between 1 and 5 can be expressed as 13 A sin 1 + B cos 1 + C = 0 2 here A = p1 sin 5 + p2 cos 5 + p3 B = p4 sin 5 + p5 cos 5 + p6 C = p7 sin 5 + p8 cos 5 + p9 n which Pi, i=1,2 , . . ., are functions of the linkage structure paameters. Let cos 1 = 1 \u2212 y2 1 + y2 , sin 1 = 2y 1 + y2 , y = tan 1 2 cos 5 = 1 \u2212 x2 1 + x2 , sin 5 = 2x 1 + x2 , x = tan 5 2 quation 2 can be converted to a quadratic equation in the form f B1y2 + B2y + B3 = 0 3 here B1, B2, and B3 are functions of the input angle or x", " quation 2 may be considered as a constraint imposed by a irtual spherical four-bar loop to the five-bar spherical indicatrix. n other words, in view of joint rotatability, a simple RCRCR echanism may be regarded as a virtual Stephenson six-bar link- ge, which contains a five-bar loop, i.e., the spherical indicatrix, 92303-10 / Vol. 130, SEPTEMBER 2008 om: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/28/201 and a virtual four-bar loop depicted in Eq. 2 with 1 and 5 as the common joint variables between the loops Fig. 13 c . It is further observed that if the input is given through 1 or 5, a dead-center position occurs when 3=0 or . This implies that the JRS boundary of 1 and 5 of the spherical indicatrix is the 1 versus 5 curve with 3 kept at 0 and . The 1 versus 5 curve intersects the JRS boundary at the branch points as the example in Fig. 11 b . The effect of the JRS of the spherical indicatrix to the virtual four-bar loop or Eq. 2 is illustrated with the following two linkages, which have the same link parameters: a12=25, a23=30, a34=40, a45=10, a51=32, 12=60 deg, 23 =45 deg, 34=35 deg, 45=30 deg, and 51=10 deg", " The sub-branch identification under such a condition, for Stephenson six-bar and geared five-bar as well, can be resolved elegantly by using the side identification criterion on an equivalent linkage depicting the relationship between loops in the linkages 20 . 4. The method outlined above for planar linkages is valid for their spherical counterparts as well. The further extension to spatial linkages via their spherical indicatrix further demonstrates the generality and importance of the concept. 5. Using Eq. 2 for the displacement or mobility analysis of a simple RCRCR mechanism without taking the JRS of the spherical indicatrix into consideration is like mistaking a Stephenson six-bar linkage as a four-bar linkage Fig. 13 c . Using the discriminant function of the I/O curve without considering the effect of JRS on the spherical indicatrix had been a pitfall for the mobility analysis of simple RCRCR mechanisms 2 . The equation representing the I/O curve may contain an invalid input domain e.g., Segment AB in Fig. 14 a . The I/O curve alone does not necessarily reflect the number of branches or sub-branches e.g., Fig. 14 and cannot express the one-to-one correspondence between points on the I/O curve and the linkage configurations", " This paper offers a simple way to establish a one-toone correspondence between the input domain and the linkage configurations. 7. Introducing parameters aij or Sii to a five-bar spherical indicatrix to form a simple RCRCR mechanism has the similar effect of imposing a virtual four-bar loop to the spherical indicatrix to form a virtual spherical Stephenson six-bar linkage, in which S33=0, 1, and 5 are the common joint parameters between the spherical indicatrix and the virtual four-bar loop while 2, 3, and 4 are joint parameters not in the virtual four-bar loop Fig. 13 c . Giving input to 2, 3, and 4 of a simple RCRCR mechanism has the similar effect and complexity as giving input to a joint not in the four-bar loop of a Stephenson six-bar linkage. This analogy with Stephenson six-bar linkages Fig. 13 c and therefore the treatment displayed in Fig. 14 are valid for all Duffy\u2019s simple Group II linkages 13 , in which 3=0 or will correspond to the dead-center position when the input is given through 1 or 5. It is also noted that RPSPR type mechanisms 33 are members of this subgroup. 7 Order of Motion Determining or rectifying the order of motion for a synthesized linkage to reach the prescribed discrete positions is a complex issue with the traditional mobility analysis methods but extremely simple with the proposed approach" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000608_155022891010015-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000608_155022891010015-Figure1-1.png", "caption": "FIG. 1. A schematic diagram of a rolling bearing.", "texts": [ " The results presented here have been obtained from a large number of numerical integrations, mainly presented in form of Poincare\u0300 maps, phase plots and frequency spectra. In this section, a mathematical model for analysis of the structural vibration in rolling element bearings is developed. Initially, the expressions for kinetic and potential energies are formulated for all components of rolling element bearing. The equations of motion, which describe the dynamic behavior of the complex model, are derived by considering these energies expression and the Lagrange\u2019s equations. A schematic diagram of rolling element bearing is shown in Fig. 1. For investigating the structural vibration characteristics of rolling element bearing, a model of bearing assembly can be considered as a spring mass system, in which the outer race of the bearing is fixed in a rigid support and the inner race is fixed rigidly with the rotor. A constant radial force acts on the system. In the mathematical modeling, the rolling element bearing is considered as a spring mass system, and rolling elements act as nonlinear contact spring as shown in Fig. 2. Since the Hertzian forces arise only when there is contact deformation, the springs are required to act only in compression" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.162-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.162-1.png", "caption": "Fig. 14.162 Schematic of robot for checking adhesion of tiles to building\u2019s elevation (1 securing robot to roof edge; 2 chain; 3 robot (hoisting gear motor and carrier of: tile adhesion diagnosing balls, analyzing circuit, and driving and communication control circuits); 4 power supply unit (performs various robot positioning functions); 5 winch controlling computer (position recording and communication control circuits); 6 output units (computer, X\u2013Y plotter); 7 100 V AC power supply plug)", "texts": [ " Robots for Assessing the Technical Condition of Building Structures Robots for checking the technical condition of building structures, referred to as inspection robots, are used for inspecting the following elements: \u2022 Exterior wall facings\u2022 Utility piping\u2022 Concrete structures such as water dams and bridges\u2022 Underwater structures The most numerous in this group are tiled elevation inspection robots. Depending on their function, they check the adhesion of tiles to the base or check for layer corrosion. As time passes, adhesion decreases and, since the tiles may start falling off the wall, it becomes necessary to inspect the tiled walls at regular intervals. A schematic of a robot for checking the adhesion of tiles to the building\u2019s elevation is shown in Fig. 14.162. The robot is drawn up on by a chain secured to the roof edge. The check is conducted by continuously tapping the lining with ten small balls arranged in a row and analyzing the sounds generated. The diagnosis results, including the tapping locations, are transmitted to a computer on the ground, saved on a diskette, and represented graphically. The robot can also be used to assess the adhesion of plasters. It can operate at a rate of 700 m2/day [14.53]. Besides tapping tiles with balls, other techniques are used to check the adhesion of tiles to the elevation, e" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001156_j.compstruc.2006.01.037-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001156_j.compstruc.2006.01.037-Figure1-1.png", "caption": "Fig. 1. Bending\u2013tension test machine and experimental set-up from [24]. To carry out this test, the sheet metal strip is first pre-bent, clamped into the machine and stretched around the roller to the corresponding form. The test itself is then preformed by pulling the strip from left to right via the clamps around the roller.", "texts": [ " On this basis, it has been applied to a number of structural simulations involving cyclic loading and strain-path changes which result in kinematic hardening. For example, consider the case of combined tension\u2013 bending loading of a sheet metal strip. Such a test is used in particular in springback studies. In this test (e.g., [20,24]), a straight sheet metal strip is pre-bent, clamped into the machine at both ends and stretched around a roller at the top of the machine, and then pulled from left to right around the roller to carry out the test (see Fig. 1). The advantage of this type of strain-path-change test lies in the minimization of friction since the roller rotates with the sheet metal strip as it is bent. The experimental results here are represented by the final geometric form of the sheet metal samples released from the clamps after the test which have sprang back. Photographs of such final sheet metal forms from the side for the sheet metal DP600 are shown in Fig. 2. The parameter identification was carried out by first fitting a purely isotropic hardening model to the uniaxial tension test data, then a purely kinematic model, and finally the combined model" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001679_gt2008-50305-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001679_gt2008-50305-Figure9-1.png", "caption": "Figure 9. 3-D image of turbine showing symmetric features of casting and final machined part. Symmetrically bent blade tips are highlighted between circles.", "texts": [ " A positive value is away from the surface, in the normal direction. This part is made with an investment cast process. Simply put, a sacrificial mold is created in a ceramic material that can endure the pouring temperatures of the metal. After the part is poured and cooled, the mold is forcibly removed from the metal part that has been created inside. The process may include chipping with hammers, blasting with media, or treating in chemical solutions. Due to the mold removal step, the blade tips in Figure 9 were bent with respect to the nominal blade shape, but evenly and symmetrically. The hub line, and therefore the flow volume of the blade pairs, is also even and symmetric about the rotor. For this reason, this part performed Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/27/201 well with respect to rotordynamics, but most likely had a reduction in aerodynamic performance. The bad rotor clearly shows that the cast surfaces are displaced approximately 750 \u00b5m (~.030 in.) radially from the axis defined by the machined diametral datum" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003845_1.4003270-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003845_1.4003270-Figure2-1.png", "caption": "Fig. 2 Six-bar linkage: state 2", "texts": [ " However, adjacency matrices cannot be used to identify type of joints, the kinematic orientation of the joints, the fixed links, or the DOF of the mechanism. 2.2 EU-Matrix Transformations. In 2005 Dai and Jones 7 proposed an EU-elementary matrix operation to represent the state changes of metamorphic mechanisms. EU-elementary matrix operations are the first type of matrix operation that can be used to reduce the dimension of an adjacency matrix. An E-elementary matrix is combined with a U-elementary matrix to form an EUelementary matrix operation. As shown in Fig. 2, consider the case in which link 6 of the six-bar linkage becomes fixed as a result of a pin P. The E-elementary matrix and the U-elementary FEBRUARY 2011, Vol. 3 / 011012-111 by ASME 016 Terms of Use: http://www.asme.org/about-asme/terms-of-use T g T a r d t m i f e n 0 Downloaded Fr A1 = E6U5,6 A0 E6U5,6 T 4 he matrix operation uses modulo-2 arithmetic and the results are iven in Eq. 5 A1 = 0 1 0 0 1 1 0 1 0 0 0 1 0 1 0 0 0 1 0 1 1 0 0 1 0 5 he EU-matrix transformation effectively transformed A0 to the djacency matrix for the mechanism in Fig. 2. Dai and Jones ealized that, upon a change in the number of effective links, the imension of an adjacency matrix changes as well as the order of he elements. In the previous example, A0 changed from a 6 6 atrix to A1, a 5 5 matrix. The EU-elementary matrix operation s useful in capturing this change. However, the EU-matrix transormation does not identify the type of joints, the kinematic orintation of the joints, the fixed links, or the DOF of the mechaism. 11012-2 / Vol. 3, FEBRUARY 2011 om: http://mechanismsrobotics.asmedigitalcollection.asme.org/ on 02/03/2 2.3 Improved Adjacency Matrix. In 2008, Lan and Du 16 proposed a \u201c 1\u201d element to indicate a fixed kinematic pair. In this way the dimension of the adjacency matrix remains the same after a topological change has occurred. In addition, information about the original state of the mechanism is not lost. For example, the improved adjacency matrix for the six-bar linkage in Fig. 2 is given in A2 as A2 = 0 1 0 0 0 \u2212 1 1 0 1 0 0 0 0 1 0 1 0 0 0 0 1 0 1 0 0 0 0 1 0 1 \u2212 1 0 0 0 1 0 6 This representation ensures that the dimension of the adjacency matrix remains unchanged due to a change in the number of effective links. However, it does not identify the type of joints, the kinematic orientation of the joints, the fixed links, or the DOF of the mechanism. An improved adjacency matrix can be used to determine the DOF of the mechanism, but it cannot be used to analyze whether or not a mechanism contains any partially locked kinematic chains 14 ", " x i,j : a link number 1 to n corresponding to a link con- nected to link j by a joint, where x+1 i,j x i,j j. 4. x i,j : the type of kinematic pair connecting link j to link x i,j . 5. x i,j : the orientation of the kinematic pair connecting link j to link x i,j . It should be noted that x i,j x i,j is identical to the joint code used or directionality topology matrices 17 . The notation for the type nd orientation of common kinematic pairs can be found in Fig. 3. 3.2 Representative Example. The transformation from the echanism in Fig. 1 to the mechanism in Fig. 2 will be used to llustrate how mechanism state matrices can be used to represent he topological characteristics of reconfigurable mechanisms. In his example, the values associated with the link code for SM 1,1 are as follows: 1,1 = 1 1,1 = 2 2 1,1 = 6 1 1,1 = R 2 1,1 = R 1 1,1 = Z 2 1,1 = Z he values for the link code can be found for every element of SM. Plugging the link code values into the generalized form of he mechanism state matrix yields the matrix given in Eq. 12 MSM = 2Z R,6Z R 3Z R 4Z R 5Z R 6Z R 2Z R,6V X 3Z R 4Z R 5Z R 6Z R 12 otice that each row of the mechanism state matrix corresponds o a distinct state in the mechanism" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001935_s11036-008-0122-9-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001935_s11036-008-0122-9-Figure3-1.png", "caption": "Figure 3 Structure of 1-coverage optimal solution", "texts": [ " a group of three sensors, they are complementary in the sense of coverage redundancy and feasible spectrum of Dreq min. 4.1 Overlaying 1-coverage optimal solutions Our first algorithm for the 3-coverage is based on the 1-coverage optimal placement presented in [14], which is proven to be optimal in the sense of minimizing the number of sensors required for the full 1-coverage of the target area. In the 1-coverage optimal placement, all of the sensors are regularly placed with the same intersensor distance of \u221a 3R as shown in Fig. 3. Our first solution for the 3-coverage is to carefully overlay the 1-coverage optimal placement in Fig. 3 three times with considering Dreq min across the three layers. Figure 4 depicts the idea of overlaying the first, second, and third layers. For the first layer, we simply use the 1-coverage optimal placement as in Fig. 4a without worrying about Dreq min. In this figure, the black dots marked by 1 are the positions of the 1st layer sensors and solid line circles are their sensing coverages. Connecting positions of adjacent three sensors, we form an equilateral triangle whose length of each edge is \u221a 3R" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000499_1.1857773-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000499_1.1857773-Figure1-1.png", "caption": "Figure 1. Front view of the cathode compartment.", "texts": [ " This is the first investigation looking at the effect of a cationic surfactant on the reactor engineering of O2 electroreduction in flow-by fiber-bed and reticulated electrodes. In addition to pulp bleaching, the potential applications of an efficient H2O2 electrosynthesis method at pH , 13 are numerous, including the paired electrosynthesis of benzaldehyde from toluene mediated by H2O2 and the V51/V41 redox couple18 or degradation of organic effluents, such as formaldehyde,19,20 phenol, cresol, catechol, p-benzoquinone, anilin, oxalic acid, and dyes by H2O2 .21,22 Experimental Figure 1 shows a front view of the cathode compartment in the flow-by electrochemical reactor equipped with either a graphite fiber bed ~GF, referred to also as graphite felt! or reticulated vitreous carbon ~RVC, ;39 ppc ~pores per centimeter!, i.e., 100 ppi ~pores per inch!#. Table I and Fig. 2 show certain physicochemical and structural characteristics for GF ~The Carborundum Co., Sanborn, NY! and RVC ~ERG Materials & Aerospace Co., Oakland, CA!. The Table I. Physicochemical characteristics of the three-dimensional electrodes", " On the other hand the surface of RVC showed oxidized carbon in the form of hydroxyl, ester, and carbonyl functional groups. The oxygen content of RVC surface was determined by XPS analysis as about 8 atom %. Furthermore, FTIR analysis revealed that GF is composed of graphitic carbon only, while RVC shows C-H and C-C bonds that are characteristic for hydrocarbons. The differences in surface-functional groups could impact the electrocatalytic properties of the carbon electrodes under investigation. The cathode thickness in the reactor ~Fig. 1! was 4.5 3 1023 m, corresponding in the case of GF to 38% compression ~uncompressed thickness 7.25 3 1023 m!. The superficial cathode area in the direction of current flow was 33 3 1024 m2, while the area in the direction of fluid flow was 1.1 3 1024 m2. The cathode was operated in flow-by arrangement ~the current and fluid flow directions perpendicular on each other! with co-current upward O2 gas-liquid flow. The two-phase pressure drop along the cathode height was measured with Bourdon gauges connected at the inlet and outlet ports of the cathode compartment. The GF three-dimensional electrode was placed on a stainless steel current feeder plate and fitted into Durabla\u00ae gaskets ~Fig. 1!. The reactor with the GF cathode was uniformly compressed using a torque wrench set to 5 foot-pound. In the case of the reticulated vitreous carbon cathode, a graphite fiber paper of 3 3 1024 m thickness was inserted between the stainless steel current feeder plate and RVC to assure good electric contact. The reactor with the RVC fixed-bed embedded tightly into gaskets was carefully sealed to avoid compression and crushing of the rigid reticulated structure. The anode for alkaline conditions was Ni felt (7 3 1024 m thick", " Effect of surfactant on the gas-phase pressure gradient in the graphite fiber bed.\u2014The pressure gradient required to drive a gas bubble through a liquid-filled capillary in the presence of surfactant can be estimated using the Herbolzheimer-Park model31 DpG,g* 5 DPG,g* Hc 5 0.942gCa1/.3 lb Hc \u2022 rb 2 @B-1# where DPG,g* and DpG,g* are the pressure drop ~Pa! and pressure gradient ~Pa m21!, respectively, associated with the movement of the bubble in the liquid-filled capillary with surfactant present, Hc is the effective height of the 3D cathode ~m! ~Fig. 1!, lb and rb are the length and menisci radius of the bubble, respectively ~m!, g is surface tension ~N m21!, and Ca is the dimensionless capillary number defined as Ca 5 mLub,eff g @B-2# with ub,eff effective gas bubble velocity in the capillary filled with liquid ~m s21! and mL the liquid dynamic viscosity ~Pa s!. The effective gas bubble velocity in the liquid-filled portion of the bed can be estimated as ub,eff 5 uG \u00abbL @B-3# where uG is the superficial gas velocity ~m s21!, bL is liquid hold-up ~Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003640_j.apm.2011.08.005-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003640_j.apm.2011.08.005-Figure2-1.png", "caption": "Fig. 2. Two dimensional model of the two wheeled skateboard. The coordinate frame is defined so that the y axis coincides to the central axis of the initial skateboard position, and the x axis becomes normal to the y axis. (X,Y) is the CoM of the skateboard, h is the orientation of its central axis from the y axis, (Xwi,Ywi) is a wheel position, ai is an orientation of the front and rear wheel, FX and FY are respectively the force form the rider on the board in the x and y direction, FN is its normal component to the current skateboard orientation, T is a moment from the rider, Fci and FFi is a component of the ground reaction force at the wheel in the normal and tangential to the wheel direction. Here, i = 1, 2 in which i = 1 denotes front side, and i = 2 does rear side.", "texts": [ " Thus, if the wheel orientation is directly regarded as input instead of the weighted position, the twist between the plates is ignored. The alternatingly back-and-forth motion of the rider\u2019s feet on the board never produces net force parallel to the board orientation because the parallel forces are canceled each other following the action\u2013reaction law of the force. This means that both the yaw moment and normal force are input forces from the rider. Based on these assumptions, the board motion is expressed by a two-dimensional model, as illustrated in Fig. 2. The motion of this model is described by a first-order, six-dimensional, differential equation with two velocity constraints (see Appendix A). In this motion equation, the inputs to the board are the yaw moment T, the normal force to the board FN, and the wheel orientations a1 and a2. To simulate the board motions, a time profile of these inputs is required. We obtain them except T by approximating measurement data of the skateboard motion, and determine T so that the measurement data and the simulation data will match well" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001210_jmes_jour_1980_022_016_02-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001210_jmes_jour_1980_022_016_02-Figure1-1.png", "caption": "Fig. 1. Mechanical face seal", "texts": [], "surrounding_texts": [ "Since Denny (I)$ demonstrated the existence of a lubricating fluid film between the mating faces of mechanical face seals (Fig. l), many attempts have been made to find mechanisms which enable face separation. Some of these mechanisms which were recently summarized by Ludwig and Greiner (2), include hydrostatic, hydrodynamic, and thermal effects. In each of these models the geometry of the mating faces is a major factor affecting seal performance. The basic geometries treated so far in the seals literature are mainly the flat parallel faces, wavy faces, flat faces with angular misalignment, and flat axisymmetric coned faces. These basic geometries were, in general, analysed separately and provided good insight into the various possible mechanisms of seals\u2019 operation. However, isolated basic geometries as described above are very unlikely in reality and interactions of the various geometries and modes of operation seem to be inevitable. The first attempt to analyse a combination of two modes was made by Pape (3) who presented an analytical expression for the pressure distribution in a seal having both waviness and coning. However, the separating force was calculated by numerical integration for a very specific sample case, thus limiting the generality of Pape\u2019s analysis. Stanghan-Batch and Iny (4) discussed the possible combination of waviness and angular misalignment and claimed that waviness is the main factor responsible for stable seal operation. StanghanBatch and Iny overlooked the hydrodynamic transverse moment shown by Etsion (5) to be inherent in misaligned seals. This transverse moment can start precession of the nutating flexibly-mounted seal ring which results i n appreciable load capacity even with flat faces. Another combination of two basic geometries was treated by Sharoni and Etsion (6). They presented closed form analytical solution for seals having both angular misalignment and coning and found that coning can drastically reduce the separating force. Coning is almost inevitable in mechanical face seals due to thermal distortion of the mating faces, as was shown by Kral(7) and also by Li (8). Hence, it is desirable to determine the effect of coning, combined with other geometries, on seal performance. The MS. of this paper was received at the Institution on 18th June i Technion-Israel Institute of Technology, Haifa. 0 References are @en in the Appendix. 1979 and accepted for publication on 9th January 1980. In this paper a combined coning and waviness model is treated where an arbitrary number of waves around the seal circumference is considered. Several investigators (e.g., Stanghan-Batch and Iny (4)) indicated only two sinusoidal waves around the seal circumference, but in some other cases (e.g., Kral(7)) more than two waves were observed. By considering an arbitrary number of waves the effect of number of waves on the separating force will be examined. In addition a comparison will be made with the separating force obtained from other geometries to evaluate their relative influence on the separating force. 1.1 Notation seal clearance along centreline wave amplitude separating force dimensionless separating force, F/6pwro2(ro/C)z dimensionless film thickness, h/C film thickness general integral defined in equation (16) number of waves pressure dimensionless radius, r/ro radial coordinate time function, e/[C + /3(r - ri)] angle of coning angle of tilt waviness parameter, e/C tilt parameter, yro/C coning parameter, /3ro/C angular coordinate viscosity transformed angular coordinate, nO shaft angular velocity Subscripts i inner radius m mid-radius 0 outer radius" ] }, { "image_filename": "designv11_20_0002052_iecon.2008.4758361-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002052_iecon.2008.4758361-Figure1-1.png", "caption": "Fig. 1: Two-dimensional fine positioning stage", "texts": [ " a modified Kalman Filter is used as disturbance observer to compensate the friction introduced by the ball bearing guides. After explaining the experimental setup, the employed control scheme is presented. Next the architecture of the disturbance observer is described. At the end it is shown by experimental results that the proposed control approach works properly over wide ranges and significantly improves the dynamical behavior of the controlled system. II. EXPERIMENTAL SET-UP The experimental set-up is a two dimensional fine positioning stage (see Fig. 1). It was constructed by members of the collaborative research centre SFB 622. As can be seen, every axis is driven by two ULIM3-2P-66 linear voice coil actuators of IDAM [8]. The motors are powered by proprietary developed analog amplifiers, which provide the needed current with the required precision. Commutation of the motors is achieved by the controlling system upon magnetic field intensity measurements provided online by Hall sensors. The operating range of this positioning stage is 200x200 mm2", " For data acquisition and control a modular dSpace\u00ae real-time hardware system in combination with Matlab/Simulink\u00ae is utilized. The position is provided by the SIOS interferometer unit as a 32-bit digital signal and is sampled by the dSpace\u00ae system at a rate of 25 kHz. The control algorithm uses a slower sampling rate of 6.25 kHz and operates on the analog amplifiers with a 16 bit resolution. For the presented study only the outer axis of the demonstrator is used. The inner axis is mechanically jammed at the position shown in Fig. 1. Neglecting the friction force based on the second Newton\u2019s axiom the dynamical behavior of the outer axis can be described as follows: ( )F m a t= \u22c5 (1) F is the applied force, m the mass and a(t) the resulting acceleration of the slider. Since the motor force has a linear relationship with respect to the applied current, the control algorithm controls the position via the current. The dynamical behavior of the amplifier can be neglected, because its cut-offrequency is higher than 10 kHz. Hence the force/current relation could be is simply modeled as a gain kA using the motor parameters provided by IDAM: ( ) ( )Ak i t m a t\u22c5 = \u22c5 (2) In state space notation the system can be expressed as: 0 ( ) 0 1 ( ) ( ) ( ) 0 0 ( ) = \u22c5 + \u22c5 \u23a1 \u23a4\u23a1 \u23a4 \u23a1 \u23a4 \u23a1 \u23a4 \u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 \u23a3 \u23a6 \u23a3 \u23a6 \u23a3 \u23a6 A x t x t i tk x t x t m (3) [ ] ( ) 1 0 ( ) = \u22c5 \u23a1 \u23a4 \u23a2 \u23a5\u23a3 \u23a6 x t x t y (4) where x(t) is the position and dx(t)/dt the velocity of the slider" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000930_robot.1985.1087262-Figure11-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000930_robot.1985.1087262-Figure11-1.png", "caption": "Figure 11. Joint Compliance Model", "texts": [ " where J is the Jacobian relating differential motion of the end effector frame E, aligned to the base frame B, and the joint frame q (See Fig. 10). EB q k Base Frame f The End Effector Applied Load During static positioning, when this additional load r is applied, the resulting motion has velocities and acceleration terms close to zero. As a result, the motions may be independently studied for each joint (dynamic effects can be ignored). The model of each joint and its compliant members is illustrated in Fig. 11. The equation of motion of the joint motor shaft fl, is given by: where T, is the motor torque, J, is the effective motor inertia, and B, is the effective motor viscous friction. The coulomb - static friction term, f,,,, is given as: The torque T, in equation (29) is the reaction torque exerted at the motor shaft due to the external load and the drive train. If the effective compliance of the gearbox is reflected to the joint drive shaft, then the motion of the final gear is given as: T2 = Jg 8g + Bg 8g + fegs + ks (8, - ear,) (31) where k, is the effective drive train joint compliance" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003828_s00453-010-9399-8-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003828_s00453-010-9399-8-Figure7-1.png", "caption": "Fig. 7 (a) The height of a book folding can be made arbitrarily small with pleat folds. (b) Alternatively the height can be reduced by adding new disks to the gap in the center to produce smaller molecules", "texts": [ " The ribbons, not considered part of any molecules, offset the shared axis of the deepest molecules so that it is \u03b5 above (and parallel to) the axis of the second-deepest molecules. We continue in this manner, building books within books, with each level of containment offset from the previous one by \u03b5, something like a layered wedding cake. See Chap. 17 of [16] for a similar construction. Now in order to cut all boundary components at once\u2014and nothing else\u2014we must fold the interiors of the molecules out of the way of the axes. To do this, we reduce the heights of the molecules so that they are all smaller than \u03b5, using the technique shown in Fig. 7 below. Now the ribbons determine the height of the entire folding (which will be about \u03b5 for a simple polygon, 2\u03b5 for a polygon with simple-polygon holes, and so forth). Finally, for the one-straight-cut problem, we can fold the entire book using folds parallel to the axes in order to bring all the axes to a single cutting line. This repaired construction also gives a (theoretical) way to perform an origami salami magic trick. In this trick, the first straight cut produces a hole in the shape of a silhouette of George Washington, the second straight cut (salami thickness \u03b5 away) produces a small John Adams, the third cut a still smaller Thomas Jefferson, and so forth, until we reach the tiniest president of all", " 6, we surround each hole with a thin \u201cribbon\u201d (offset polygon) of width \u03b5, where \u03b5 is smaller than polygon\u2019s minimum feature size. In the disk packing step of the construction, we pack the polygon minus the ribbons, and then project the tangency-point folds from the molecules perpendicularly across the ribbons. Thus the interior boundaries of the ribbons fold to the axis, and the boundaries of the holes fold to a parallel line, a new axis, distance \u03b5 from the original axis. By pleating down all the molecules (Fig. 7) to have height less than \u03b5, we can ensure that the interior of P embeds between the two parallel axes in the book folding. An alternative way to reduce a molecule\u2019s height is to add small disks to the 3- or 4-sided gap in its center. As shown in Fig. 7, we can add a disk to the center of a 3-sided gap, tangent to all three disks around the gap, and break that molecule into three smaller molecules with the same folds extending to the outside. This step can be repeated to reduce the height to less than any \u03b5. Proof of Theorem 1 Let M be a compact, orientable, genus-g, PL 2-manifold without boundary. If necessary, we triangulate the faces of M into small triangles in order to enable all subsequent steps. Using Lemma 3, we cut M\u2019s handles. Then as in the case of a topological sphere, we open one more path e to serve as the outer boundary, thereby obtaining a PL manifold M \u2032 homeomorphic to a disk with g pairs of holes, one pair for each handle loop", " 5(a)) cross a bridge, so that a tour around the perimeter of the tree of molecules (Fig. 5(b)) visits each hole and its paired hole in succession, and no pairs of holes cross (or even nest). In the book formed by all the molecules, Q and Q\u2032 thus appear within the same \u201cchapter\u201d, and R and R\u2032 appear together in some other chapter. The book flaps between Q and Q\u2032 are formed by the bridge of small molecules, chosen sufficiently small so that none of them stick out \u201cabove\u201d Q and Q\u2032. (If some molecule did stick out, we could of course pleat it back down as in Fig. 7.) Lemma 4 now folds M \u2032 so that cutting-forest edges all appear along one axis, and hole boundaries all appear along a second parallel axis. The construction is such that cutting-forest edges can be taped across the bottom of the book folding and holes can be taped across the top of the book folding, without any crossed pairs of tapings. A final taping closes the puncture path e. The ordering condition in the definition of a flat folding allows us to break connections at vertices and edges and displace faces into a third dimension, so that no two parallel faces interpenetrate in E 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003147_ijmee.38.2.5-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003147_ijmee.38.2.5-Figure4-1.png", "caption": "Fig. 4 Suggested beam cross-sections (top); assembled beam with bulkheads (middle); and loading schematic (bottom). Full dimensioned drawings are available at http://home. sandiego.edu/~dmalicky.", "texts": [ "comDownloaded from International Journal of Mechanical Engineering Education 38/2 students learned to use the following machines for this project: 52\u2033 stomp shear (Pexto), 6\u2033 corner notcher (Enco), 40\u2033 box/pan brake (Grizzly G0578), hand drill, and manual and pneumatic pop-rivet guns. The students were constrained to making the main beam cross-section from an 11\u2033 \u00d7 16\u2033 sheet of .032\u2033-thick 5052-H32 aluminum. Bulkheads were made from additional material. Students were encouraged to try different cross-sections: rectangular, U, I and triangle sections (Fig. 4). Most teams chose a 2\u2033 \u00d7 3\u2033 rectangular or I-section. To prevent collapse of the crosssections, bulkheads were specifi ed at each of the three load application points, though the detailed design of each bulkhead was up to the students. Beam construction generally took one lab period. All beams were tested in a three-point bending test fi xture of 15\u2033 span; students recorded the maximum load and mode of failure, and the test specimens were saved for later use in Mechanics of Materials. at USD & Wegner Health Science Information Center on April 9, 2015ijj" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001920_1.2988480-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001920_1.2988480-Figure4-1.png", "caption": "Fig. 4 Coordinate systems Om-xmymzm and O-xgygzg", "texts": [ " The positions of the gear axis and the datum plane must be etermined by measurement independent of the tooth surface easurement because the gear is set arbitrarily. We can make the rigin O and gear axis zg in the coordinate system O-xgygzg atached to the gear coincide with the origin Om and axis zm in the oordinate system Om-xmymzm of the CMM, respectively. Howver, the angle by which the gear is rotated about its axis is un- ertain. Therefore, we must define an unknown angle between ournal of Mechanical Design om: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/17/201 the xg and xm axes as shown in Fig. 4. Therefore, the tooth surface of the straight bevel gear and its unit normal are rewritten as xm and nm in Om-xmymzm as follows: xm u, ; = C xg u, + 0,0, T nm ; = C ng 6 where is the error of the apex to back l. Figure 5 shows that the gear tooth surface xm and a spherical probe of radius r0 of the CMM are in contact with each other at point Q. The coordinates of the probe center P are theoretically expressed as a position vector P Px , Py , Pz in Om-xmymzm as follows: P u, ; = xm u, ; + r0nm ; 7 On the other hand, the coordinates of the probe center P are measured by the CMM" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002187_15502280701759218-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002187_15502280701759218-Figure6-1.png", "caption": "FIG. 6. Bumper\u2019s CAD model.", "texts": [ "007 seconds and moved down, after reaching the maximum displacement, it returned back through its static position by a much smaller amplitude in another time span. After that the bumper experienced several small oscillations until it reached its static position. In computer simulation, a CAD bumper model was first created using IDEAS. All important geometries of the bumper system were measured and variant modeling techniques such as reflecting, lofting, sweeping, and shelling were applied to create the bumper model with a long curve surface. Figure 6 shows the bumper\u2019s CAD model. After finishing the CAD model, the bumper model\u2019s database was then transferred to an IGS file and imported into ANSYS environment to create a finite element model, which would be used for the computer analyses. In generating the finite element model, the bumper was meshed with the 2D shell element, SHELL181, which is a 4-node element with six degrees of freedom at each node and is suitable for analyzing thin to moderately thick structures. Two shocks were modeled using the 3D elastic beam element, BEAM188, which is a linear (2-node) or a quadratic 3D beam element that includes shear deformation effects" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000308_0094-114x(87)90058-9-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000308_0094-114x(87)90058-9-Figure7-1.png", "caption": "Fig. 7. The solution curve of compatibility equations; the curve of tan(/~/2) vs ~2. The solution points lying in between two straight lines which are drawn by the user are", "texts": [ " Those cases will lead to the synthesis of triads which have sliding pair(s) THE INTERACTIVE SOLUTION OF A FORTRAN program has been established for interactively choosing appropriate triads from the six sets of infinite number of solution triads. A description of this program is given below. Program input The input of the program contains the desired end point displacements, and the angular displacements of two links of a triad. Solution curves of the compatibility equations According to the input data, the program will calculate the values of tan(~3/2) with respect to each of the arbitrarily chosen values of/~2 which may range from 0 to 360 \u00b0 and display a curve which contains all values of tan(/~ 3/2) with respect to/~2 as shown in Fig. 7. i ::' APPLICATION EXAMPLE An example is given in this section to demonstrate the application of the triad synthesis program to the synthesis of an eight-bar linkage. This example shows the dimensional synthesis of the type of an eight-bar two-loop linkage as given in Fig. 10, so that the coupler point p of the eight-bar linkage will pass through six prescribed motion positions pj for j -- 1-6 at the specified angular displacements of ground links 1-2, 4-5, and 7-8 which may be geared together to form a one-degree-of-freedom mechanism" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002767_j.polymer.2010.05.003-Figure12-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002767_j.polymer.2010.05.003-Figure12-1.png", "caption": "Fig. 12. Scheme of the rolling mechanism of PANI nanotubes. a) oligoaniline nanosheet, b) oligoaniline nanosheet and PANI nanosheet of emeraldine base form, c) semirolled PANI nanotube, and d) PANI nanotube.", "texts": [ " In addition, the tubular morphology of the pure PANI is not destroyed by removal of the oligoaniline through methanol washing, as shown in Fig. 1. Based on the above results, the final products are simple mixtures of pure PANI nanotubes and oligoaniline complexes composed of oligoanilines and aniline sulfate salts. Thus, the oligoaniline complexes produced in the early stages do not grow into the PANI chains, but function as only templates for 1D PANI formation. In addition, KSBAwas introduced only to regulate the pH of reaction medium and to control oligoaniline morphology. Fig. 12 illustrates the scheme of the rolling mechanism of PANI nanotubes synthesized in the presence of the KSBA: a) first, after adding APS, the oligoaniline nanosheets composed of aniline sulfate salts and oligomers of phenazine-like units are synthesized; b) oligoaniline nanosheets act like templates for the growth of PANI nanosheets of emeraldine base form; c) PANI nanosheets start to be rolled up by the conformational change due to the protonation of PANI chains under acidic condition; d) finally, PANI nanotubes are fabricated as a result of the rolling process" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002816_es2009-90480-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002816_es2009-90480-Figure2-1.png", "caption": "Figure 2 Schematic of a wind turbine", "texts": [ " The wireless telemetry is well proven and is easily demonstrated in a laboratory. SECTION 1: Operational Evaluation Investment in a structural health monitoring system for wind turbine blades requires economic justification for its purchase, installation and maintenance. Additionally, damage must be defined, normal operating environments must be identified and any limitations on data acquisition should be noted. The economics of a three blade horizontal axis wind turbine is considered (schematic shown in Fig. 2). The three blade horizontal design constitutes most \u201cutility-scale\u201d turbines on the global market. Wind turbines work by converting wind energy into electrical energy. The wind spins the blades of the turbine which turns a drive-train that turns a generator. Utility lines collect the electricity generated and then distribute it to consumers. Wind turbine profitability depends on the amount of electricity produced, which in turn depends on both the size of the turbine and the wind speed. Utility scale rotor sizes range from 50 to 90 meters" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000098_b:frac.0000021022.48417.a6-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000098_b:frac.0000021022.48417.a6-Figure2-1.png", "caption": "Figure 2. The general view of a subsurface crack with several cavities and closed segments.", "texts": [ " These conditions depend on whether a crack is of a surface or subsurface nature, and whether its faces are in contact with each other or not. First, let us consider a subsurface crack. Obviously, at the subsurface crack tips the jumps of the normal and tangential displacements of its faces are equal to zero v(\u00b1l) = u(\u00b1l) = 0. (4) At the open segments of the subsurface crack the normal stress is zero, and it is non-positive at the segments of the crack which are in contact with each other (see Figure 2) pn(x) = 0, v(x) > 0; pn(x) \u2264 0, v(x) = 0 for y0 < \u2212l sin |\u03b1|. (5) Now, let us consider the case of a surface crack which mouth (x = l) emerges at the halfplane boundary at the point with the global x0-coordinate equal to x0 + l cos(\u03b1)sign(\u03b1). At the subsurface tip of the surface crack the jump of the tangential displacements of its faces is equal to zero. At the mouth of such a surface crack the derivative of this jump along the crack is finite. Thus, we have u(\u2212l) = 0, du(l) dx \u2212 finite. (6) The more accurate statement is: the shear stress intensity factor k+ 2 is zero (see (15))" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001744_j.optlastec.2006.12.009-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001744_j.optlastec.2006.12.009-Figure3-1.png", "caption": "Fig. 3. Schematic of the relationship between the coordinates of the points at the cross-sectional contour line and those at the common boundary of the three phases (gaseous, solid and liquid phase).", "texts": [ " If edge B0 is expressed as function z \u00bc f(x,y), every edge contributing to formation of the cross-sectional profile can be expressed as function z \u00bc f(x+s,y), where s is defined as a value within the range 0\u2013DS, where DS is distance in x direction from an intersection point of edge B0 with plane x\u2013y to the corresponding one of edge Bn with plane x\u2013y. Therefore, all points at the above-mentioned crosssectional profile can be defined by z \u00bc f \u00f0x\u00fe s; y\u00de, x \u00bc 0. \u00f01\u00de For any cross-sectional profile, x \u00bc X, the points at it can be expressed as z \u00bc f \u00f0x\u00fe s; y\u00de, x \u00bc X . \u00f02\u00de Fig. 3 shows the link between the coordinates of the points at the cross-sectional profile of the clad bead and those of the corresponding points at the existing pool edge in steady-state laser cladding. On the surface of the clad bead, the lines aa0, bb0, cc0, dd 0, ee0, ff 0 and gg0 are located on the longitudinal profiles of the clad bead and parallel to each other. The points on the same line have the same y-coordinates. In steady-state laser cladding, the pool shape and size remain stable, the shape of the interface of solidification remains constant, thus the pool edge, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002994_acc.2010.5530875-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002994_acc.2010.5530875-Figure1-1.png", "caption": "Fig. 1. Definitions of axis, control surfaces, and main variables", "texts": [ " The propulsion system consists of a RIMFIRE 22M-1000 brushless motor with a 10x3.8 propeller coupled to a 25A electronic speed controller. The aerodynamic surfaces are controlled using a standard 72 MHz receiver and micro servos. The vehicle attitude is measured by a wireless IMU (Microstrain 3DM-GX2). It has 3 accelerometers, 3 rate-gyros, and 3 flux-magnetometers. The IMU calculates and transmits orientation via a cosine matrix at 100 Hz. A vision system is used to measure the altitude in hovering mode. As shown in Fig. 1, there are four manipulated variables: ailerons (\u03b4a), elevator (\u03b4e), rudder (\u03b4r) and throttle (\u03b4t). XB , YB , and ZB are the body axis with the origin at the center of gravity (CG) of the aircraft. \u03c6, \u03b8, and \u03c8 are the body frame angular attitude; p, q, and r are the body frame angular rates; ax, ay , and az are the body frame accelerations; u, v, and w are the velocity components of CG in the body reference frame. For navigation, the NED axis system (North-EastDown) is used as the frame of reference on the Earth" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003147_ijmee.38.2.5-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003147_ijmee.38.2.5-Figure10-1.png", "caption": "Fig. 10 Sample student-machined compressed-air engine.", "texts": [ " Students learned to use the following machine tools for this project: 12\u2033 \u00d7 36\u2033 engine lathe with digital read-out (DRO) (Birmingham, Sony), manual milling machine with DRO (Lagun, Sony), horizontal bandsaw (jet), tool grinder, drill press, dial caliper, and vernier micrometer. Before starting machining operations, all students developed operation sheets for each part, which were reviewed by the instructors. On the lathe, almost all students successfully achieved the +/\u22120.0005\u2033 tolerance for their piston diameter. Additional practiced skills included print-reading, use of machinist tables, press-fi tting, tapping, assembly, and shop professionalism. A sample student engine is shown in Fig. 10. The importance of holding tolerances becomes clear to many students during the assembly phase. For example, holes intended for a press-fi t often required rework due to a reaming operation that cut oversize. Or, the smoothness of running was at USD & Wegner Health Science Information Center on April 9, 2015ijj.sagepub.comDownloaded from International Journal of Mechanical Engineering Education 38/2 at USD & Wegner Health Science Information Center on April 9, 2015ijj.sagepub.comDownloaded from International Journal of Mechanical Engineering Education 38/2 occasionally hampered by a non-perpendicular cylinder bore" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.53-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.53-1.png", "caption": "Fig. 14.53 Multipoint pneumatic vibrating beam", "texts": [ " Vibrating beams equipped with attachable electric vibrators are often adapted to a safe voltage supply to eliminate the electric shock hazard. Vibrating beams of the type shown in Fig. 14.52 are manufactured 2.7\u20135.7 m long and can compact 10\u201330 cm thick bases. The development of vibrating beams is directed towards increasing their length, obtaining uniform compaction of concrete mix along the whole length of the vibrating beam, and improving operational safety and transportability on the construction site. To a large extent the above requirements are met by the multipoint pneumatic vibrating beam shown in Fig. 14.53. The vibrating beam is 6.1 m long and consists of two end segments and one middle segment, joined together by bolts. The individual segments have a lattice structure triangular in cross section. Two stringers, one in the form of an angle and the other a T-bar, form the vibrating beam\u2019s base. The upper stringer is made from a pipe, which also serves as the compressed air conduit that supplies the vibrators. The vibrating beam is equipped with 16 pneumatic vibrators spaced at different intervals on both the lower stringers" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002565_j.tws.2008.08.010-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002565_j.tws.2008.08.010-Figure6-1.png", "caption": "Fig. 6. Cushion disc CAD model.", "texts": [ " Furthermore, the elastic springback which occurs after the forming process of the cushion disc affects the geometry accuracy. Besides, the cushion disc geometry is difficult to control with conventional metrological tools because it bends under the load applied by any mechanical probe. All this influences the accuracy and repeatability of experimental correlation with FE models. A FE model was implemented by applying ANSYS Parametric Design Language (APDL). Only two paddles of the cushion disc are modelled taking into account the symmetry (Fig. 6). The cushion disc (Fig. 7) is meshed with shell elements (Shell 181 in the ANSYS library) according to its thin thickness (0.7 mm) and computed using a linear elastic material law. The pressure plates are modelled by two flat rigid plates that compress the disc axially to simulate its behavior during the gear re-engagement. A master\u2013slave surface to surface contact pair is used between the cushion disc (master surface) and each pressure plate (slave surfaces). The contact elements associated to master and slave surfaces are, respectively, Contact 173 and Target 170 in the ANSYS library" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002349_s12239-009-0039-8-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002349_s12239-009-0039-8-Figure5-1.png", "caption": "Figure 5. Drivetrain layout.", "texts": [ " However, separate lubrication is needed if the ISCVT is installed because traction fluid is not designed to lubricate the engine. The ISCVT prototype was installed using several existing parts of the motorcycle: the engine, the clutch system, the primary reduction gear, the final reduction chainsprocket, and the fifth gear of the stepped transmission. Therefore, the primary reduction and the fifth gear of the stepped transmission of the existing motorcycle play the role of primary reduction of the ISCVT prototype, which has a reduction ratio of 0.3. The schematic of the drivetrain layout is depicted in Figure 5. The main shaft of the ISCVT prototype is concentrically a threefold structure; i.e., the outer shaft contains the inner shaft with bearing support, and the inner shaft contains an additional inner shaft with additional bearing support. In this structure, the engine power input occurs through the \u03c1 = h3 h2 ---- = f \u03c6( ) shaft, and the ISCVT output flows from the outer shaft. The final reduction ratio by the chain-sprocket is 0.32. The specifications for the motorcycle and the design concept are listed in Table 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003663_oxfordjournals.jbchem.a128592-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003663_oxfordjournals.jbchem.a128592-Figure3-1.png", "caption": "FIG. 3. Succinatc oxidase system in \"shockate\", membrane and supernatant.", "texts": [ "005 M, the presence of which resulted in the retention of approximately 50% of protein, 40% of RNA and 30% of DNA in the membrane system. Characterization of Cellular Components in the Membrane System i. Enzymes\u2014The membrane system prepared in the presence of 0.005 M MgCl2 contained one half the total spheroplast protein as indicated in Table I. Glucose-6-phosphate dehydrogenase was, however, detected mostly in the supernatant fraction as shown in Fig. 2. On the contrary, succinate oxidase system was found to be specifically localized in the membrane and almost absent in the supernatant fraction as demonstrated in Fig. 3. This result is in good agreement with the finding by S e d a r and B u r d e who used the combined techniques of cytochemistry and electron microscopy (11). TABLE I Effects of Magnesium Ions on the Distribution of Cellular Components between Membrane System and Supernatant Fraction Mg++ concentration during preparation of membrane system (mM) O.D.660m, RNA ( / \u00ab \u2022 ) DNA (/*8-) Protein (mg.) Membrane Supernatant Sum Membrane Supernatant\" Sum Membrane Supernatant Sum 0 0.730 247 993 1,240 34. 1 144 178 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000975_6.2004-5308-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000975_6.2004-5308-Figure4-1.png", "caption": "Figure 4. Generalized coordinates of the system: \u03b1i, \u03b2i and li", "texts": [ " For formations spinning in a plane normal to the orbit, those coordinates are not convenient, because when \u03b8 is not zero, a change in the \u03c6 angle does not represent angular motion in the y-z plane. It is therefore not possible to examine the spin rate using these coordinates when the spin plane is not the orbital plane. For that reason, the generalized coordinates are chosen to be \u03b1i, \u03b2i and li, where \u03b1i is the angle between the tether and the z-axis when the tether is projected onto the y-z plane, \u03b2i is the angle between the tether and its projection in the y-z plane, and li is the distance between the central body and a peripheral satellite (Fig. 4). The figure shows the ith tether between the parent body P and a peripheral satellite i. The projections of the bodies onto the y-z plane are shown on the figure as P \u2019 and i\u2019. The form of the equations obtained are the same as those for the orbital plane. Lagrange\u2019s equations can once again be written in the following form: d dt ( \u2202T \u2202q\u0307j ) \u2212 \u2202T \u2202qj + \u2202Vg \u2202qj + \u2202Vsp,in \u2202qj + \u2202Vsp,ex \u2202qj = Qd,in,qj + Qd,ex,qj (25) where qj = \u03b1j , \u03b2j , lj , j = 1, . . . , N . The following equations of motion are obtained: N\u2211 k=1 lj(GjkS\u03b1jk + RjkM\u03b1jk) = N\u2211 k=1 lj\u03a0jkM\u03b1jk (26) N\u2211 k=1 lj(GjkS\u03b2jk + RjkM\u03b2jk) = N\u2211 k=1 lj\u03a0jkM\u03b2jk (27) N\u2211 k=1 (GjkSljk + RjkMljk) + kin m \u00b7 (lj \u2212 l0,j) = \u2212cin m \u00b7 l\u0307j + N\u2211 k=1 \u03a0jkMljk (28) The components of the position and velocity vectors, as well as the various terms in the equations of motion above, are detailed in Appendix B" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000462_978-1-4020-2249-4_38-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000462_978-1-4020-2249-4_38-Figure8-1.png", "caption": "Figure 8 Coronal Articulations the Dodecahedral Zig-Zag Linkage", "texts": [ " The Octahedral Zig-Zag Linkage built up with Flat Ring Articulations is certainly mobile with one degree of freedom, but it could be that it is not overconstrained and is movable with more then one degree. Therefore we determine F with 1 = 13 (disregarding the loops in the Nuremberg Scissors), j = 6 x 8 (disre garding the joints in the Nuremberg Scissors) and I./; = j + 12 = 60 and find: F = I./; - 61 = 60 - 6 x 13 = -18, which proves overconstraintness. The forth kind of articulation of the Nuremberg Scissors will be called the \"Coronal Articulation\" and demonstrated on the dodecahedron. Fig.8 shows that the three Nuremberg scissors, positioned \"upright\" on the do decahedral edges, are connected by three gussets with nonintersecting rotary joint axes whose orientation now depends on the angle cp. With the set of parameters {a,s,d,t,f3=20.91\u00b0},h(cp)=acoscp-ssincp and the components 1\\ = kcosf3 + hsinf3 + (kcosf3 - hsinf3)/ 2,12 = (kcosf3 - hSinf3)..J3 /2 and 13 = 2hcosf3 of the gusset vector t. the distance k(cp) can be found from the condition I~ + t~ + Iff = 12 and therewith the variable length of the dodecahedral edges is given by: L( cp) = 2~-5h( cp)2 + 212 - 3h( cpi cos2f3 /(cosf3..J6)+ 6(a sincp + scoscp). For the mechanism in Fig. 8 we find F = 6 + 3- 6 x 1 = 3 and for the Do decahedral Zig-Zag Linkage (Fig.9): F = I./; -61 = (20x6+30)-6x31 = -46. For thkosahedral Zig-Zag Linkage we use the Flat Ring Articula tion. With the {a,s,d,t,{3 = arccos[(-3+,j5)/~(20-8,j5) -n/2 - 31, 71\u00b0}and h(cp) = acoscp-ssincp the edge length ofthe icosahedron can be calculated: L( cp) \"\" 2[(1 I 2) + h( cp )cos( n I 5)]/[ cos{3 sine n I 5)] + 6(a sin cp + s cos cp). The mobility of the mechanism shown in Fig.10 (Flat Ring Articulation and 5 Nuremberg scissors) is found by F = ~/; - 6/ " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003369_j.mechmachtheory.2010.12.002-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003369_j.mechmachtheory.2010.12.002-Figure4-1.png", "caption": "Fig. 4. Example of a mechanism with eight assembly configurations (the arrows show the points of contact between surfaces).", "texts": [ " The solution chosen for the first position should be continued for following positions to avoid a jump to another assembly configuration during the calculations of successive positions. The term \u201cassembly configuration\u201d means a possible mechanism configuration where the input and output members can move (at least in a certain proximity to the assumed position of the input member), and at least one possible direction of movement exists. The dependence \u03b22(\u03b21) is the position function of RTTR mechanism. The minimum distance between the surfaces, e.g. \u03b22, may change from one to eight times (Fig. 4). This is also the number of the assembly configuration of this mechanism. There are situations where two assembly configurations have common positions. Such situations are widely known in the theory ofmechanisms. They occur even in the simplestmechanisms, e.g. a flat four-bar linkage. These are points of bifurcation [14], and a greater accuracy of calculations is required to obtain correct solutions. When analysing the three-link mechanism of RTTR with a higher kinematic pair in the form of two contacting surfaces in the shape of tori joined with the frame by means of revolute pairs, the definitions of singular points may be divided into four types", " Assuming the pairs joining the elements with the frame can be any pairs (R, P or H) and any element can be a frame, hundreds of known structures of single-race, single-loop mechanisms may be created. In the case of some mechanisms, additional mobility can occur. This happens, inter alia, if more than 3 prismatic pairs (or cylindrical and prismatic pairs) exist in the structure. All the obtained structures are special cases of the presented mechanism model. Table 4 includes examples of mechanisms. As an example of kinematic synthesis of a mechanism with a higher pair in the form of two tori, a RTTR mechanism was considered (Fig. 4). The following guidelines have been accepted (according to the denotations in Figs. 2 and 3): \u2022 the input link performs a full rotation 0\u2264\u03b21\u2264360\u00b0; \u2022 the output link performs a rotation (the angle between the axes: \u03b12=0\u00b0 and distance a1=10) with constant transmission ratio k=\u03c92/\u03c91=0.1 in the range of the driving link rotation 0\u2264\u03b21\u2264180\u00b0, where\u03c91=\u2202\u03b21/\u2202 t and \u03c92=\u2202\u03b22/\u2202 t; \u2022 pressure angle \u03bc1\u226460\u00b0. The driving axle overlaps with the OZ-axis of fixed system when \u03b11=0 and a1=0 (Fig. 3). The index of the mechanism transmission ratio was accepted as the optimization criterion K = \u2211 180 i=1 ki\u22120:1\u00f0 \u00de2 calculated on the basis of the value of themechanism transmission ratio ki" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001126_j.aca.2006.09.037-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001126_j.aca.2006.09.037-Figure3-1.png", "caption": "Fig. 3. Schematic diagram of an electrochemical flow-cell made up by the glass substrate with sputtered silver alloy and platinum and other parts: a,", "texts": [ " Howver, the cyanide served as a reactant raises problem of environ- ental pollution and may influence the enzyme activity of FAO, PO, or protease. Another is sodium lauryl sulfate-hemoglobin SLS-Hb) method [44,45], which is cyanide free and has a good orrelation with the HiCN method. himi c s t t F t H d H c ( 2 2 1 o f ( f d D a E ( i ( h d w c F t c f w ( f B m h r M U b a ( 2 e t t a U f l 2 U t a p l b s w i b c 2 g ( a c with ethanol, and then a PTFE adhesive tape (d) with a square hole (13.5 mm \u00d7 13.5 mm) was set up as shown in Fig. 3. Then, a 40 L portion of 1,4-dioxane solution containing 1% (w/v) acetyl cellulose and 0.01% (w/v) vinyl acetate was spread over Y. Nanjo et al. / Analytica C This paper describes a FIA system comprised of an electrohemical detector with a FAO or FPO reactor and a flow-type pectrophotometer, which makes it possible to determine simulaneously fructosyl peptide (or/and fructosyl amino acid) and otal hemoglobin. By the combination of protease which releases VH but not FV from the HbA1c, this FIA method was applied o the simultaneous assay of FVH and total hemoglobin in the bA1c blood samples with different HbA1c values", " The preparation was washed with distilled water and then hosphate buffer (0.1 M, pH 7.0). The FAOX-TE (13 U) was oaded into the glutaraldehyde-activated Uniport C, by incuating in enzyme-sodium phosphate buffer (0.1 M, pH 7.0) olution overnight at 4 \u25e6C. The immobilized enzyme support as packed into a 30 mm long column (OSI, 3 mm i.d.). Simlarly, FPOX-CET (23 U) was also loaded onto aminoalkylonded Uniport C. Likewise, it was packed in a 30 mm long olumn. .3. Preparation of hydrogen peroxide electrode In the OSI electrochemical flow-cell shown in Fig. 3, the lass substrate (e) with sputtered silver alloy (e1) and platinum e2 and e3) was covered with acetyl cellulose membrane. It cted as a sensing part of the hydrogen peroxide in the flowell. The surface of the glass substrate was initially washed lectrical connector; b, upper cell part; c, rubber ring; d, PTFE adhesive tape 17.5 mm \u00d7 17.5 mm) with the square hole (13.5 mm \u00d7 13.5 mm); e, the glass ubstrate (17.78 mm \u00d7 22.86 mm); e1, sputtered silver alloy as a reference elecrode; e2, sputtered platinum as a working electrode; e3, sputtered platinum as counter electrode; f, base cell part" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000978_2004-01-2911-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000978_2004-01-2911-Figure2-1.png", "caption": "Figure 2 - Piston geometry: definition of te and be", "texts": [ " Figure 1 shows the forces and moments acting on the piston. These forces and moments come from interactions of the piston with the liner, rings, wrist-pin, connecting rod, cylinder pressure, and inertia. The nomenclature section as well as References [1,2,22] explain in detail the definition of parameters and the forces and moments on the piston. The primary motion - the piston position, speed and acceleration along the axis of the cylinder - can be calculated as a function of crank speed [1]. Figure 2 shows the piston geometry and definition of piston transverse movements, te and be , i.e. the lateral displacements of the piston from the cylinder axis in the direction of the crank axis at the top and bottom of the skirt respectively. Therefore, the balance of forces and moments about the wrist- pin yield the set of governing differential equations of piston secondary motion. The transverse inertial forces and moments depend on the piston transverse acceleration, te&& and be&& , as shown in equation (1) which is explained in detail in Ref" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001120_05698197608982807-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001120_05698197608982807-Figure7-1.png", "caption": "Fig. 7-Measured and corrected pressure curves for Fluid C.", "texts": [ " T h e pressure taps have holes (I,, = 0.3 mm in diameter compared to h,,, = 0.5 nim, the ~ n i n i m u m film thickness. So the real pressure d r o p d u e to the \"hole effect\" shoulcl be expected between tlie two limits of the magnitudes obtained from the average a n d wall shear rates. By superposition of the measured pressure distribution ancl thc pressure d r o p d u e to the \"hole effect,\" the two limits of the real pressure distl-ibution a re obtained, and s l i o ~ \\ ~ n by the vertical lines in Fig. 7. From this figure, by separating into an antisymmetrical pressure a n d the symmetrical part, one gets in Figs. 8a and Bearing Angle 8 (degrees) Flg. &Wail and average ahear rates In the bearing's fluid film. n c I : 40- t? l C I' i r W ' ~ ~ ~ ~ ~ RATE AT THE BEARING WAU i 3 , a i 0 ! I a z YI 0 3 0 6 0 9 0 120 150 180 :\\o(; A. HAKNOV A N D 81) the Nc\\vtonian anel tlie two limits of tlie elastico-viscous p;1rts or t he pl-essurc distribution. The data for the pressure c ~ l r \\ ~ c s of' tlic clastico-viscous fli~icls is as follows" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002261_j.mechmachtheory.2009.09.006-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002261_j.mechmachtheory.2009.09.006-Figure3-1.png", "caption": "Fig. 3. Scheme of the two different calculations.", "texts": [], "surrounding_texts": [ "In this paragraph, we propose creating a \u2018\u2018pseudo\u201d base of functions that can be used in the case of spiral bevel gears. The term \u2018\u2018pseudo\u201d base of functions is used because the K independent functions are generative only in a K dimension base. The efforts are applied on the surface of the teeth. Similarly, the bending displacements are calculated on the same surface, it is therefore possible to use functions with only two parameters: L10 and L20, which are the point locations along the tooth width and along the tooth height respectively (Fig. 4). It was decided to separate the two parameters of the functions. Each function is a product of two functions having only one parameter. fk\u00f0L10; L20\u00de \u00bc hj\u00f0L10\u00de gi\u00f0L20\u00de \u00f022\u00de With J functions according to L10: hj(L10), and I functions for L20: gi(L20), it is possible to obtain a base of K = I J functions. In order to obtain a well-conditioned matrix, both kinds of functions are reduced to a standard space on the tooth surface. Thus the functions fk are written in the following form: fk\u00f0L10; L20\u00de \u00bc Hj\u00f0v\u00de Gi\u00f0u\u00de \u00f023\u00de with u \u00bc L20 L20f \u00f0L10\u00de L20a\u00f0L10\u00de L20f \u00f0L10\u00de and v \u00bc L10 L10min b If L20f(L10) corresponds to the root tooth point for plane L10, if L20a(L10) corresponds to the top tooth point for plane L10, if L10min is the minimum of plane L10 corresponding to the tooth edge, and if b is the tooth width, the entire surface of the tooth is covered when parameter u varies from 0 to 1 and when parameter v also varies from 0 to 1. Parameters L20f(L10), L20a(L10) and L10min are presented in Fig. 5. It was decided to use polynomial functions for the G functions, because the behaviour of the tooth along its height is similar to the behaviour of a fixed-free beam (Eq. (24)). Gq\u00f0u\u00de \u00bc uq 1 \u00f024\u00de Along the tooth width, its behaviour is close to the behaviour of a free-free beam. That is why we have chosen to use the resonance functions of such a beam for the H functions as in [9] (Eq. (25)). if q \u00bc 1 Hq\u00f0v\u00de \u00bc 1 if q \u00bc 2 Hq\u00f0v\u00de \u00bc 1 2u if q > 3 Hq\u00f0v\u00de \u00bc sin\u00f0lqv\u00de \u00fe sinh\u00f0lqv\u00de gq\u00bdcos\u00f0lqv\u00de \u00fe cosh\u00f0lqv\u00de with gq \u00bc sin\u00f0lq\u00de sinh\u00f0lq\u00de cos\u00f0lq\u00de cosh\u00f0lq\u00de lq \u00bc q 3 2 p 8< : \u00f025\u00de" ] }, { "image_filename": "designv11_20_0000457_bfb0075007-Figure11-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000457_bfb0075007-Figure11-1.png", "caption": "Figure 11", "texts": [ " 4\u2022 \u2022 Trivial knot orthogonal to the plane figure 8 38 - -'t-r- ';:::.l:>----+=-et===t:+===I- ?-1- -1 ;==========t==t=-i---tli-_a... ?C=========i==I=t+==1- Ithe band c +2Number of turns of =========- ---= =====t=t====-JE. It- \" - -13l't-:=================_/6 figure 9 39 J \"f ... r> r: h \":, - _ r ;::. :> f - q - 1(; .... . 11 Ir 16 1[' /6 /0 Number of turns of the band = +4 figure 10 40 f(-- '/- J- fr -( \"\"\"\"\"'I r:. - - b I I\"\" , _I.> - 0 denotes the viscous friction coefficient and \u03c1c > 0 denotes the Coulomb friction level. Suppose that the uncertainty term \u03b4(t, q, q\u0307) is bounded by growing terms as in (27)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001657_bfb0119410-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001657_bfb0119410-Figure1-1.png", "caption": "Figure 1: The Field and Space Robotics Laboratory (FSRL) Rover (without cover)", "texts": [ " Future planetary explorers will need to navigate rugged terrain and travel substantial distances, while exercising a high degree of autonomy [2,3]. The focus of our research at the MIT Mechanical Engineering Field and Space Robotics Laboratory (FSRL) is on developing methods and algorithms for the design, planning and control of high-performance robotic planetary explorers based on the physics of these systems. To validate the effectiveness of the developed algorithms, experimental evaluation is essential. To perform this evaluation a low-cost rover test-bed has been developed (see Figure 1). The FSRL rover is a 6-wheeled rocker-bogie vehicle that is similar kinematically to the JPL Lightweight Survivable Rover (LSR) (see Figure of rock samples. The experimental system chassis contains shape memory alloy (SMA) actuated variable geometry mechanisms that re-configure the system to improve its ability to traverse difficult terrain. All power is provided by on-board batteries. Low-level control and planning are performed using an on-board PC/104 computing architecture. A wireless modem is used for external communication to obtain high-level commands from a task planner" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000608_155022891010015-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000608_155022891010015-Figure3-1.png", "caption": "FIG. 3(b). Wave of the race.", "texts": [ " Waviness is realized in the form of peaks and valleys of varying height and width. Therefore, for mathematical modeling using waviness effect, a statistical approach is necessary in order to have a complete solution. If the races are assumed to bend due to rolling element loads then the flexural vibrations of the races as well as the rigid body motion have to be considered. To avoid these problems the races are assumed bend-less under these loads and a sinusoidal wavy surface is assumed as shown in Fig. 3(a). The wavelength is assumed to be much longer than the roller to race foot print width and the wave geometry itself is assumed to be unaffected by contact distortion. Waves are described in terms of two parameters: the wavelength (\u03bb), which is the distance taken by a single cycle of the wave and its amplitude ( i ). The amplitude of wavy surface is often measured with respect to central point at a certain angle from the reference axis. Hence the amplitude of sinusoidal wave is: ( )i = p sin ( 2\u03c0 L \u03bb ) (2) The race has circumference sinusoidal wavy surface, therefore the radial clearance consist of a constant part and a variable part. Hence the amplitude of the wave of race is: ( )i = ( O ) + ( p) sin ( 2\u03c0 L \u03bb ) (3) Where p is the maximum amplitude of wave and o is initial wave amplitude (or constant clearance) as shown in Fig. 3 (b). The arc length (L) of the wave at the contact angle is: L = r\u03b8 j (4) The wavelength is the ratio of length of the race circumference to the number of waves on circumference, which is: \u03bb = 2r\u03c0 Nw (5) The amplitude of the race waves at the contact angle is: ( )i = ( O ) + ( p) sin(Nw\u03b8 j ) (6) Hence the contact angle is: \u03b8 j = 2\u03c0 Nb ( j \u2212 1) + \u03c9cage \u00d7 t (7) The cage speed \u03c9cageis: \u03c9cage = 1 2 \u03c9inner [ 1 \u2212 \u03c1 j Rp ] + 1 2 \u03c9outer [ 1 + \u03c1 j Rp ] (8) The varying compliance frequency is: \u03c9vc = Nb\u03c9cage (9) Hence the instantaneous amplitude of waviness at the contact angle is: ( )i = ( O ) + ( p) sin [ Nw {2\u03c0 Nb ( j \u2212 1) + \u03c9cage \u00d7 t }] (10) Hertz considered the stress and deformation in the perfectly smooth, ellipsoidal, contacting elastic solids" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003444_j.proeng.2010.04.132-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003444_j.proeng.2010.04.132-Figure1-1.png", "caption": "Figure 1: Double-support and single-support phases", "texts": [ " Keywords: Hammer throw, Hula-hoop, Energy pumping, Hetero-parametric excitation The motions of a hammer are commonly regarded as a circular movement which a hammer undergoes around a turning human body. The distance of a throw depends on the speed of the hammer and the angle of its trajectory at the instant of release. The speed of the hammer increases gradually during preliminary winds and turns. The hammer throw is a sequence 1 of turns that involves moving the hammer while turning on either one leg (single-support phase) and turning on both legs (double-support phase)(Fig. 1). During each turn the hammer rises up toward the high point of the trajectory and passes through the low point. There are many studies which describe such hammer and thrower movement with experimental analysis [1, 2, 3, 4]. However, theoretical analysis which describe the mechanisms of acceleration are few. Our interest in this study was to understand the mechanism of acceleration of the hammer in terms of energy pumping. The motions of hammer were analyzed and numerical experiments were performed to examine our theory" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000936_978-1-4020-4941-5_24-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000936_978-1-4020-4941-5_24-Figure4-1.png", "caption": "Figure 4. Cardanic self motion when the legs are concurrent and make equal angles.", "texts": [ ", if L1 and L2 are not parallel, then the elliptic curve degenerates if, and only if, its curvature is zero for any \u03c6. The curvature, \u03ba, of E can be derived as \u03ba = x\u0307y\u0308 \u2212 y\u0307x\u0308 (x\u03072 + y\u03072)3/2 = sin(\u03b81 \u2212 \u03b82 + \u03c0/3) D3 , (11) where D is a relatively large expression in \u03b81, \u03b82, and \u03c6. Therefore, when L1 and L2 make a 60\u25e6 angle, \u03ba = 0, and E degenerates to a line. In other words, if, and only if, the point of intersection between L1 and L2, denoted by P12, lies on the curcumcircle of the base, denoted by C, E degenerates to a line (Fig. 4). In fact, it degenerates to a doubly traced line segment of length 4/ \u221a 3 (Tischler et al., 1998). This line segment is centered at P12 and passes through O. When P12 \u2261 O, the doubly traced line segment is parallel to O1O2. As we said before, the direct kinematic problem is equivalent to finding the two intersection points between L3 and E , of which one is always O. We will not present an actual algorithm for determining the other intersection point (x, y) and the corresponding platform orientation \u03c6, but only investigate the singular configurations corresponding to all particular cases in which there is a single or infinitely many solutions", " Case 1b: cos \u03b81 = 0 and cos \u03b82 = 0 In this case, L1 and L2 are parallel to the y-axis and E degenerates to a single line parallel to L1 and L2, and passing through O. If L3 is parallel to L1 and L2, the platform vertices can slide along L1, L2 and L3, with \u03c6 = 0, even though all actuators are blocked. If L3 is not parallel to L1 and L2, the platform can assume only the trivial solution q = 0, and the configuration is Type 1 and Type 2 singular (the platform can rotate infinitesimally). Case 2: sin(\u03b81 \u2212 \u03b82 + \u03c0/3) = 0 In this case, L1 and L2 make a 60\u25e6 angle and their intersection point, P12, lies on C (Fig. 4). The curve E degenerates to a doubly traced line segment passing through P12 and O (if P12 \u2261 O, E is parallel to O1O2). If L3 is collinear with E , then point B3 can slide along E while the platform changes orientation simultaneously (Fig. 4), even though all actuators are blocked (as in Reuleaux straight-line mechanism). If L3 is not collinear with E , then the platform can assume two possible poses. The first one is the trivial solution q = 0, while the second one is q = [180\u25e6, 0, 0]T . For both poses, the corresponding configurations are only Type 1 singular (the platform is jammed). Case 3: sin(\u03b81 \u2212 \u03b82 + \u03c0/3) sin(\u03b81 \u2212 \u03b82) = 0 In this case, L1 and L2 intersect at a point that does not lie on C, and E is an ellipse (Fig. 3). If L3 is tangent to E at O, then the platform can assume only the pose q = 0, and the corresponding configuration is both Type 1 and Type 2 singular (there is no self motion)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002155_j.msea.2009.10.058-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002155_j.msea.2009.10.058-Figure8-1.png", "caption": "Fig. 8. Simulated temperature distributions (\u25e6C) for group R1, 600th step.", "texts": [ " 7. The temperature curve in one scaning period has been enlarged and illustrated in the inset graph f Fig. 7. At the end of one single scanning line where the laser eam changes its scanning direction, the temperature exhibits draatic fluctuation. After temperature fluctuation, the temperature n the sintering zone decreases stably because the temperature nfluence of the previous sintered area on the current sintering zone s decreasing slowly with elapsed time due to the heat transfer, as hown in Fig. 8. In other words, the pre-heating temperature of the urrent sintering zone is changing. With the effect of the previous intered line, the base temperature has been changed to a higher alue instead of the initial temperature 200 \u25e6C, as shown in Fig. 8. hat means that with invariable laser sintering parameters, contant temperature sintering cannot be achieved. The elapsed time hich is determined by the scan speed has a significant influence n the range of temperature. In the case of R1, the range of temperture is around 70 \u25e6C. It also proves that active temperature control or the laser sintering process is necessary to attain a constant sinering temperature as well as homogeneous microstructure and hase compositions [22]. Linear regression analysis has been conducted on the simulated \u25e6 emperature curve", " 10a and b, which are produced by balancing the hermal expansion of material in the sintering area. The maximum alues for tensile stress of X and Y components are 26.1 MPa and 1.0 MPa, respectively. The distribution of the Z component stress s different from X and Y component. Tensile stress in Z direcion is present under the laser sintering zone, as shown in Fig. 10c nd the maximum is 87.7 MPa which is much higher than those n x and y directions. Transient thermal stresses are mainly conrolled by the temperature gradient. It can be observed in Fig. 8 hat the temperature gradient in z direction is larger than those n x and y directions. Moreover, the temperature gradient in x irection is larger than y direction which is the laser scanning irection because higher residual heat in this direction causes small temperature gradient, and consequently small thermal tress. .2.2. Transient internal stress In our fabrication process, the ceramic slurry was deposited on a re-heated tile and then dried before laser sintering. The distortion r displacement caused by thermal stress is completely constrained y the substrate", " In the upper part of the model, ensile stresses occur except in the laser sintering zone. But in the ower part of model, compressive stresses arise. The patterns of he stress distribution in x, y, and z directions are also different. erpendicular to the laser sintering direction, x-axis, the compresive stress in the lower part of the model has a larger influence rea in the middle of the cross section than that in both edges. It s similar with the temperature distribution pattern in the cross ection, as shown in Fig. 8. Parallel to the laser sintering direction, -axis, the compressive stress distribution is more flat than that in x irection. This is probably due to the relatively small temperature radient in the y direction. As far as z direction is concerned, the ensile stress appears right under the laser sintering zone because f the expansion actions by the material around the sintering area. he maximum compressive stresses in x, y directions are similar, 56.0 MPa and 264.0 MPa, respectively, and little higher than that n z direction, 245" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002775_optim.2010.5510562-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002775_optim.2010.5510562-Figure1-1.png", "caption": "Fig. 1. Vector diagrams with active flux a d\u03a8 , for PM-RSM", "texts": [ " The paper is organized as follows: Section II The proposed active flux balance for voltage amplitude stabilizing loop; Section III The proposed voltage angle stabilizing loop; Section IV The proposed control system; Section V Case study for high saliency PM-RSM drive and Section VI Conclusions. II. THE PROPOSED ACTIVE FLUX ERROR-BASED VOLTAGE STABILIZING LOOP The core of the proposal consists in driving to zero, through a closed loop, the active flux error between its reference a* d\u03a8 and estimated value a d\u03a8 . The active flux a d\u03a8 [24] of any AC machine is defined as: a d S SqL i\u03a8 =\u03a8 \u2212 \u22c5 (2) Where S\u03a8 is the stator flux vector, Si the stator current vector and Lq is q axis machine inductance. As by this concept all ac machine models \u201cloose\u201d their magnetic saliency [24], (Fig.1): * *( )\u03a8 = \u2212 \u22c5 a d s sc dL L i -for induction motor (IM) in rotor flux coordinates (3) * *( ) a d PM d q dL L i\u03a8 =\u03a8 + \u2212 \u22c5 -for interior PMSM (IPMSM) and permanent magnet reluctance synchronous motor (PM-RSM) in rotor coordinates (4) * *( ) a d dm f d q dL i L L i\u03a8 = \u22c5 + \u2212 \u22c5 -for dc excited synchronous motor (SM) (5) Ls, Lsc -are no-load and short-circuit IM inductances; Ld, Lq - d, q synchronous inductances; Ldm-d axis magnetisation inductance; \u03a8PM - PM flux linkage; If -field current; For active flux estimation (2), stator flux estimation is crucial" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003517_09544062jmes1958-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003517_09544062jmes1958-Figure1-1.png", "caption": "Fig. 1 Rolling bearing structure and characteristic frequencies", "texts": [ " In fact, because the rolling elements experience some random slip [2] (this is partly due to the fact they have a faster spin rotation in the load zone than in the unloaded zone), the periodicity of the shock (Ho and Randall [2] called it a series of equally spaced force impulse excitations, whereas in the present article, it is called the periodicity of the shock) is not exact. This will cause the frequency spectrum of the vibration signal to be made of discrete and continuous parts. There are five basic frequencies related to different fault locations. These frequencies and the parameters required to derive the frequencies are illustrated in Fig. 1. Db and Dc are the ball diameter and the pitch diameter; \u03b8 is the bearing contact angle. Vi, Vc, and Vo are the linear velocities of the inner race, the cage, and the outer race. fr, fc, fb, fi, and fo are the shaft rotation frequency, fundamental cage frequency, ball rotational frequency, ball pass inner race frequency, and ball pass outer race frequency, respectively. The possible fault locations and related frequencies are presented in Table 1. These characteristic frequencies are quite clear and can be used as a reliable source of information for bearing diagnosis [2]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002341_j.snb.2008.03.036-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002341_j.snb.2008.03.036-Figure1-1.png", "caption": "Fig. 1. Structure of bromophenol blue.", "texts": [ " Many reports have demonstrated that quinonimine dyes such as neutral red, toluidine blue, methylene blue, methylene green, and brilliant cresyl blue can be used for preparing chemically modified electrodes (CMEs) [23]. Bromophenol blue (BPB) is a triphenylmethane derivative and an effective electron redox mediator due to its conjugated system. To our knowledge, rare works deal with the determination of NO with this dye molecule. In this work, we reported a new method for preparing sensitive NO electrochemical microsensor by electropolymerizing PBPB composite (Fig. 1) and coating Nafion on a CFME. The Nafion/PBPB composite film-modified microelectrode (Nafion/PBPB/CFME) showed excellent electrocatalytic activity towards the oxidation of NO. The oxidation current linearly increased with NO concentration in the range of 3.6 \u00d7 10\u22128 to 8.9 \u00d7 10\u22125 mol/L with a low detection limit of 3.6 \u00d7 10\u22129 mol/L (S/N = 3). It is clear that the analytical performance of this NO microsensor is comparable or even better than most NO electrochemical sensors reported previously [15,29,14] (Table 1)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002397_2013.23130-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002397_2013.23130-Figure3-1.png", "caption": "Figure 3. (Top) blade and (bottom) blade equilibrium diagram.", "texts": [ " Based on the d\u2019Alembert principle, the rotational equilibrium around the X3 axis is: =.\u2212 Tb JbFC .. (4) where Fb (N) is the axial force on the conrod, and b (m) is the minimum instantaneous distance between the conrod axis and crank rotational center, measured on the X1X2 plane: ))\u2019*sin(*cos )\u2019*sin(1*sin )*sin( 2 + + + \u2212 = + = r rb ( . .. . (5) where \u03b2 (rad) is the conrod angle. Dividing Fb into its X1 and X2 components yields: )\u2019*sin(sin )\u2019*sin(cos 2 2 1 + = = + \u2212= = bbb bbb FFF 1FFF . . .. . (6) So, the blade translational equilibrium in the X1 direction is (fig. 3): 01 =\u2212\u2212\u2212 RFmF aTb x .. 1 . (7) where mT x .. 1 is the inertia force of the translational bodies applied to the blade center of gravity Gl. The translational equilibrium along the X2 axis is: sbs FFFR +=+ 22 (8) where R2 (N) is the counterbar reaction in the X2 direction, while the clamping force (Fs, N) of the counterbar on the blade appear on both sides of the equation since it is autoequilibrated. Hence, the friction force (Fa) due to the total force in the X2 direction is: )( 2 sba FFF + = (9) where \u00b5 is the friction coefficient", " The trend of the function R, as obtained in equation 14, has to be limited to a certain window, considering the counterbar teeth dimensions on the overall path. Finally, the hypothesis of a cutting line is assumed, so that the cutting resistance is idealized to lie on it. This line is parallel to the X1 axis and is assumed to be placed on the bottom surface of the blade teeth, at a distance of 2/3 of the total tooth height from the tooth tip. The friction force (Fa) has ideally the same action line, while the counterbar reaction (R2) belongs to the vertical plane through this line and the X2 axis, in the middle of the blade (fig. 3). Figure 4 shows the shape of the cutting resistance at different values of the total filling percentage (p = p1\u00b7p2) and at different values of the average running speed (\u03c9). In agreement with other authors (Sitkei, 1986) and practice, it can be seen that R decreases with increasing average running speed. As shown in figure 5, analyzing the blade rotational equilibrium, it is possible to identify three couples arising from the vertical and horizontal offsets of the conrod small end from the cutting line: HFh\u2019\u2019h\u2019mh\u2019FC d\u2019\u2019mdFC dFC bTb Tb b " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000262_tmag.2004.824897-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000262_tmag.2004.824897-Figure9-1.png", "caption": "Fig. 9. Contours of hysteresis loss (1000 rpm). (a) Stator core. (b) Rotor core.", "texts": [ " 6 shows the primary current characteristics. It is found that the results with skew are almost the same as those without skew. Fig. 7 shows the torque characteristics. It is found that the results with skew are larger than those without skew. The calculated results agree with the measured ones [7]. Fig. 8 shows the contours of eddy-current loss. It is found that there is a lot of eddy-current loss at the upper and lower parts in the stator teeth surface, and the upper part in the rotor surface. Fig. 9 shows the contours of hysteresis loss. It is found that the hysteresis loss at the upper part in the stator teeth surface and the upper part in the rotor surface are greater than the other part. Table II shows the discretization data and CPU time. We developed a new mesh modification method for the skewed motor analysis using the 3-D FEM. In our method, the auto-mesh generator is used for creating only the initial mesh and it is limited in the air gap of the rotor region. Consequently, it is not necessary to use the auto-mesh generator at each rotation angle, and the computational cost is very small" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003722_lindi.2011.6031137-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003722_lindi.2011.6031137-Figure4-1.png", "caption": "Figure 4. The cross section of the SRM in study", "texts": [ " The output voltage of the linear coupler is proportional to the time-rate-of-change of the phase current, which is modifying due to winding faults [10]. IV. EFFECTS OF FAULTS ON THE EMFS GENERATED IN All the winding faults of a SRM cause unsymmetrical field distribution inside the machine. The best way to emphasize these changes is to perform a numeric field analysis of a sample SRM [11]. The main data of the simulated simple SRM is: i.) Rated power 350 W ii.) Rated voltage 300 V iii.) Rated current 6 A iv.) Rated speed 600 1/min v.) Number of stator poles 8 vi.) Number of rotor poles 6 The cross section of the SRM in study is given in Fig. 4. The numeric field computations were carried out by using the finite elements method (FEM) based Flux 2D program package [12]. The simulation based study was performed for the healthy machine and for a faulty condition where 20% of the A winding's turns were shorted. The magnetic flux distribution in the three significant rotor positions of the SRM (aligned, semi-aligned and unaligned) are given in Fig. 5. \u2013 144 \u2013 As it can be seen in the plots the changes in the magnetic flux distribution due to the minor stator winding fault taken into study are not too significant, but there can be sensed by finely tuned detection circuits" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003521_amm.26-28.676-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003521_amm.26-28.676-Figure1-1.png", "caption": "Figure 1. Schematic diagram of the bearing test setup.", "texts": [ " A series of experiments was conducted on a specially designed rolling bearing test rig. The genetic algorithm based fuzzy neural network (GAFNN) was used as a multivariate information fusion technique to identify the rolling bearing conditions. The aims of this work were to examine comparisons between the integration and separation of wear particle and vibration analyses in machine condition monitoring. Experimental analysis of the 6306 type rolling bearing with man seeded fault has been carried out using a special designed experimental setup (see Fig. 1). The test rig assembly consisted of a transmission shaft driven by an electric motor. The transmission shaft was supported by two 6306 type rolling bearings. A simple hydraulic circuit was used to lubricate the testing bearing. The loading ring and tail spindle were adopted to introduce radial and axial loads to the shaft. The testing bearing chock design includes mounting of the accelerometers in order to obtain the direct transmission path for the vibration data, and the location of a plastic pipette in order to collect the oil samples" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002998_1.3213552-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002998_1.3213552-Figure3-1.png", "caption": "Fig. 3 Summary of flow-field main features from flow visualization", "texts": [ " The flow velocity f the wind tunnel was measured directly using a vane anemomter as 3.5 m s\u22121 0.5 m s\u22121 and indirectly as 3.5 m s\u22121 from a anometer reading of 7.5 Pa and by applying the Bernoulli equaion. The Reynolds number based on the separation between the trips was 2700. Pressure measurements close to the backing ring uggested peak velocities in the region of 20\u201322 m s\u22121. Small ufts were placed on the backing plate, the rotor surface, and the exible strips. The flow-field captured is illustrated in Fig. 3. The ey flow features seen are as follows: 1. downward fluid motion approaching leading edge 2. flow turning to align with strips 3. swirl induced by the inclination of the strips 4. flow separation vertically beneath strip tips 5. flow exiting strips with swirl 6. separation under backing plate and flow contraction 7. flow near floor aligned with axial direction ig. 2 Blow-down of thin widely spaced flexible leaves in a ow speed wind tunnel 41004-2 / Vol. 132, OCTOBER 2010 om: http://turbomachinery", " Consider first the approach to the strips the leading edge of the plates . Before the strips, the flow is aligned with the shaft surface in the axial direction and starts to turn vertically downwards toward the gap under the backing plate as would be the case if there were no strips . Just after the leading edge of the inclined strips, the flow is forced to travel between the strips. This change in direction indicates that a pressure force is acting on the strips here. Consider a stream filament between points 1 and 2 in Fig. 3 , which is illustrated in detail in Fig. 8, showing a typical change in direction of a fluid element at the leading edge of the plates element shown is on the upper surface . The flow along the small stream-tube can be considered like flow round a pipe bend. Hence the momentum equation can be applied in a similar way to that used in the analysis of a pipe bend. Note that , the inlet approach angle, is measured in the plane of the inlet swirl defined by angle . , the leaf entry angle, is measured in the plane of the leaves", " The dashed outline shows where the strips would be positioned. This confirms the presence of a flow direction change suitable for generating a blow-down load, as postulated in Fig. 8. The density and angle of the streamlines in Fig. 9 also imply that the maximum blow-down at the leading edge occurs at the same position, as indicated by the numerical model. Now consider the exit from the strips. The flow here sweeps off the inclined plates and under the backing plate with significant swirl. Consider a stream filament regions 3 and 5 in Fig. 3 showing a change in direction of a fluid element at the trailing edge of the plates. Note that , the leaf exit angle, is measured in the plane of the leaves. The outlet escape angle, , is measured in the plane of the outlet swirl, . When applied in the X horizontal , Y vertical , and Z axial directions for the exit side of the strip, the momentum equation gives OCTOBER 2010, Vol. 132 / 041004-3 ?url=/data/journals/jotuei/28766/ on 03/27/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use T T n T e r d p t f t t 0 Downloaded Fr FX = p3 + 3u3 2 cos sin A3 \u2212 p5 + 5u5 2 cos sin A5 5 FY = p3 + 3u3 2 cos cos A3 \u2212 p5 + 5u5 2 sin A5 6 FZ = \u2212 p3 + 3u3 2 sin A3 + p5 + 5u5 2 cos cos A5 7 he trailing edge blow-down is thus fb = p5 + 5u5 2 A5 cos sin cos \u2212 sin sin 8 he angle is small and can be assumed to be zero for some, but ot all, of the stream-tubes" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002849_j.elecom.2010.10.021-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002849_j.elecom.2010.10.021-Figure2-1.png", "caption": "Fig. 2. Voltammetric curves of RXs in DMF+TBAPF6 at GC electrodes. A1) 1- iodooctadecane (saturated solution). Recurrent sweeps. A2) Cathodic response of the same solution at the modified electrode according to A1 after sonication. In the inset, voltammetry of 1,2;4,5-tetracyanobenzene (concentration 7.8 mM) at polished GC (a) and at modified carbon surface (b). Scan rate in all cases: 0.1 V s\u22121. B1) Behaviour of 4-bromobutyronitrile (concentration: 12 mM) in the course of recurrent scans (scan rate: 0.1 V s\u22121) and B2) after sonication and rinsing.", "texts": [ " Let us recall that such charged material may exhibit both a reducing and nucleophilic behaviour [7,8]. Therefore, at GC electrodes, the classical reduction process already reported with RXs [13] should be reviewed since, in certain cases, the \"electrochemical\" reduction of primary RXs might occur according to a kind of redox catalytic process implying mediator species belonging to the surface. A good support for this proposal is that primary RIs and RBrs (leading to potential shifts upon scans) are often reacting with the GC surface within a potential range with a threshold close to \u22121.7 V. Fig. 2 illustrates this point: curves A1 and B1 relative to 1- iodooctadecane and 4-bromobutyronitrile respectively exhibit Ep/2 equal to \u22122.13 V and \u22122.12 V (first scan). One checks (Fig. 2, A2 and B2) that the negative shift after several scans is large (\u0394EN0.2 to 0.3 V) together with a spectacular decay of peak currents, always noticed with very long alkyl chain RXs. The blocking of GC surface by alkyl chains certainly contributes to slow down the heterogeneous ET. This fact is confirmed by voltammetric responses of some \u03c0-acceptors used as probes (Fig. 2, inset). Moreover, it was established by SECM that the organic layer turned to forbid electron transfers. When the reactivity of long chain RXs is considered, mass increases were clearly noticed. Let us stress, however, that the negative shift would not be Charge of graphitised area (Potential E1, TAAX) Nucle substi RX X Start of alkylation E1 (I) (III) Graphatized zone charge [Cn M +]m [C]m MX e- RX E1 < -1.5 V ET SN Scheme 1. Schematic representation of GC alkylation due to the charge of graphitized zones ( the case where the RX compound is less electroactive than graphitized zones dispersed at the material and that of nucleophilic displacement (SN) are supposed to depend on M+, the le exclusively provoked by the progressive covering of the GC surface (checked by means of SEM imaging), but quite presumably by alteration of graphene-like moieties due to the alkylation process: charged graphitized mediators dispersed at the interface turn to be less and less reducing (because the interfacial SN progressively increases the LUMO level) and therefore, the electron exchange rate towards RX progressively slows down until a limit" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001980_sisy.2008.4664900-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001980_sisy.2008.4664900-Figure1-1.png", "caption": "Figure 1. Illustration of internal and external synergy", "texts": [ " The motion of the joints (change of the internal coordinates) changes the relative position of the links, which results in a simultaneous change of the relative position of the humanoid robot with respect to the environment (change of the system position described in terms of external coordinates). In order the change of internal coordinates yielded a desired change of external coordinates, it is necessary that the robot-ground contact has to be steady during the walk. II. INTERNAL AND EXTERNAL SYNERGY Let us focus on the humanoid robot sketched in Fig 1. Let the coordinate frame representing environment to which external synergy has been defined be xGROGRzGR (Fig. 1a) and let it be fixed to the ground. Let the internal synergy be defined with with respect to the coordinate frame xFOF zF fixed to the foot tip (Fig. 1a). Both coordinate frames are presented only in the sagittal plane: external by the axes xGR and zGR, and the frame fixed to the foot tip by the axes xF and zF, the y axes not being shown in either of the cases. Let us further suppose that the humanoid is walking (Fig. 1 shows the single-support phase) while trying to reach the object O shown in Fig. 1b. The object position requires that the hand (represented by the point H) approach trajectory is parallel to the x-axis of the frame xFOFzF. In case the foot realizes \"proper\" contact with the ground, (i.e. contact over surface, not line or point) the coordinate frames xGROGRzGR and xFOFzF will coincide and hand trajectory will also be parallel to the x-axis of the external coordinate frame xGROGRzGR. Suppose further that a disturbance occurred and that the humanoid started falling forward by rotating about the front foot edge (Fig. 1d). In such a case the frames xGROGRzGR and xFOF zF will not coincide any more (Fig. 1e). They are being rotated by the angle . If the humanoid is continuing to realize joints trajectories exactly as before the occurrence of disturbance the hand trajectory will remain parallel to the x-axis of the frame xFOF zF, but will not be parallel to the x-axis of the external frame xGROGRzGR. In other words, in the case of a \"non-proper\" foot-ground contact the internal and external synergies will not coincide. This means that the hand trajectory with respect to the environment (external synergy) will not depend now only on the joint angles qi, i=1, " ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003955_1747-9541.7.2.371-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003955_1747-9541.7.2.371-Figure1-1.png", "caption": "Figure 1. The Ball Release Angle of a Basketball Shot is Positively Related to the Angle of Ball entry into the Basket An increase in release angle typically leads to an increase in entry angle.", "texts": [ " Tran and Silverberg [8] argued that a 52\u00b0 release angle allowed a great possibility for a successful free throw. Release angle values suggested for successful free throws in these studies are in the range of 49\u00b0 to 60\u00b0, but inter-study comparison cannot easily be made due to differences in release height values. The release angle of the basketball free throw is positively related to the angle of ball entry into the basket [13, 14, 16]; an increase in release angle typically leads to an increase in entry angle (Figure 1). However, when shooting free throws and to have the best chance of passing through the hoop, the ball should approach the hoop with the highest possible angle, allowing for the greatest possible margin for error in all directions [13, 14]. Indeed, shooting the ball at a higher release angle offers the ball a wider effective diameter as the ball passes through the center of the rim cleanly without touching either the rim or the backboard. Furthermore, Brancazio [13] stipulated that there is a trade-off between the respective advantages to be gained from an increased error margin when the ball passes through the basket, resulting from larger release angles, in which the ball enters the basket more steeply, and those from an increased error margin in release speed and angle for values of the latter which require close to the minimum release speed" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002478_jjap.46.4698-Figure15-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002478_jjap.46.4698-Figure15-1.png", "caption": "Fig. 15. Examples of breakdown vibrator shape, when upper mode was driven by high input power.", "texts": [ " The highspeed revolution is due to the large vibration velocity of the stator edge and the thinness of the rotor shaft; nonetheless, the rising time of the rotor is very long because the starting torque is very small. The efficiency is also small because the rotor slips owing to the small preload force. However, the efficiency is about threefold larger than that of the conventional piezoceramic motor reported in ref. 10. An increase in input power increases the vibration velocity and internal stress of the stator vibrator, and finally, the vibrator reaches self-destruction. Examples of the vibrator shape after breakdown examinations are shown in Fig. 15. Additionally, the internal stresses of the coupling modes are analyzed, and the von Mises (equivalent) stress distributions are shown in Fig. 16. We assumed as below that the breaking down of the vibrator starts from a point exceeding the limit strain energy, and it can be expected from the von Mises stress. Figure 16 shows the points of the maximum stress. There are two points appearing out side the center, different from the L1-mode in a simple rectangular plate. The analyzed maximum stress points are also marked in Fig. 15. The breakdown planes of the examples shown in Figs. 15(a) and 15(b) include the maximum stress point. Hence, we assumed that the breakdown starts from the maximum stress point. On the other hand, for an instance, the tip of the vibrator is broken, as shown in Fig. 15(c). In this situation, the contact point appears like a wearing dimple, and the shaft acts as a wedge that triggers a crack. We analyzed breaking strength for the X 128 -rotated Ycut LiNbO3 plate vibrator using a FEM (ANSYS). The relationships between the vibration velocity v of the tip and the maximum internal stress Tmax for some vibrators are shown in Figs. 17\u201320. The breakdown vibration velocity of a rectangular plate whose longitudinal direction is aligned in the x-axis, which is called the x-dir" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002780_cae.20393-Figure7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002780_cae.20393-Figure7-1.png", "caption": "Figure 7 (a) Open water intake. (b) Coupled brakes. (c) Bomb turned on. [Color figure can be viewed in the online issue, which is available at www.interscience.wiley.com.]", "texts": [ " In this mode, a single start command causes a sequence of activities in which the necessary auxiliaries are started, and the necessary safety checks are carried out at each stage of the sequence. All through the startup and shutdown procedures, textual details are delivered to the user in the screen of the virtual environment. In the startup Automatic Startup Procedure sequence, the machine is in the state of \u2018\u2018Stopped Unit.\u2019\u2019 After the startup command is given, the three-dimensional simulation of the startup commences, and a series of requisite conditions is verified automatically simultaneously: the open water-intake gate (see Fig. 7a), the closed distributor, and the lowered generator rotor. In the Figure 7b,c, some of the requisite conditions of startup, necessary for the unit to be ready for the mechanical turn that causes the opening of the distributor are shown. This process begins the rotational movement of the machine, or in other words, the startup. With the requisite conditions of mechanical turn being satisfied, the distributor is ready to open, and the machine enters into unloaded gear without excitation. At 80% of the rating speed, the requisite conditions for excitement are satisfied and the automatic closing of the field circuit takes place; subsequently, the unit enters into excited unloaded gear" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003538_s10847-010-9752-1-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003538_s10847-010-9752-1-Figure4-1.png", "caption": "Fig. 4 Cyclic voltammograms (in 1 mM K3Fe(CN)6 in 0.1 M KCl, scan rate 100 mV/s) obtained at bare gold electrode (a), deposition of CDPSH monolayer (b) and addition of three successive ALP\u2013ADA layers (c\u2013e)", "texts": [ " No decrease in baseline response (corresponding to the CDPSH layer) was observed during this modification\u2013regeneration sequence, indicating that the support CDPSH layer remains stable. Electrochemical characterization of multilayer deposition Electrochemical techniques provide a sensitive and accurate method of characterizing interfacial phenomena. This is usually done by studying the variations in the electrochemical behavior exerted by surface modifications on an electroactive probe in solution. Gold electrodes were prepared with increasing numbers of ALP\u2013ADA and CD-Au layers. Figure 4 shows the CV of a [Fe(CN)6]3-/4- redox probe obtained for the formation of successive layers of ALP\u2013ADA. In CV, an increase of peak-to-peak potential splitting (DEp) associated with a decrease in current intensities are an indication of a blocking effect caused by substance deposition on the electrode surface which inhibits electron transfer from the electroactive probe. Chemisorption of CDPSH at the gold surface provoked a 15 mV increase in DEp with respect to the bare surface (DEp = 80 mV). In contrast, addition of a first, second and layers of ALP\u2013ADA gave DEp values of 350, 410 and 510 mV, indicating the occurrence of a strong blocking effect due to enzyme/nanoparticle deposition, although the current response is not totally suppressed demonstrating that the surface is still permeable to electron transfer" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002605_vppc.2008.4677591-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002605_vppc.2008.4677591-Figure2-1.png", "caption": "Figure 2. Transverse cross section of the LIM", "texts": [ " The influence of the transverse edge effect can be included by an appropriate decrease in the conductivity of the secondary [6]. Therefore, the conductivity of the aluminum plate is corrected by the following Russell and Norsworthy coefficient tanh 21 1 tanh tanh 2 2 Al t OV w k w wk w \u03c0 \u03c4 \u03c0 \u03c0 \u03c0 \u03c4 \u03c4 \u03c4 \u239b \u239e \u239c \u239f \u239d \u23a0= \u2212 \u239b \u239e\u239b \u239e \u239b \u239e \u239b \u239e+\u239c \u239f \u239c \u239f \u239c \u239f\u239c \u239f\u239d \u23a0 \u239d \u23a0 \u239d \u23a0\u239d \u23a0 (2) in which 1 1.3 OV t t d k d \u2212 = + (3) \u03c4 and w are the pole pitch and the width of back iron, respectively. Transverse cross section dimensions OVw , OVt and d are shown in Fig. 2. Similarly, to take into account transverse edge effect in the back iron, the conductivity of back iron is decreased with Gibbs\u2019s coefficient corek 2core wk w \u03c0 \u03c0 \u03c4 = + (4) The 2-D time-stepping finite element analysis was conducted by using Galerkin\u2019s method [7]. Only one pole pitch model is inappropriate, as the LIM is not periodically symmetrical due to end effect, therefore, the entire motor need to be studied. Fig. 3 shows the meshes for the main part of analysis model. In order to make flux density and force calculation more accurate, it can be seen from Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003369_j.mechmachtheory.2010.12.002-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003369_j.mechmachtheory.2010.12.002-Figure1-1.png", "caption": "Fig. 1. Distance between two circles in 3D space.", "texts": [ " An application of a homogenous matrix transformation and Denavit\u2013Hartenberg's notation [11] constitute the mechanism model analysis method. In order to solve the analysis task, a closed-loop vector equation is established at the contact point of the elements which form a higher kinematic pair [4]. This method is widely applied in the analysis of gear mechanisms, among others [12]. Calculation of the distance between two tori located freely in space can be determined by calculating the distance between their basic circles as shown in Fig. 1. The circles, freely located in relation to one to another in space, can be describedwith the help of coordinates of both circle centres Qi and directional unit vectors of their axes Ti (i=1,2\u2212 circle no.). The distance of l=A1A2 is the shortest distance between the axes of the circles, distances h1=P1A1 and h2=P2A2 are distances between points Ai and the centres of these circles Qi. Angle \u03b7 is the angle between the axes of the circles. The coordinates of points on the circles can be presented in the form of the following vectors: T1 = 0 0 1 0 2 664 3 775; T2 = 0 \u2212sin \u03b7 cos \u03b7 0 2 664 3 775;R1 = r1cos \u03c31 r1sin \u03c31 h1 1 2 664 3 775;R2 = r2cos\u03c32 + l r2 sin \u03c32 cos \u03b7\u2212h2 sin \u03b7 r2 sin \u03c32 sin \u03b7 + h2cos \u03b7 1 2 664 3 775: \u00f01\u00de \u03c3i\u2014 angular location of Ri points on the circles, between which distance d is the extreme distance (Fig. 1). The distance where between any Ri points of the circles will be then presented with the following formula: d = ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R2x\u2212R1x\u00f0 \u00de2 + R2y\u2212R1y 2 + R2z\u2212R1z\u00f0 \u00de2 r : \u00f02\u00de The necessary condition for this distance to be extreme is that the following derivatives: \u2202d \u2202\u03c31 = r1 d r2 cos\u03c32 + l\u00f0 \u00desin\u03c31 + h2 sin \u03b7\u2212r2 sin \u03c32 cos \u03b7\u00f0 \u00decos \u03c31\u00bd \u00f03\u00de \u2202d \u2202\u03c32 = r2 d r1cos\u03c31\u2212l\u00f0 \u00de sin \u03c32\u2212 r1 sin \u03c31cos \u03b7 + h1sin \u03b7\u00f0 \u00decos\u03c32\u00bd \u00f04\u00de imultaneously equal zero, based on the assumption that d\u22600" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003619_s11012-010-9331-y-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003619_s11012-010-9331-y-Figure9-1.png", "caption": "Fig. 9 Gear teeth addendum collision (a) and free vibration caused by teeth impact (b)", "texts": [ " In gear pair meshes between the meshed teeth all of these processes exist. A few kinds of teeth impacts arise in the mesh. More important is the addendum impact which is the result of gear pitch difference caused by elastic deformations. 4.1 Gear system restorable free vibrations Teeth deformations are proportional to teeth load and teeth stiffness. Deformations replace the first point of contact from the right position A, to position A\u2032 which is ahead of point A. The contact of teeth pair starts with intensive addendum impact (Fig. 9a). Collision speed vc is proportional to teeth deformation, speed of rotation n and gear design parameters. By analyzing teeth geometry, deformations and speeds, collision speed at the first point of teeth contact is defined and presented in [10]. Every individual teeth impact produces natural free vibration of the gears (Fig. 9b) with natural frequency fn. By strong inside damping these vibrations attenuate in a short time. The next teeth pair entering the mesh collides again and again. Vibrations become restorable after every teeth impact. For relatively slow gear rotation the measured time function of restorable free vibrations is presented in Fig. 10a. The teeth mesh frequency f corresponds to gear revolutions and to gear teeth number. If the speed of the gear rotation is slow enough, the time between the two teeth impacts 1/f is higher than the time necessary for the free vibration attenuation", " The objective of modeling was to prove the hypothesis that gear vibration is a restorable natural one and then to present the vibrations in the form suitable for use as constraints in design parameters definition in gear drive design process. The total level of gear vibration x\u0308 is divided into two parts, continual x\u0308a and transient x\u0308b (Fig. 10b). x\u0308 = x\u0308a + x\u0308b = x\u0308a(1 + \u03b6T sin\u03d5) = A c\u03b3 x0 me f (1 + \u03b6T sin\u03d5) (4) The gear pair is modeled as a single mass oscillator, where me is equivalent mass reduced in collision direction (Fig. 9a), c\u03b3 is the mean gear teeth stiffness, x0 is teeth deformation amplitude in the moment of teeth collision, \u03b6T is transfer function between force transferred to the gear masses and collision force Fc with phase angle \u03d5, A\u2014is coefficient of energy absorption inside the vibration system, and f is the gear teeth mesh frequency. The model presented by (4) and by the diagram in Fig. 10b is developed using the theory of singular systems. According to this theory, the singular process consists of the two processes, continual and transient" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003340_bf00925745-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003340_bf00925745-Figure8-1.png", "caption": "Fig. 8. Reentry vehicle coordinate system.", "texts": [], "surrounding_texts": [ "No results are displayed. The results of the examination of Method 2 of Section 4 of Ref. I are shown in Fig. 5. Since P was fixed a priori , the method does not converge as rapidly as the ordinary QL-method although it does converge over a somewhat much wider region. The test problem was the brachistochrone problem with e --= 0.5. The results from Method 3 are shown in Figs. 6 and 7. This method of extending the range of convergence works far better than the other methods examined and considerably better than the usual OL-method. In Fig. 6, two iterations with Method 3 (e == 0.5) were followed by three iterations with the usual QL-procedure, for a total of five iterations. The results are compared with five iterations with the usual QL-procedure. In Fig. 7 four iterations with Method 3 (E = 0.25) were followed by three iterations with the usual ~L-procedure, for a total of seven iterations. The same standard of comparison is used as in Fig. 6. Both Figs. 6 and 7 show that, at the expense of computing time, a much wider range of convergence can be obtained with Method 3 than is available with the ordinary ~L-method and that, the higher the price, the wider the range of convergence obtained.\n3. R e e n t r y T r a j e c t o r y P r o b l e m\nThe reentry vehicle trajectory problem is an appropriate choice for a more complex problem. It is well known that this problem is computationally difficult because of integration instability and sensitivity of the adjoint variables. The equations of motion are those used by Scharmack (Refs. 3 and 4) as well as Breakwell, Speyer, and Bryson (Ref. 5) and are similar to those used by Payne (Ref. 6). These equations are simplified considerably but are still realistic physically. The reentry vehicle is assumed to have a low lift-drag ratio. The simplifications include the use of an exponential model for the atmosphere and the use of a simple lift-drag polar.\nCases Inves t iga ted . Two cases were formulated to see what differences would arise. The first case is similar to that studied by Breakwetl, Speyer, and Bryson (Ref. 5) and Payne (Ref. 6); the second case is that solved by Scharmack (Refs. 3 and 4). The primary differences are in the initial and final conditions, the control polar, and, with respect to Payne's paper (Ref. 6), the use of a more accurate gravity approximation. The criterion function also varies: Scharmaek used both convective and radiative heating, Breakwell used velocity, while Payne considered convective heating and sensed acceleration. These cases are defined explicitly in Table 1.", "The quantity to be optimized (minimized) is the time integral of the total stagnation point heating per unit area and includes either the convective term alone or both the radiative and convective terms.\nThus, the following formulas hold:", "Convergence rate from a piecewise linear approximation t o the final solution for unbounded control." ] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.80-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.80-1.png", "caption": "Fig. 14.80 Truck-mounted elevating platform", "texts": [ " 1 2 3 4 A mobile elevating platform with a scissor extending structure. 2 3 1 A truck-mounted elevating platform whose extending structure is usually in the form of a two-stage boom. The elevating platform performs the following motions: \u2022 Raising and lowering \u2022 Slewing relative the vertical axis perpendicular to the base 1 2 3 Key: 1 platform; 2 extending structure; 3 chassis; 4 struts Truck-Mounted Elevating Platform The truck-mounted elevating platform consists of the following parts shown in Fig. 14.80. The supporting frame is a body to which protractible struts, a hydraulic feeder, and a slewing gear are fixed. The supporting frame is secured to a truck chassis. Part B 1 4 .5 The slewing column is attached to the supporting frame through a crown-bearing. The column is a slewing welded-construction frame with a transmission gear and a brake mounted on it. The lower boom stage and the cylinder are attached to the column by articulated joints. The boom consists of two stages connected by an articulated joint" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002589_tmag.2008.2005137-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002589_tmag.2008.2005137-Figure6-1.png", "caption": "Fig. 6. Mode shapes of the base. (a) First mode, (b) second mode.", "texts": [ " And when the flying height is smaller, the effect of intermolecular force will be more obvious. This is because when the flying height is smaller (before contact happens), the attractive force will be more dominant. The response of the first five modes response of the base for the 250 g, 1 ms shock is shown in Fig. 5. It can be seen that the first two modes give the main contribution to the base deformation. The corresponding natural frequencies are 2844 and 3592 Hz, respectively. These two mode shapes are shown in Fig. 6. Since the base stiffness will directly affect the natural frequencies of the base, the effect of base stiffness on the shock response can be investigated by numerically changing the stiffness matrix in (4) through changing the natural frequencies of the base. If the mass of the base remains unchanged, a higher natural frequency means a higher stiffness. In the case when the shock amplitude is 230 g and the duration time is 1.0 ms, Fig. 7 gives the comparison of flying attitudes and dimple spacing for HDD with a soft base and with a stiff base" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003704_0954409711427835-Figure13-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003704_0954409711427835-Figure13-1.png", "caption": "Figure 13. Wheel cross representation of the experiment. As a function of the hunting oscillation experimented in each moment and the lateral displacement suffered, the point of contact differs and the diameter and perimeter are slightly different.", "texts": [ " Moreover, both wheels will suffer the same movements, so the relationship will be similar. There is another strategy to analyse the data. For the case of long distances, the effect of topography and the displacements and deviations suffered by the wheels affect the measures. This can be due to many factors, such as hunting oscillations of the railway units, nonperfect curvature of the wheels, variations in the speed, rigidity of the railway units, etc. Because of this, the area of contact between the rail and wheel changes slightly, as shown in Figure 13. Therefore, the diameter will also change in order to adapt itself to the conditions of the rails and consequently the rotational speed of the wheel. The conclusions drawn from the analysis of some causes of the rolling for all of the tests are presented in this section. There is a direct relation between the deviation of the rotational speeds and the type of route. Several methods are available for analysing railway transition curves and transition forms,18 where the type of transition, length of the curved tracks and train speed are decisive factors that have significant influence on the dynamic performance" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003167_iecon.2009.5415436-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003167_iecon.2009.5415436-Figure1-1.png", "caption": "Fig. 1. Parameters of the kinematic chain.", "texts": [ " Finally, acknowledgments and references complete the paper. An n-trailer system can be considered as a kinematic chain consisting of n + 1 units. Unit 0 is the tractor vehicle, and trailers are numbered from 1 to n. Local coordinate frames are attached to each unit i, with the Yi axis coincident with the longitudinal axis and the Xi axis lying on the rear axle. The state of this system can be defined by the position and orientation of the tractor, plus the angle of each joint. The parameters of the kinematic chain are shown in Fig. 1. Distances Li f and Li b are always positive constants for 978-1-4244-4649-0/09/$25.00 \u00a92009 IEEE 2385 trailers with off-axle hitching. The relative angle of the ith trailer with respect to the (i \u2212 1)th unit is represented by \u03b8i. \u03b80 is the tractor\u2019s heading with respect to the global coordinate system XY . All these angles are considered counterclockwise positive. Heading \u03c6i of the ith trailer with respect to the global frame can be computed as: \u03c6i = i\u2211 j=0 \u03b8j . (1) Besides, the relative angular velocity \u03c9i of the ith unit with respect to the preceding one is defined by: \u03c9i = d\u03b8i dt ", " The angular \u03a9i and linear vi speeds of trailer i depend on how v0 and \u03a90 are propagated through the kinematic chain: vi = \u2212vi\u22121 cos (\u03b8i) + \u03a9i\u22121 Li\u22121 b sin (\u03b8i), (5) \u03a9i = \u2212vi\u22121 sin(\u03b8i) + \u03a9i\u22121 Li\u22121 b cos(\u03b8i) Li f , (6) where dynamic effects are neglected under the assumption of not overloaded trailers moving with moderate velocities and accelerations [1]. B. Virtual tractor for backward motion control Backward motion control of an articulated vehicle can be implemented by considering the last trailer n as a virtual tractor that moves forward (see Fig. 1). In this way, the actual tractor is commanded to act as the last trailer so that the articulated system moves forward virtually. The virtual tractor local axis XY v 0 is the same as the last trailer local frame XYn but rotated 180\u25e6. Thus, the position and heading of the virtual tractor with respect to the global coordinate system XY are: xv 0 = xn, yv 0 = yn, \u03c6v 0 = \u03c6n + \u03c0. (7) Each control interval for the trailer system in backward motion requires three steps: 1) Compute position (xv 0, yv 0) and orientation \u03c6v 0 of the virtual tractor with respect to the global coordinate system XY ", " The vehicle top speed coincides with this limit in straight-line motion, but it decreases to zero according to the increase in the demanded curvature. The track speed controller runs in an on-board DSP every 10 ms, which also provides odometric data every 30 ms. The tractor can be controlled as a differential drive vehicle using the approximated kinematic model presented in [13]. This also allows to define the virtual rear axle of this vehicle, which is relevant to determine hitch parameter L0b (see Fig. 1). A Sick LMS 200 time-of-flight laser scanner is mounted on the forward part of the vehicle at a distance of 0.5 m ahead of its coordinate center. To correct odometric estimations of the actual tractor, an accurate laser scan matching technique has been employed every 270 ms [14]. The dimensions of the two wheeled trailers are similar to the tractor. The first trailer is employed for carrying loads, while the second one is for spraying (see Fig. 3). Each angle \u03b8i is indirectly obtained with a draw-wire displacement sensor" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000224_1.2114987-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000224_1.2114987-Figure10-1.png", "caption": "Fig. 10 Nominal, actual, and measured profiles, m=4 mm.", "texts": [], "surrounding_texts": [ "The measurement of tooth profiles for each gear was made six times. For each test gear five teeth were chosen to compare the results obtained from the developed measuring technique and a well-established method on a toolmaker\u2019s microscope. The number of examined teeth is limited by the measurement and evaluation time required by the traditional technique. The profile error obtained from the toolmaker\u2019s microscope is considered as the actual profile error. Tables 3\u20135 list actual profile errors of some selected teeth on the 3-, 4-, and 6-mm-module gears, respectively. The developed software was used to evaluate the tooth profile error and plot the nominal, actual, and measured flank profiles. Figure 8 shows the profile shapes for all teeth of the 6-mm-module gear. Figures 9\u201311 show the profile plots for the results of the first tooth flank of the 3-, 4-, and 6-mm-module gears, respectively. The profile errors obtained from the measurement system for all teeth of the same gear were found to be equal. Table 6 lists the profile error obtained from the developed system for those gears. Examination of the profiles shown in Figs. 9\u201311 shows that the actual and measured profiles are quite similar. Further investigation shows that the smaller irregularities seen in the actual profiles are not present in the measured profiles; this is also indicated by the values listed in Table 6. This can be attributed to the relatively large image size, which tends to produce an averaging effect that overrides minute variations in the tooth profile. Variations observed when using the toolmaker\u2019s microscope may be attributed to measurements being made at different lateral gear sections or to edge effects. Figures 12\u201314 plot the relationship between the actual and measured profile developed system errors for the five selected teeth on the 3-, 4-, and 6 -mm-module gears, respectively. Examination of those figures shows that the profile error values evaluated using the developed system and the toolmaker\u2019s microscope method are close to each other. It can also be seen that the profile errors evaluated using the developed system for all teeth of a single gear are almost equal in value. These two observa- tions may be attributed to the inability of the system, in its Optical Engineering 103603-6 ed From: http://opticalengineering.spiedigitallibrary.org/ on 05/15/2015 Term Fig. 12 Actual and measured profile errors, 3-mm-module gear. Fig. 13 Actual and measured profile errors, 4-mm-module gear. October 2005/Vol. 44 10 s of Use: http://spiedl.org/terms R b t a p l m i a t m m n e a q i n Download current configuration, to sense the minute variations on the tooth profile, which was previously discussed." ] }, { "image_filename": "designv11_20_0000413_robot.2003.1242208-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000413_robot.2003.1242208-Figure2-1.png", "caption": "Fig. 2: Experimental setup", "texts": [ " In Section 5, we propose applying hybrid position/force control to metal spinning, and experimentally verify the effectiveness of our proposed control method. 2 Experimental setup First, we developed an experimental setup for conducting basic experiments on metal spinning. Forming of different materials under various parameters can be done to gather forming data, e.g. success or failure of forming, precision of product, wall thickness, surface roughness, and forming force. In particular, generation of defects (wrinkles or breaks) can be analyzed by monitoring the conditions around their occurrence. Figure 2 illustrates our experimental setup. The linear motion of 2: and y axes is driven by the ballscrews (2 mm/rev) and DC servo motors (60W). The mandrel (0 axis) is rotated by a DC servo motor with a planetary gear (redoct,ion rate: 1/10), The 0 axis is slanted relative t,o t,he z axis by a/3 rad. Each motor has an incremental encoder (4000 p/r) for detecting t~he rotation angle. The specificat,ions of each axis are listed in Table 1. The diameter of the forming roller is 70 mm. The roundness of the edge is a 9" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003619_s11012-010-9331-y-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003619_s11012-010-9331-y-Figure8-1.png", "caption": "Fig. 8 Bearing failure probability testing", "texts": [ " teeth flanks hardness effectuates all indicators of ultimate flank stress distribution. In Fig. 7 some of the chosen examples are presented. Increase of the teeth flank hardness significantly increases ultimate flank stress, i.e. translates the range of failure probability distribution in the area of higher stresses. 3.3 Failure probability identification of bearings and seals Failure probability testing of the bearings is carried out according to the principle similar to the gear wear probability testing. In Fig. 8a one of the possible testing rigs is presented. The range of failure probability distribution is bounded by inclined straight lines in a double logarithmic coordinate system (Fig. 8b). For the straight lines definition, it is necessary to make the tests at the two levels of force acting at the bearing. Similar to the gear testing at the both force levels, it is necessary to test the collection of the same bearing type. Using the revolution numbers to the bearing failure N , the Waibull\u2019s functions of failure distribution PF (N) are defined. For those functions, in the next step, the range of bearing failure probability distribution is bound. Boundary lines for failure probability PF = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003998_icat.2013.6684058-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003998_icat.2013.6684058-Figure2-1.png", "caption": "Fig. 2. Schematic diagram of the inverted pendulum", "texts": [ " Indeed at each step of evaluation, we transform the nonlinear reference and plant models to constant LTI models and by using MRAC scheme, we try to control it and force the output of plant ( states) to track the reference model output ( states). The main objective of the proposed algorithm is to adapt the LQR (or Pole Placement) controller designed for a nonlinear reference model (known) to a given nonlinear plant dynamics (unknown). IV. AN APPLICATION: INVERTED PENDULUM The proposed algorithm shall be applied to a physical system. The model used here is an inverted pendulum model that has been studied in [12]. Fig. 2 shows the schematic view of the system. The equations of motion for the inverted pendulum mechanism are given as follows, *\u03b8\u0308 = ./123 \u03b8 \u2212 .1 \u03b8\u0308 \u2212 .1\u0308452\u03b8 \u0308 = \u2212 .6\u0308 + 1\u03b8\u03084527 \u2212 1\u03b8\u0307 23 \u03b88 \u2212 9\u0307 Then 7\u0308 and \u0308 can be written as \u03b8\u0308 = (. + )./123 \u03b8 \u2212 .1 452\u03b8 \u2212 (.1\u03b8)\u0307 23 \u03b8452\u03b8 + .19452\u03b8\u0307 (* + .1 )(. + ) \u2212 . 1 452\u03b8 \u0308 = \u2212. 1 /23 \u03b8452\u03b8 + (* + .1 ) + .1\u03b8\u0307 23 \u03b8(* + .1 ) \u2212 9(* + .1 )\u0307 (* + .1 )(. + ) \u2212 .212452\u03b82 where, and m denote the mass of the cart and the pendulum respectively, is the applied force to the cart which is admitted as the control signal" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001569_2007-01-0136-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001569_2007-01-0136-Figure3-1.png", "caption": "Fig. 3. Force distribution for bearing loadings evaluation", "texts": [ " Differentiating the above equation with respect to time, we get the instantaneous piston velocity pist 2 2 cosu ( ) r sin 1 1 sin (4a) while differentiating once again with respect to time, we get the instantaneous piston acceleration 2 4 pist2 2 2 3 / 2 2 ucos2 sin 1b( ) r cos r(1 sin ) (4b) The last term on the right hand side of Eq. (4b) takes into account the crank\u2019s angular acceleration \u2018 \u2019 in the piston acceleration. The loading of the bearings can then be computed from the following equations, with reference to Fig. 3 [25,26] 0x rod.l 0y rod.l B m bsin F( )B m b cos cos (5) for the connecting rod small end bearing, 1x rod.r pist 2 1y rod.r B m u cos F( )B m r cos( cos ) (6) for the connecting rod big end bearing, 2x 2 2y rod.r F( )B sin( ) cos F( )B cos( ) m cos r (7) for the crank pin, 3x 2 3y rod.r crank F( )B sin( ) cos F( )B cos( ) (m m ) cos r (8) for the crank journal1, and 2 4x rod.r crank 2 4y rod.r crank B F( ) tan (m m ) r sin B F( ) (m m ) r cos (9) for the main crankshaft bearing. The corresponding total bearing force is then 2 2 i ix iB B B y (10a) and the angle , shown in Fig. 3, is given by 1 ix i iy B tan B (10b) with i = 0\u20264 according to the bearing studied. In Eqs (5-9), corresponds to the connecting rod angle (see also Fig. 2), i.e. 1 2 2cos [ (1 sin )] . The total force acting on the piston is composed of the gas and the inertia force, i.e. , which then propagates into the thrust force and the force in the direction of the connecting rod . The gas force is determined by while the reciprocating masses (inertia) force by , with the reciprocating mass and , i.e. the connecting rod is assumed equivalent to two masses, one reciprocating with the piston assembly and the other rotating with the crank" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001239_roman.2006.314421-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001239_roman.2006.314421-Figure3-1.png", "caption": "Fig. 3. Model of the pneumatic passive element with a rotation axis : (a) laminated sheets and (b) friction plane", "texts": [ " 2 illustrates the two modes of the pneumatic passive element system. When the air is caused to flow from the air buffer to the passive element by a force external to the air buffer, the newly developed passive element can rotate freely around the rotation axis. (See the free mode in Fig. 2.) When the air returns from the passive element to the air buffer removal of the force external to the air buffer, the air pressure presses the laminated sheets down, and the passive element becomes firm. (See the constraint mode in Fig. 2.) Fig. 3 illustrates the detail model of the pneumatic passive element with a rotation axis. The constraint torque T on the laminated sheets is expressed in the following equation: T = (n + 1) \u222b 2\u03c0 0 \u222b R 0 \u03bcr2Pdrd\u03b8 = 2 3 \u03c0\u03bc(n + 1)R3P (1) where \u03c0, \u03bc, n, R and P are the ratio of the circumference of a circle to the diameter, the coefficient of static friction, the number of sheets, the radius of the friction plane, and the vacuum pressure. The parameters dr and d\u03b8 denote the infinitesimal radius and the infinitesimal angle as illustrated in Fig. 3(b). Table I presents the configuration of the pneumatic passive element with a rotational axis. Substituting these parameters into (1), the theoretical maximum constraint torque is calculated as 4.07 [Nm]. We developed an AFO using a pneumatic passive element with a rotational axis. (Hereafter, we refer to this as the \u201cdeveloped AFO.\u201d) Fig. 4 presents the prototype AFO, which was made by cutting a conventional shoehorn-type AFO off at the ankle and attaching the proposed pneumatic passive element with a rotational axis at the ankle" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000507_2006-01-0582-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000507_2006-01-0582-Figure5-1.png", "caption": "Figure 5. Definition of rotational phase angle from the housing end view", "texts": [ " Using the defined coordinates of the tri-axial forces shown in Figure 3, one can obtain the combined rotational transformation matrix. By multiplying each individual transformation matrix, as depicted in Figure 4, in sequence, the following equation that relates the measured internal forces to the global forces in accordance with the housing coordinate is given by: Fz Fy Fx zF yF xF )90(sincos)90(cossin)90cos(]cos1[)90sin()90cos( sin)90cos(cossin)90sin( ]cos1[)90sin()90cos(sin)90sin()90(sincos)90(cos 22 22 (1) Where, the CV joint rotational phase angle is defined as shown Figure 5, from the housing end view. Thus, by using this relationship one can calculate the net friction coefficient along the housing groove by using the relationship as shown in the equation (2) below. xF yF P Q (2) The three major friction components and friction coefficients inside of the tripod CV joint are as depicted in Figure 6. Skidding friction, which is represented by SK, is generated along the housing groove. This friction is believed to be the dominant friction that causes the GAF. The rolling friction, represented by the rolling friction coefficient RL, is the pure rolling friction term of the roller, and this value could be estimated from the slip to roll ratio during the plunging motion", " To investigate experimental repeatability, several experiments under identical experimental conditions were performed and confirmed. The grease used was a standard petroleum hydrocarbon with solid EP additives CV joint grease and was applied inside of the CV joint housing. In these experiments the rubber seal (boot) that seals the grease in the CV joint housing was not installed for easy accessibility and visualization. Note that all tests were conducted at the position when the trunnion with the tri-axial force sensor is in the top position, which is defined as a 90 \u00b0 of rotational phase angle in Figure 5. Typical experiments captured over two complete cycles (60 seconds) under 160 Nm static torque are depicted in Figure 8, at two different articulation angles of 0 \u00b0 and 15 \u00b0. As shown in Figure 8 (g), the experiments start with the CV joint spider being in the center position of the housing (at time t = 0), then moving inwards in the housing for 15 mm (at t = 7.5 sec it reaches the end position), and then moving all the way towards the outward position of the housing, passing through the center position (at t = 22", " Thus, it is reasonable to assume that the friction properties at higher plunging velocities (estimated to be about 330 mm/s under maximum engine rpm) will be similar to the values obtained under the present test conditions. As in an actual vehicle the CV joint is rotating, next we evaluate the effect of rotational phase angle on the friction coefficient. Tests were conducted at 26 equally spaced locations (phase angles) in the 360 \u00b0 of overall phase angles (definition of phase angle is shown in Figure 5). Figure 13 shows the summary of the test results under grease lubricated conditions for the articulation angle of 15 \u00b0. In Figure 13 (a), error bars represent the fluctuation of the friction coefficient during a cycle of CV joint stroking motion, and the solid circles the mean value at a phase angle. The results clearly show that the friction coefficient is changing in terms of the rotational phase angle due to the CV joint kinematics and orbiting trajectory of its center position. Interestingly, the friction coefficient mean values, represented by the solid circles show a sinusoidal trend with phase angle changes, and one can thus use a simple sinusoidal function to curve-fit and describe this behavior. By using the experimental data shown in Figure 13 (a), the amplitude of the friction coefficient loop is plotted separately as shown in the bar graph of Figure 13 (b). From this graph, one readily observes that the friction coefficient amplitude of the loop is highest when the spider (that includes the tri-axial force transducer) is at 90 \u00b0 and 270 \u00b0 (top and bottom positions shown in Figure 5). Using the friction coefficient behavior as a function of the rotational phase angle (sinusoidal curve fit in Figure 13 (a)), Figure 14 is obtained and shows the trends of averaged friction coefficient at each trunnion as well as the combined result. As a tripod type CV joint has three trunnions, each individual trunnion exhibits sinusoidal friction behavior with 120 \u00b0 of rotational phase angle difference. Thus, the combined friction coefficient becomes constant with the friction coefficient value of about 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003022_1542362.1542415-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003022_1542362.1542415-Figure1-1.png", "caption": "Figure 1: (a)Two bodies connected by a hinge in 3-dimensional space. (b)A body-and-hinge framework.", "texts": [ " In past years, despite the absence of the rigorous proof of the Molecular Conjecture, empirical data have been accumulated that support this conjecture [2, 24, 31]. In this paper we are able to settle the Molecular Conjecture affirmatively in R3 and in higher dimensions that provides the theoretical validity of the algorithms behind such software as FIRST, FRODA, etc. Body-and-hinge frameworks. A 3-dimensional body-andhinge framework is the collection of rigid bodies in R3 connected by hinges, where the bodies are allowed to move continuously in R3 so that the relative motion of any two bodies connected by a hinge is a rotation around it (see Figure 1). The framework is called rigid if every such motion provides a framework isometric to the original one. This mathematical model of physical structures can be naturally extended to d-dimensional case, where a d-dimensional body-and-hinge framework consists of d-dimensional rigid bodies connected by (d \u2212 2)-dimensional affine subspace, i.e. pin-joints in 2- space, line-hinges in 3-space, plane-hinges in 4-space and etc. The formal definitions of body-and-hinge frameworks and the rigidity will be given in the next section", " In this paper, for our special interest in (D \u2212 1)G, we shall use the simple notation eG to denote (D \u2212 1)G and let eE be the edge set of eG. Tay [22] and Whiteley [28] independently proved that the generic infinitesimal rigidity of body-and-hinge frameworks (defined in Section 2) is determined by the underlying graphs as follows. Proposition 1.1. ([22, 28]) A graph G can be realized as an infinitesimally rigid body-and-hinge framework in Rd if and only if eG has D edge-disjoint spanning trees. For example Figure 2(a) shows the underlying graph G of the body-and-hinge framework illustrated in Figure 1(b). Figure 2(b) indicates six spanning trees in G. Observe that each edge of G is contained in at most five spanning trees, which implies that eG(= 5G) contains six edge-disjoint spanning trees. Therefore Tay and Whiteley\u2019s Theorem (Proposition 1.1) ensures that G can be realized as an infinitesimally rigid body-and-hinge framework (G,p) in R3. The Molecular Conjecture. A body-and-hinge framework (G,p) is called hinge-coplanar if, for each v \u2208 V , all of the (d\u22122)-dimensional affine subspaces p(e) for the edges e incident to v are contained in a common (d\u22121)-dimensional affine subspace (i" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001635_978-1-4020-6114-1_10-Figure10.7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001635_978-1-4020-6114-1_10-Figure10.7-1.png", "caption": "Fig. 10.7. Schematic representation of formation of a complete path by successive B-Spline segments projected on the horizontal plane.", "texts": [ " The remaining two control points are allowed to take any position within the scanned by the radars known space, taking into consideration given constraints. When the next path segment is generated, only the first control point of the B-Spline curve is known; it is the last control point of the previous BSpline segment. The second control point is not random, since it is used to guarantee at least first derivative continuity of the two connected curves at their common point. Hence, the second control point of the next curve lies on the line defined by the last two control points of the previous curve as shown in Figure 10.7. It is also desirable that the second control point is near the first one, so that the UAV may easily avoid any obstacle suddenly sensed in front of it. This may happen because the radar scans the environment not continuously, but at intervals. The design variables that define each B-Spline segment are the same as in the off-line case, i.e., seg_lengthk,j , seg_anglek,j, and zk,j (k=2, 3, and j=1,\u2026,N). The path-planning algorithm considers the scanned surface as a group of quadratic mesh nodes. All ground nodes are initially assumed to be unknown" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000774_1.7143-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000774_1.7143-Figure1-1.png", "caption": "Fig. 1 Deformation of membrane.", "texts": [ " Detailed explanations of the preceding criteria are found in Refs. 5 and 20. In this subsection, we establish the basic equations of membrane kinematics, neglecting the wrinkling effects. Derivation using the tensorial components in the convected coordinate system21 is presented. The equations presented here are directly applicable to taut regions of membranes because wrinkles are absent in these regions. Consider a deformation of a membrane placed in threedimensional Cartesian coordinate system ei (i = 1, 2, 3), as shown in Fig. 1. The convected coordinate system r\u03b1 (\u03b1 = 1, 2) is employed to locate a point on the membrane midsurface. Here and in what follows, we suppose that Greek indices run through values 1 and 2. The covariant base vectors in the undeformed and deformed configurations of the membrane are defined, respectively, as G\u03b1 = \u2202X(r\u03b1) \u2202r\u03b1 (2) g\u03b1 = \u2202x(r\u03b1) \u2202r\u03b1 (3) where X(r\u03b1) and x(r\u03b1) are the position vectors of a point on the membrane midsurface in the undeformed and deformed configurations, respectively. The contravariant base vectors G\u03b1 and g\u03b1 can be determined so that the following relations are satisfied: G\u03b1 \u00b7 G\u03b2 = \u03b4\u03b1 \u03b2 (4) g\u03b1 \u00b7 g\u03b2 = \u03b4\u03b1 \u03b2 (5) where \u03b4\u03b1 \u03b2 denotes the Kronecker delta (\u03b4\u03b1 \u03b2 = 1 for \u03b1 = \u03b2 and \u03b4\u03b1 \u03b2 = 0 for \u03b1 = \u03b2)" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002304_amr.47-50.13-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002304_amr.47-50.13-Figure1-1.png", "caption": "Fig. 1: Residual Stress Components", "texts": [ " Various authors [1, 2] have studied the effect of machining parameters in respect of process variables like residual stress and surface finish. It has been found reasonably good enough to select low (less severe), medium (severe) and high (more severe) levels of machining parameters to study their effects on these process variables. Residual stress is essentially confined to surface. Hence, bi-axial stress analysis by x- ray diffraction technique [3-8] is generally adopted. Residual stress has been described in terms of in-plane stress, \u03c3\u03c6 as well as its principal stress components \u03c31 and \u03c32 and angle of rotation, \u03c6 (Fig. 1). From a knowledge of principal stresses and their orientation with respect to the milling direction (\u03b8), residual stresses along the milling direction (\u03c3m) and normal to milling direction (\u03c3n) as well as shear residual stresses (\u03c4mn) were evaluated by using equations \u03c3m + \u03c3n = \u03c31+\u03c32 and \u03c3m - \u03c3n = (\u03c31-\u03c32) Cos 2\u03b8 (1) \u03c4mn= [(\u03c31-\u03c32)/2] Sin 2\u03b8 (2) All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000071_iecon.2004.1433424-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000071_iecon.2004.1433424-Figure2-1.png", "caption": "Fig. 2. Typical turbine power relationship for various wind speed.", "texts": [ " It is clear h m the figure that there is a value of A for which Cp is maximized, thus maximizing the power for a given wind speed. Every wind speed has a variable turbine speed value that gives a maxi\" output power as shown in Fig. 3 for various wind speeds. As seen in the figure the peak power for each wind speed occurs at the point where Cp is maximized. To maximize the power generated from the wind turbine, it is therefore daimble for the generator to have a power characteristic that will follow the maximum C, line in Fig. 2. Then, the available optimum values at every turbine speed are the output turbine power data. 111. SruDIED SYSYTEM MODELING As shown in Fig.3, the utility grid with the ACIDCIAC converter i s utilized for the purpose of both excitation control of induction generator and energy storage. The equivalent circuit of this system is divided into two parts. The fmt is the equivalent circuit of the AC/DC/AC converter-grid combination referred to the induction generator. The second is the equivalent circuit of the induction generator" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002088_978-3-540-30738-9_14-Figure14.2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002088_978-3-540-30738-9_14-Figure14.2-1.png", "caption": "Fig. 14.2 Treadwheel-type boom derrick placed on erected building\u2019s wall", "texts": [ "1 shows a derrick commonly used in both ancient Greece and Rome, with a mast in the form of the letter \u201cA\u201d. The hoisted block is grabbed by a crampon. The winch is turned by means of two levers and the mast\u2019s inclination is adjusted by guy-ropes. As late as the second half of the 19th century, i. e., until the industrial revolution, cranes were driven by people or, less commonly, by animals. Treadmills, including treadwheels with steps to be climbed, or pole windlasses were used as the driving devices [14.1] (Fig. 14.2). Historically, the next class of construction machines to be invented was pile-drivers. These were needed to build structures (including temples) on weak ground and bridges. The first pile-drivers were very similar in their design to cranes. A simple medieval pile-driver consisted of a tripod, a rope sheave, and a rope with a hook on which a heavy stone block was hung. The dropped block would drive a pile into the ground. Devices working on the double-arm lever principle and a hoisting winch were employed to pull out such piles" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003491_1.3456118-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003491_1.3456118-Figure2-1.png", "caption": "Fig. 2. Diagram for finding point P\u2019s acceleration as it passes into and out of contact with the surface.", "texts": [ " Thus, in addition to the customary r F terms, its time derivative has the additional phantom torque term, ph = \u2212 i mi ri \u2212 rS aS = \u2212 M rcm \u2212 rS aS, 14 where aS is the acceleration of point S. In the previous example, we denoted S by P. We now find the phantom torque for the rocking hoop. As the semicircular hoop rocks, the trajectory of point P, which is in contact with the horizontal surface in Fig. 1, can be described in terms of its displacement from the point of contact. We use the quantities exhibited in Fig. 2 to describe this displacement. The fixed angle specifies the angular displacement of the line intersecting the contact point from the line that intersects the center of mass C. The angle specifies the angular displacement of the center of mass from the vertical. We use these definitions and take into account the horizontal motion of point O as the hoop rocks to obtain the horizontal and vertical components of the displacement, velocity, and acceleration of P from its point of contact, xP = R \u2212 \u2212 sin \u2212 , 15 yP = R 1 \u2212 cos \u2212 , 16 x\u0307P = R cos \u2212 \u2212 1 \u0307 , 17 \u02d9 \u02d9 yP = \u2212 R sin \u2212 , 18 906L" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000413_robot.2003.1242208-Figure12-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000413_robot.2003.1242208-Figure12-1.png", "caption": "Fig. 12: Clearance between mandrel and roller", "texts": [ " In Figure 11, F,y is plotted for the height of the flange during one turn (No. lo), where the phase of the 0 axis was mutually shifted by rad a5 the range sensor was located at the opposite side of the roller. The unevenness of the flange and the fluctuation of the feeding force are related almost linearly. Therefore, the growth of the wrinkles can be detected from the feeding force. 4 Clearance between mandrel and roller Among the various forming parameters in metal spinning, the clearance between the mandrel and roller (Fig. 12) can be considered the most difficult to set, in view of controlling the forniing machinery. In shear spinning, the wall thickness after the process is represented as Eq. (l), and the clearance should be exactly controlled equal to the thickness. When the clearance is too large, the precision is degraded because the material does not fully contact the mandrel. This upsets stable forming and wrinkles are likely to occur. Conversely. too small a clearance makes the forming force very large, and the flow of the material sometimes improperly deforms the product" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003133_aero.2009.4839606-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003133_aero.2009.4839606-Figure3-1.png", "caption": "Figure 3 \u2013 Geometry of the guidance law", "texts": [ " Moreover, the flight of the vehicle has to be performed under the requirements of a maximum rate of turn and with a constant airspeed, in order to avoid loss of lift. The latter requirement is fulfilled by considering a guidance law that acts only on the lateral acceleration, i.e. the acceleration perpendicular to the vehicle velocity respect to air, while leaving the longitudinal acceleration at zero. Now, by defining RD as the minimum radius of a circle centered over the target position and such that the pursuertarget distance cannot be lower, it results (see Figure 3) sin D D R if R R R \u03d5 = \u2265 (12) where R is the target-pursuer distance; \u03d5 is the angle between the target-pursuer line of sight and the direction of the tangent to the circle with radius RD taken from the pursuer position. If the pursuer is inside the circle, DR R< and expression (12) is not defined. In this case \u03d5 is arbitrarily set to 2 \u03c0 , the value it has over the circumference. Summarizing, it holds: asin 2 D D D R R R R R R \u03d5 \u03c0\u03d5 \u23a7 \u239b \u239e= \u2265\u239c \u239f\u23aa\u23aa \u239d \u23a0\u23a8 \u23aa = <\u23aa\u23a9 (13) The aim of the guidance law is to force the pursuer to steer toward the closest tangent-to-circle direction, that can be \u03bb \u03d5\u2212 or \u03bb \u03d5+ (14) or, depending on the direction of the pursuer velocity vector" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000192_02678290412331282109-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000192_02678290412331282109-Figure1-1.png", "caption": "Figure 1. (a) Unperturbed fibre with radius a is aligned in the z-axis of a cylindrical coordinate system (r, h, z). Cross-sectional view in Cartesian coordinates (x, y) shows the unit vectors ir and ih and azimuthal angle h. (b) Periodically deformed axisymmetric fibre with unit surface normal vector N, radius R and wavelength l. Fibre radius R and unit surface normal N periodically change with wavelength l along the z-direction. The figure is representative of the peristaltic mode.", "texts": [ " Geometry and texture of nematic liquid crystalline fibres To define the state of a nematic liquid crystalline fibre completely, both the geometry of the fibre and the spatial orientational order of the nematic liquid crystal must be specified. More specifically we define: Nematic Liquid Crystalline Fibrew{n, R, N}, where n is the nematic director field [22], R is the fibre radius, and N is the unit surface normal vector. For an isotropic material fibre, only the geometry is necessary, i.e. {R, N}. Figure 1 shows definitions of the fibre geometry. Figure 1 (a) shows that the fibre is initially a uniform cylinder with radius a, with its axis collinear with the z-axis of a cylindrical coordinate system. In the crosssectional view, unit vectors ir and ih are shown in the direction of the r- and h-axes, respectively. Figure 1 (b) shows the periodically deformed fibre with unit surface normal N, radius R and wavelength l. The fiber radius R and unit surface normal N periodically change with a wavelength l along the z-direction. In this paper, we apply our analysis to three characteristic nematic textures of initially constant director fields, denoted as axial, onion, and radial textures, and accordingly the nematic fibre with each texture is called axial fibre, onion fibre, and radial fibre, respectively. Figure 2 shows the schematic of undeformed fibres with (a) axial, (b) onion and (c) radial textures", " Likewise, chiral non-axisymmetric distortion modes are not taken into account in this paper. 2.2. Static energy analysis We consider the thermodynamic stability of an infinitely long cylindrical nematic LC fibre subjected to infinitesimal periodic surface disturbances. The nematic liquid crystal is assumed to be incompressible, and its initial orientation is homogeneous in each of three characteristic textures: axial, onion and radial. Then, the director fields evolve as the shape of nematic fibres changes. As seen in figure 1 (b), the fiber radius R and the unit surface normal N change along the z-direction. In the static analysis, the fibre shape with a periodic surface disturbance is given at any position z by R z\u00f0 \u00de~R0zj z\u00f0 \u00de: \u00f01\u00de The periodic surface disturbance j is expressed as j z\u00f0 \u00de~j0 cos kz\u00bd \u00f02\u00de where j0 is the initial amplitude of the disturbance and k the axial wavenumber. The average fibre radius R0 in equation (1) is required, for a fixed volume, to be R0~a 1{ j2 0 2a2 !1 2 &a 1{ j2 0 4a2 ! \u00f03\u00de where the approximation is valid in the linear regime of the capillary instability, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003059_cdc.2010.5717728-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003059_cdc.2010.5717728-Figure1-1.png", "caption": "Fig. 1. Two-link revolute manipulator", "texts": [ " Due to the adaptive nature of the protocol (24), knowledge of vehicle dynamic model is not required. This is a highly desirable feature in practical applications where some dynamics effects such as friction can not be modeled accurately. As an example of networked Lagrangian systems, a group of robot manipulators is considered here. For ease of plotting, we simulate a multi-manipulator systems composed of six two-link revolute manipulators. The profile of the two-link planar manipulator is shown in Fig. 1. Fig. 2 shows the communication topology among the manipulators, which has a spanning tree. In what follows, all the functions are defined componentwise for a vector. According to the physical properties of the manipulator [14], the input signals of the NN for manipulator i are taken as xi = (sin(qi) T ,cos(qi) T , q\u0307T i ,(\u2211 j\u2208Ni ai jq\u0308 j + biq\u03080) T ,(\u2211 j\u2208Ni ai jq\u0307 j + biq\u03070) T )T in simulation. In the case that the variations of velocities are slow, the acceleration information, which is used as input signals of NN, can be indirectly calculated by numerical differentiation of the velocities" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001669_1-4020-2204-2_10-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001669_1-4020-2204-2_10-Figure8-1.png", "caption": "Fig. 8: Manipulating the controller to (a) nod the head; (b) incline the head; (c) turn the head; (d) bow the body; (e) keep the head upright; (f) walk the puppet.", "texts": [ " For example, Figures 8(a), 8(b) and 8(c) show how to nod, incline, and turn the head respectively by maneuvering the controller [8]. Such kind of manoeuvring usually produces fine puppet movement because string lengths are unchanged. Figures 8(d), 8(e), 8(f) illustrate how to use the combined hand and controller motion to make the puppet bow [8], lift his head, and walk. To make large and prominent gestures, such as walking, plucking the strings or pulling the detachable horizontal bars is necessary (Fig. 8f). Usually one puppeteer can only operate one marionette from the controller unless the design of the controller can cater for multiple puppet figures. The puppeteer uses one hand to hold and maneuver the controller and the other to pluck or pull the strings and the detachable bars. With an ergonomically designed paddle controller (Fig. 5d), manoeuvring the controller and plucking the strings can be achieved using one hand simultaneously [10]. Manipulating the marionette using a Chinese Gou Pai is similar" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001529_15567260701715396-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001529_15567260701715396-Figure1-1.png", "caption": "Figure 1. The physical model.", "texts": [ " NOMENCLATURE hst latent heat of fusion, J kg 1 hiv latent heat of vaporization, J kg 1 k thermal conductivity, W m 1 C 1 q\u02dd heat flux, W m 2 Rg gas constant, J/kg K ss solid-liquid interface location, m sl liquid-vapor interface location, m t time, s tm time at which melting begins, s tp half width of the laser beam pulse at q\u02dd0/2, s tr time at which melting and resolidification begins (when vaporization ends) ts time at which solidification ends, s tv time at which vaporization begins, s x coordinate measured from the particle surface, m Greek Symbols thermal diffusivity, m2 s 1 thermal penetration depth, m s thickness of thermal layer at the time at which solidification ends, m g specific heat ratio (T \u2013 Ti), relative temperature, C density, kg m 3 relative time, t/tp Subscripts i initial l liquid phase m melting point s solid phase sat saturation In this article, preheating, melting, vaporization, melting and resolidification, and thermalization of a single powder particle subjected to a nanosecond pulsed laser heating will be modeled. Temporal Gaussian heat flux of a laser beam is considered and the origin of time is chosen to be at the time when the heat flux is at its maximum (see Figure 1). q00 t\u00f0 \u00de \u00bc q000e ln2t2 t2p \u00f01\u00de where q\u02dd0 is the maximum heat flux, and tp is the half-width of the laser pulse at half maximum. The diameter of the metal powder particle is much smaller than the diameter of the laser beam, which is in turn much smaller than the dimension of the final part. Since the laser radiation penetrates the powder bed over a distance of several powdersphere diameters, it can be assumed that multiple scattering of the radiation leads to a nearly homogeneous distribution of the heat flux within the optically penetrated layer [13]; this leads to an almost normal incidence of the radiation on the surfaces of the grains in the underlying layers. Figure 1 shows the physical model of solid-liquidvapor phase change process under consideration. Because of symmetry of the spherical particle as well as the assumption of uniform heat flux distribution around the particle, the problem can be assumed to be 1-D in the r-direction. The heat flux at the particle surface, q\u02dd (W/m2) can be related to the total energy flux by one pulse, or the laser fluence J (J/m2) by: J \u00bc Z1 1 q00 t\u00f0 \u00dedt \u00bc Z1 1 q000e ln 2t 2 t2pdt \u00bc 2q000tp ffiffiffi pffiffiffiffiffiffiffi ln 2 p \u00f02\u00de Providing sufficient pulse energy, a single power particle in the SLS process may undergo five stages: (1) preheating, (2) melting, (3) vaporization, (4) melting and resolidification, and (5) thermalization" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002239_j.euromechsol.2008.06.008-Figure18-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002239_j.euromechsol.2008.06.008-Figure18-1.png", "caption": "Fig. 18. Results obtained for Problem 5 with stability constraint (Nm = 3, N p = 7, Na = 6). (a) The end-effector\u2019s goal orientation is imposed. (b) The end-effector\u2019s goal orientation is free.", "texts": [ " Note also that calculations now are carried out using Na = 6 (instead of 5) free nodes for the path of the manipulator. Two distinct versions (A and B) of this problem are considered. In version A, the finale state is defined completely: U G e = (1.222,7.071,0.65,135\u25e6,0\u25e6,0\u25e6) while in version B it is only partly defined in the sense that the end-effector\u2019s final orientation is left unspecified: U G e = (1.222,7.071,0.65, free,0\u25e6,0\u25e6). The solutions calculated by RPA for both versions, while considering stability, are shown in Fig. 18. The runtimes are approximately the same (187 seconds). In both cases, the ZMP is kept within the BSR (not shown). Comparing the two runs, we clearly notice the significant difference in the platform paths and the opposite unfolding of the arms. As the end-effector\u2019s goal orientation is not imposed in version B, the RPA was able to take advantage of this extra freedom by choosing a faster path for the platform but also by determining the proper unfolding of the arms to compensate the platform inertia forces" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001916_j.fusengdes.2008.05.042-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001916_j.fusengdes.2008.05.042-Figure6-1.png", "caption": "Fig. 6. Detection of interference. The parts that interfere \u2013 module and gripper \u2013 change those colors.", "texts": [ " In this case, the simulator can provide such infor- ation by making the next module transparent on the screen, as hown in Fig. 5. .4. Testing a taught sequence (off-line) The simulator can also be used for testing a taught sequence efore actual operation. Using the teaching points in the sequence s inputs to the control system, the 3D model\u2019s motion can be eviewed to confirm the interaction between the manipulator nd other components. If the manipulator interferes with any omponents, the simulator will give an alarm showing the relvant parts in red, as shown in Fig. 6. The simulator can also how the distance between components that potentially interfere ith each other, as shown in Fig. 7. These interference-related unctions are also useful for real-time viewing during actual opertion. .5. Operator training (off-line) Because there are many joints, the operation of the manipulaor is complex, and operator training is important. The simulator an be operated by the same controlling device as the manip- 1840 N. Takeda et al. / Fusion Engineering an F s u f w t i w a m 4 c t c b m t i n o p b l t o 5 k T t d m f v b I m o fi k c A C t m m t c R [ [ [ [ [ [ ig" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000844_iemdc.2005.195712-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000844_iemdc.2005.195712-Figure1-1.png", "caption": "Fig. 1. 6-slot, 4-pole brushless permanent magnet motor.", "texts": [ " \u2022 By reducing harmonics of order n C Ni s= , (n=1, 2, 3\u2026) in the cogging torque produced by a single slot, the resultant cogging torque can be reduced significantly. The resultant cogging torque waveform can be synthesised from the cogging torque waveform which is associated with a single stator slot, either numerically or analytically. However, analytical synthesis provides much more insight into the nature of the cogging torque waveform. Hence, an analytical method is used in the following analysis. III. 6-SLOT, 4-POLE MOTOR In this section, a 6-slot 4-pole surface-mounted magnet motor, Fig. 1, is considered to illustrate the influence of the slot number on the waveform and harmonic spectrum of the resultant cogging torque. The outer stator diameter and axial length are 140mm and 40.6mm, respectively, while the airgap length and magnet thickness are 1.5mm and 3mm, respectively. The width of the slot openings is 2mm. The magnet pole-arc/pole-pitch ratio is 1.0, and the magnets are parallel magnetized NdFeB having a remanence of 1.2T and a relative recoil permeability of 1.05. The cogging torque waveforms and harmonic spectra which result when the stator has only 1 slot and 2, 3, 4, 5 and 6 uniformly distributed slots are analysed, and analytically synthesised and finite element predicted results are compared with measurements" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003147_ijmee.38.2.5-Figure8-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003147_ijmee.38.2.5-Figure8-1.png", "caption": "Fig. 8 Assembly drawing of the compressed-air engine.", "texts": [], "surrounding_texts": [ "Inductive learning begins with concrete experience, observations or a question, and then develops knowledge, skills and theory from that basis [1]. Research has demonstrated it improves on the more traditional, deductive learning methods, in that inductive learning promotes deep knowledge structures, critical thinking and intellectual development [1]. Industrial employers have often called for mechanical engineers who have handson skills and integrated knowledge [2, 3]. These skills and knowledge are enhanced by project activities that span theory, design and construction [4, 5], as well as direct experience with actual machine models [6]. Integrated computer-aided design/manufacture (CAD/CAM) activities are important in learning the principles of design for manufacture and concurrent engineering [7], and a practical understanding of machining operations enhances an engineer\u2019s ability to design cost-effi cient machined components. This paper examines how the Machine Shop Practices and Solid Modeling course (MENG 351) has been integrated with other concurrent and future courses, utilizing an inductive and active learning model [1, 8]. The literature indicates that the integration of hands-on machining/fabrication, design, experimentation, analysis and theory may result in less compartmentalization of knowledge, greater student enthusiasm and deeper learning of concepts. Integration of MENG 351 occurs across a number of courses, including Systems Laboratory, Mechanics of Materials, Machine Design, and Capstone Design, as well as within the course via design\u2013build\u2013test activities. at USD & Wegner Health Science Information Center on April 9, 2015ijj.sagepub.comDownloaded from International Journal of Mechanical Engineering Education 38/2 Projects were carefully chosen to achieve the learning objectives of MENG 351 and to interface with future courses in the inductive learning process. The shop portion of MENG 351 is aimed at developing skills in woodworking, manual machining and sheet-metal fabrication. In a later course (Manufacturing Processes), students develop computer numerical control (CNC) and welding skills. Students worked in teams of two for almost all projects. In the shop, this buddy-system arrangement helped ensure students were attentive to each other\u2019s safety; no signifi cant injuries occurred throughout the course. Woodworking projects: fast-return actuator and acoustic guitar As their introductory project to woodworking equipment, students constructed a simple fast-return actuator mechanism (Fig. 1). This actuator (an inversion of the slider-crank mechanism) was also analyzed in the concurrent Dynamics class. This project taught skills on the miter saw, table saw, drill press, sander and band-saw. The basic design was adapted and modifi ed from Levy [9]. Mechanical engineering students sometimes have pre-existing skills in woodworking; this project was designed to allow both basic and advanced versions, to provide challenge to all levels. This project typically took one lab period. For their second woodworking project, students designed and built simple acoustic guitars (Fig. 2). This project interfaces well with a Vibrations course, incorporating vibrating strings, resonance and acoustic coupling [10]. A schematic neck-and-box design was provided; most students enjoyed modifying the basic design towards styling or greater size. The soundboard was reinforced on its backside. Instrumentquality guitars are made from expensive tone-woods such as spruce, cedar, mahogany and maple, but adequate resonance properties can be obtained from less expensive materials: 1/8\u2033 Baltic-birch plywood was specifi ed for the soundboard and back of the body, 1/4\u2033 poplar for the body sides and 3/4\u2033 poplar for the neck. Strings were made from nylon fi shing line (30\u201380 lb test) and small eyebolts func- at USD & Wegner Health Science Information Center on April 9, 2015ijj.sagepub.comDownloaded from International Journal of Mechanical Engineering Education 38/2 tioned as tuners. The guitar project developed similar skills as the fast-return actuator project, but required greater precision. This fi rst generation of the guitar project took two lab periods, and generated student enthusiasm and creativity. In a second iteration of the course, the guitar design was greatly refi ned in aesthetics, playability and tonal quality (Fig. 3), incorporating frets, steel strings, spruce soundboard bracing in a Martin X-layout, and a laser-cut mandolin body shape. This guitar design cost approximately $US20, took 3.5 lab periods to build and delivered sound quality similar to many mass-production guitars. Refi nements in-process include a standard two-bout guitar-shaped body. Not surprisingly, the new design generated even more student enthusiasm. Sheet-metal project: reinforced hollow beam To practice sheet-metal forming methods, and to provide an experiential basis for beam theory in Mechanics of Materials, pairs of students designed and fabricated hollow aluminum beams with various cross-sections and bulkhead designs [11]. The at USD & Wegner Health Science Information Center on April 9, 2015ijj.sagepub.comDownloaded from International Journal of Mechanical Engineering Education 38/2 students learned to use the following machines for this project: 52\u2033 stomp shear (Pexto), 6\u2033 corner notcher (Enco), 40\u2033 box/pan brake (Grizzly G0578), hand drill, and manual and pneumatic pop-rivet guns. The students were constrained to making the main beam cross-section from an 11\u2033 \u00d7 16\u2033 sheet of .032\u2033-thick 5052-H32 aluminum. Bulkheads were made from additional material. Students were encouraged to try different cross-sections: rectangular, U, I and triangle sections (Fig. 4). Most teams chose a 2\u2033 \u00d7 3\u2033 rectangular or I-section. To prevent collapse of the crosssections, bulkheads were specifi ed at each of the three load application points, though the detailed design of each bulkhead was up to the students. Beam construction generally took one lab period. All beams were tested in a three-point bending test fi xture of 15\u2033 span; students recorded the maximum load and mode of failure, and the test specimens were saved for later use in Mechanics of Materials. at USD & Wegner Health Science Information Center on April 9, 2015ijj.sagepub.comDownloaded from International Journal of Mechanical Engineering Education 38/2 The students\u2019 beams held between 405 and 1188 lb, substantially below their theoretical capacities. For the rectangular-section beams, failure was typically local buckling of the thin aluminum sheet (Fig. 5). Specifying a thicker material or smaller cross-section would help prevent premature failure from local instability. This buckling clearly illustrates the compressive stresses acting on the upper fl ange. For the I-section beams, failure was typically lateral-torsional buckling (Fig. 6). Seeing both of these non-ideal failure mechanisms fi rst hand helps prepare students for the complexities of real-world design. The beam design, construction and testing directly related to Mechanics of Materials, where students calculated moment of inertia, shear forces on the rivets and theoretical maximum load. Students could then better understand their design and failure mechanisms. The construction and testing of the beams provided a direct experiential basis for the development of beam theory, following inductive learning principles. Machining project: compressed-air engine The primary project of MENG 351 is the machining, assembly and testing of small one-cylinder compressed-air engines, building on similar work at other institutions [12\u201316]. In addition to teaching machining skills (Fig. 7), this project initiated inductive learning pathways through multiple courses: Machine Design, Manufacturing at USD & Wegner Health Science Information Center on April 9, 2015ijj.sagepub.comDownloaded from International Journal of Mechanical Engineering Education 38/2 at USD & Wegner Health Science Information Center on April 9, 2015ijj.sagepub.comDownloaded from International Journal of Mechanical Engineering Education 38/2 Processes, Capstone Design, and others. The engine design was based on previous simple oscillating air/steam engines [e.g. 17] with updates such as a bronze bearing for the crank journal. The students followed dimensioned drawings developed by the instructor, incorporating some geometric dimensioning and tolerancing (GD&T) notations (Figs 8 and 9). Students were encouraged to modify this design for more power or better vibration characteristics. Students learned to use the following machine tools for this project: 12\u2033 \u00d7 36\u2033 engine lathe with digital read-out (DRO) (Birmingham, Sony), manual milling machine with DRO (Lagun, Sony), horizontal bandsaw (jet), tool grinder, drill press, dial caliper, and vernier micrometer. Before starting machining operations, all students developed operation sheets for each part, which were reviewed by the instructors. On the lathe, almost all students successfully achieved the +/\u22120.0005\u2033 tolerance for their piston diameter. Additional practiced skills included print-reading, use of machinist tables, press-fi tting, tapping, assembly, and shop professionalism. A sample student engine is shown in Fig. 10. The importance of holding tolerances becomes clear to many students during the assembly phase. For example, holes intended for a press-fi t often required rework due to a reaming operation that cut oversize. Or, the smoothness of running was at USD & Wegner Health Science Information Center on April 9, 2015ijj.sagepub.comDownloaded from International Journal of Mechanical Engineering Education 38/2 at USD & Wegner Health Science Information Center on April 9, 2015ijj.sagepub.comDownloaded from International Journal of Mechanical Engineering Education 38/2 occasionally hampered by a non-perpendicular cylinder bore. Students then learned various techniques for overcoming such diffi culties and, in the end, all engines ran well. Engines were tested on a simple Prony-brake dynamometer (Fig. 11) for minimum pressure required to run (typically 1\u20132 psi), minimum speed at that minimum pressure (200\u2013300 RPM), speed at 30 psi (1200\u20133000 RPM), and shaft output torque and rpm at 30 psi, from which the students calculated power (3\u20136 watts). The kinematics of the inverted slider\u2013crank mechanism are analyzed in the Dynamics as well as Machine Design courses. Power calculations build on concepts in Physics and Dynamics. Power output and valve port fl ow were directly related to theoretical concepts in Thermodynamics and Fluid Mechanics. Numerous manufacturing methods were drawn upon the following semester in Manufacturing Processes and the following year in Capstone Design. Print-reading and operations planning skills were taught and developed for each component of the engine, and later exercised in Capstone Design. The engine design\u2013build\u2013test project developed precision machining skills and formed the basis for multiple inductive learning processes. Systems laboratory design project: catapult In our Mechanical Engineering Systems Laboratory, teams of three or four students built a catapult to serve as a Taguchi design-of-experiments project [18]. The at USD & Wegner Health Science Information Center on April 9, 2015ijj.sagepub.comDownloaded from International Journal of Mechanical Engineering Education 38/2 catapults were of the students\u2019 own design, but had to allow variation of three parameters in order to maximize the distance, accuracy and precision of each launch. Further, the catapults had to fi t within an 18\u2033 cube. The teams built Pro/Engineer models of their design and then constructed and tested their catapults from their drawings (Fig. 12). Teams selected tension springs and torsion bars for the energystorage device. Instructors found that combining the efforts of their courses around one project resulted in a more intensive, integrated and effective learning experience." ] }, { "image_filename": "designv11_20_0003961_detc2013-12361-Figure9-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003961_detc2013-12361-Figure9-1.png", "caption": "FIGURE 9. Initial Orientations (\u03b8c = 0) a) UMD b) OSST", "texts": [ " All measurement discussed were conducted at full mast loading. Tachometers were used on both the input and output shafts. The output tachometer signal is used to partition the measured data into individual carrier cycles as well as count them to produce a reset pulse indicating when the planetary geometry has returned to its initial orientation. The planetary system has a repeat cycle of 105 carrier cycles. The initial planetary orientation, number of accelerometers, and the accelerometer locations differed between the two test programs as shown in Figure 9. UMD: Accelerometer A1 was mounted to the input end of the transmission just above the pinion. The input spiral bevel pinion has a thrust load against its tri-plex ball bearing and single roller bearing. This is the direction of A1\u2019s sensitivity axis and is 8 Copyright \u00a9 2013 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/77583/ on 03/04/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use aligned to the rotational axis of the input shaft [35]" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003263_s12204-011-1104-9-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003263_s12204-011-1104-9-Figure1-1.png", "caption": "Fig. 1 Sketch of forbidden vector", "texts": [ " In TA logic, some problems of the propulsion system require to be taken into consideration, such as the maneuverability and hydrodynamic interaction with current or hull of vessels. Actually thrust loss due to thruster-thruster interaction is proved to be very terrible, especially when two adjacent thrusters are parallel aligned in the extreme sea conditions. Consequently, aiming to reduce thrust loss, one boundary in optimal TA logic named forbidden vector was builded and proved to be feasible in the past decades. In this article, we will give out a math model to depict the forbidden vector, as shown in Fig. 1. The sketched forbidden angle is 15\u25e6. We take No. 5 thruster in Fig. 1 as an example, and suppose that the forbidden angle is 15\u25e6. When both the azimuth angles of No. 5 and No. 1 thrusters are \u221226\u25e6, the two thrusters in the Fig. 1 are aligned in a line, so the azimuth angle domain of \u221241\u25e6 to \u221211\u25e6 is forbidden. All azimuth angles are in the interval of [0\u25e6, 360\u25e6]. As a results, the azimuth angle \u03b1 of No. 5 thruster has to satisfy two equations, as follows: \u03b1 \u03b1u \u03b1 \u03b1l } , (2) where \u03b1u is the upper limit of forbidden vector, and \u03b1l is the lower limit. Consequently we give out one inequality instead of Eq. (2), g(\u03b1) = ( \u03b1 \u2212 \u03b1l + \u03b1u 2 )2 \u2212 ( \u03b1u \u2212 \u03b1l 2 )2 0. (3) In Fig. 1, \u03b1u equals to 349\u25e6, \u03b1l equals to 319\u25e6, F = [Fx Fy] is the sea load vector of platform, and Mz is the moment of force due to sea load. 1.4.1 Definition of Pseudo-Inverse In the linear algebra, the pseudo-inverse A+ of an m \u00d7 n matrix A is a generalization of the inverse matrix[7]. More precisely, this article is about the Moore-Penrose pseudo-inverse, which was independently described by Refs. [8-9]. Earlier, Fredholm[10] had introduced the concept of a pseudo-inverse of integral operators in 1903" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000230_bf02843970-Figure1-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000230_bf02843970-Figure1-1.png", "caption": "Figure 1 Top and rear (section A\u2013A) views of the basketball rolling on the rim. A moving orthogonal right-handed co-ordinate frame A has unit vectors a1 and a2 in a vertical plane. a2 is in the direction from the rim contact point to the basketball centre and the third unit vector a3 = a1 x a2 is horizontal. Two angles \u03b8 and \u03c6, and one displacement r describe the basketball position on the rim.", "texts": [ " In addition, the radial compliance of the basketball must be taken into account to model accurately the radial deflection and bouncing that can occur. The goal is to make a model of ball contact with the rim, without constraint to the sagittal plane, and including ball deflection, with general initial conditions. Preliminary versions of this model have been presented by Okubo & Hubbard (2002; 2003). In this paper, we use the model to study shot sensitivity to initial conditions and specifically to investigate initial conditions required for long-rolling trajectories. Fig. 1 shows the co-ordinate system. The intersection of the horizontal plane of symmetry and the axis of symmetry of the toroidal basketball hoop of radius Rh is labelled O. We assume that the basketball with radial compliance 1/k maintains contact with the rim at a variable point B\u0302 on the ball. Although general basketball motion also includes phases of gravitational flight and rolling contact without slipping, we first restrict ourselves to the most general case in which the velocities of the two contact points on the ball and rim are different, i", " where r is the distance between B* and B\u0302 , and Rr is the radius of the rim. Unit vectors a1 and a2 of a righthanded orthogonal reference frame A lie in the plane determined by O, B* and B\u0302 , with a1 perpendicular to the line through B* and B\u0302 and directed toward the outside of the hoop when the tilt angle into the hoop, \u03c6, is zero. The unit vector a2 points from B\u0302 to B*, is normal to the a1a3 contact plane, and is vertical when \u03c6 is zero. The roll-around angle \u03b8 is the angle from the inertial X-axis to Section A-A (See Fig. 1). In later problems it may be convenient to choose the X-axis to be parallel to the backboard plane but in this paper it is taken to be perpendicular to the vertical shot plane. The configuration of the basketball with radial compliance has six degrees of freedom. Three determine the position of the ball centre; \u03b8, \u03c6 and r. The other three describe ball orientation. Although spin is important, the orientations themselves are neglected here since the ball surface is assumed to be uniform. Define three angular velocities, \u03c91, \u03c92 and \u03c92, to be the components of the inertial angular velocity N B of the basketball body-fixed frame B in the a1, a2 and a3 directions, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002711_00124278-200702000-00003-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002711_00124278-200702000-00003-Figure2-1.png", "caption": "FIGURE 2. Mechanical structure of 3d-ISMS.", "texts": [ " The 3d-ISMS device included a signal processor, 12-bit analog-to-digital/digital-to-analog card, and 3 displacement transducers (Gefran Spa, Italy), one of which was a rectilinear displacement transducer, and the others were ballpoint rectilinear displacement transducers. A 3d-ISMS device was attached to the end part of the Olympic-style barbell by a cylindrical roller bearing and measured three dimensional displacement data of whole motion relative to the ground within an accuracy of 0.01 cm (Figure 2). Customized software was used to record the displacement-time data sampled at 100 Hz. Calibration of 3d-ISMS was checked before all testing sessions while it was attached to the Olympic-style barbell positioned statically on the bench. The mean displacement data for all 3 dimensions were found under 0.005 cm in the static position. Before the power and linear momentum were calculated, the barbell displacement for each trial was first filtered using a low-pass filter with a cut-off frequency of 8 Hz and then differentiated by finite difference algorithm in MatLab software (MathWorks Inc" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002690_978-3-540-30301-5_19-Figure18.7-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002690_978-3-540-30301-5_19-Figure18.7-1.png", "caption": "Fig. 18.7a\u2013c An experimental microassembly workcell", "texts": [ " Therefore, it is often important to consider environment control in system design. Assembly Process Flow: An Example The micromachined metal parts (Fig. 18.6c) are horizontally transferred to the workcell on the vacuum-release tray (Fig. 18.6b). The wafer is placed on a wafer mount perpendicular to the horizontal plane (Fig. 18.6b). This configuration does not require the flipping of the thin metal parts and is advantageous in terms of reliability and efficiency. There are two major operations in each assembly cycle: pickup and insertion. Under the configuration shown in Fig. 18.7a, all the operations are performed by the same workcell. Complex packaging operations must often be decomposed and performed by multiple workcells [18.76]. General System Architecture An automated microassembly system typically consists of the following functional units. Part B 1 8 .6 Large workspace positioning unit. A large workspace and long-range positioning motion are required in most microassembly operations. The large workspace is necessary to accommodate different functional units, part feeders, and various tools. In general, off-the-shelf motion control systems developed for automated IC packaging equipment can be adopted directly. For the task described in Sect. 18.6.1, the DRIE-etched holes are distributed on wafers up to 8 inches in diameter. This requires the assembly system to have a commensurate working space and high positioning speed. The coarse positioning unit has four DOFs (Fig. 18.7a). Planar motion in the horizontal direction is provided by an open-frame high-precision XY table with a travel of 32 cm (12 inch) and a repeatability of 1 \u03bcm in both directions. Position feedback with a resolution of 0.1 \u03bcm is provided by two linear encoders. A dual-loop PID plus feedforward control scheme is used for each axis. The internal speed loop is closed on the rotary encoder on the motor. The external position loop is closed on the linear encoder. Each wafer is placed on the vertical wafer mount that provides both linear and rotational control (Fig. 18.7b). Vertical motions of the wafer mount are provided by a linear slide with a travel of 20 cm (8 inch) and a repeatability of 5 \u03bcm. It is also controlled using a PID plus feedforward algorithm. Both the XY table and the vertical linear slides are actuated using alternating-current (AC) servo motors. The rotation of the wafer mount is actuated by an Oriental PK545AUA microstep motor with a maximum resolution of 0.0028\u25e6/step. All low-level controllers are commanded and coordinated by a host computer [18", " In general, the development of IC manufacturing towards the deep submicron level provides a major driving force behind the development of these motion control techniques. Micromanipulator unit. Fine pose (position and orientation) control is required in operations such as 3-D precision alignment and assembly. For example, the six DOFs required by the task introduced in Sect. 18.6.1 are implemented on separate structures. The three Cartesian DOF are provided by an adapted Sutter MP285 micromanipulator, which also provides yaw motions with its rotational DOF (Fig. 18.7c). Roll motions are implemented on the wafer mount (Fig. 18.7b). The pitch movement of the metal part after pickup is not motorized and is implemented through manual adjustment and calibration before assembly. More discussion of micromanipulator configuration can be found in [18.61]. Two principles need to be considered in implementing motion control for automated 3-D microassembly. The first is the partition of large-workspace coarse positioning unit and the fine positioning micromanipulator unit. The second is the decomposition and distributed implementation of multiple DOFs", " Microgripper development is often closely related to the development of fixtures that can also have microscale sizes. Microscope optics and imaging unit. The function of the microscope optics and imaging unit is to provide noncontact measurement of the geometry, motion, and spatial relations of assembly objects. Typical configurations of commercial device bonding systems use one or two vertical microscopes. An inverted microscope configuration is commonly used for backside alignment. On the other hand, 3-D microassembly may require two camera views in a stereo configuration. In the system shown in Fig. 18.7a, a total of four different views can be provided to its human operator: a global view of the entire assembly scene, a vertical microscopic view for part pickup, and two lateral microscopic views for the fine position and orientation adjustments during the final microassembly operations. Each view uses a CCD camera with a matching optical system. All images are captured using a Matrox Corona PCI frame grabber. Microscopic visual feedback is crucial for precise 3-D alignment. Provided that resolution requirements of the assembly task can be satisfied, microscope optics with larger working distances is desirable" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003965_acc.2012.6315561-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003965_acc.2012.6315561-Figure3-1.png", "caption": "Fig. 3. Initial attitudes and their identical forbidden zones in 3D space (case 1)", "texts": [], "surrounding_texts": [ "In this section, we present simulation results for two cases where six spacecraft are required to align their attitudes in presence of attitude forbidden zones. These spacecraft are assumed to carry light sensitive interferometers with a fixed bore-sight along the z-body axis. It is further assumed that they share attitude information as well as auxiliary variables through a bidirectional communication network. In the first case, we consider the scenario where the attitude forbidden zones are identical over all six spacecraft. In the second scenario, we consider the case where each spacecraft has a distinct attitude forbidden zone. Note that in the second case, it is implicitly assumed that the intersection of all permissible zones is non-empty in order to be able to reach a consensus orientation. The communication network is given as a complete graph for case 1 and for case 2, five edges are randomly eliminated from the complete graph while keeping the graph connected. These graphs are shown in Fig. 2. All initial attitudes, shown in Table 1, are randomly 1In this paper, we use \u2018?\u2019 for denoting the Kronecker product in order to avoid confusion with the symbol for quaternion multiplication \u2018\u2297\u2019. selected but they satisfy the attitude permissible zone DPi for all i. Although it is generally not necessary that all initial angular velocities are zero, we assume that this condition holds for our simulation examples. Figs. 3 and 4 describe initial attitudes of the multiple spacecraft and the attitude forbidden zones in three dimensional space. Figs. 5 and 7 on the other hand, depict the unit quaternion trajectories for case 1 and case 2, respectively, while Figs. 6 and 8 depict the corresponding angular velocities over time. Note that all spacecraft converge to the same attitude regardless of the configuration of their respective attitude forbidden zones. Figs. 9 and 10 trace the pointing directions of the light sensitive instruments on the cylindrical projection of celestial sphere for case 1 and 2 respectively, where \u2018\u25e6\u2019 denotes initial orientations while \u2018\u00d7\u2019 denotes desired orientations. 0 10 20 30 40 50 60 \u22121 0 1 0 10 20 30 40 50 60 \u22121 0 1 0 10 20 30 40 50 60 \u22121 0 1 0 10 20 30 40 50 60 \u22121 0 1 S/C 1 S/C 2 S/C 3 S/C 4 S/C 5 S/C 6 Fig. 5. Attitude trajectories(unit quaternions) over time in case 1 0 10 20 30 40 50 60 \u22121 0 1 0 10 20 30 40 50 60 \u22120.5 0 0.5 1 0 10 20 30 40 50 60 \u22121 0 1 0 10 20 30 40 50 60 \u22121 0 1 S/C 1 S/C 2 S/C 3 S/C 4 S/C 5 S/C 6 Fig. 7. Attitude trajectories(unit quaternions) over time in case 2" ] }, { "image_filename": "designv11_20_0001964_tmag.2007.894332-Figure10-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001964_tmag.2007.894332-Figure10-1.png", "caption": "Fig. 10. Experimental setup. The yoke is stabilized by electromagnet, and its stationary position is adjusted by composite energized by a dc-dc converter.", "texts": [ " Our strategy is a combination of composite and electromagnet. In the system, the composite is used to adjust the gap at stationary, and thus does not require a high-power amplifier. Instead, the motion of levitated yoke is stabilized by magnetic force control by the electromagnet. The system will be, therefore, equipped with zero power characteristics of the composite confirmed at the foregoing experiments and high robustness by the electromagnet. We verified this feature by an experimental setup as shown in Fig. 10. The composite and an electromagnet aligned exerted magnetic forces on levitated yoke affixed on a linear slider (total weight: 300 g). The electromagnet was iron yoke (gap area of 10 8 mm ) with coil of 200 turn wound and permanent magnet of 10 8 1 mm . A gap sensor was placed between both devices. The relations between the gap and magnetic force with parameters of fixed voltage (0, 100, 200 V) to the PZTs and current to the coil ( A) is depicted in Fig. 11. The magnetic force increasing with the decrease of the gap is summation of that of two devices including permanent magnets" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000309_s10569-004-1508-z-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000309_s10569-004-1508-z-Figure6-1.png", "caption": "Figure 6. Failed attempt to design the periodic orbit directly.", "texts": [ " Note that the LL-8-1 reference orbit (see Table I) also can be used for obtaining the intermediate 260-day transfer shown in Figure 5. The following problem will be considered here as an example: find a planar periodic orbit simply rounding the Earth with period 5\u20136months and passing through the point r0 \u00bc r1 \u00bc f 0:2 106; 1:2 106; 0g \u00f019\u00de Let us take the orbital period equal to 165 days as a first guess and use LL-1-1 reference orbit (see Table I) for the problem solution; this reference orbit is to give two revolutions around the Earth. However, the problem cannot be solved directly, in one step: the transfer shown in Figure 6 with the transfer time 330 days is the direct problem solution using LL-1-1 reference orbit. In order to solve the problem a 300-day transfer can be found first between r0 given by (19) and r1 \u00bc f 0:2 106; 1:2 106; 0g using the LL-1-1 reference orbit. This transfer is shown in Figure 7(a) Now the end of this orbit can be propagated in 30 days more and the respective orbit shown in Figure 7(b) can be used as the reference one for the 330-day transfer between r0; r1 given by (19). Although this is not a periodic orbit yet" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002567_robot.2007.364203-Figure4-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002567_robot.2007.364203-Figure4-1.png", "caption": "Fig. 4. (a) The normal vectors of the left and right plane. (b) This figure show the cones as viewed along n1+n2. The vectors that define the double tangent of F1 and F2. The vector t1a and t1b are the results of rotating n1 around r1 by \u03b8 and \u2212\u03b8, respectively.", "texts": [ " Let P be the plane containing n1 and n2. Let ri \u2208 P be the vector that is perpendicular to ni. We restrict r1 to point toward n2 and r2 to point toward n1, i.e., r1 \u00b7 n2 > 0 and r2 \u00b7n1 > 0. Obviously, r1 and r2 are the normal vectors of the left and the right plane, respectively. The vectors on F1 that define the double tangents to F2 can be calculated by rotating n1 around r1 by \u03b81 and \u2212\u03b81. Also, The vectors on F2 that define the double tangents to F1 can be calculated by rotating n2 around r2 by \u03b82 and \u2212\u03b82 (see Fig. 4). C. R3-Positive Span of Four Force Cones This question is just an extension of the previous question. Let W = \u03a8(F1 \u222a F2 \u222a F3). First, let us consider the intersection of W with the unit sphere. Again, we extend each cone by the half-angle of F4, and check whether the axis of F4 lies inside W \u2032 = \u03a8(F \u2032 1 \u222a F \u2032 2 \u222a F \u2032 3) where F \u2032 i is defined in the same way as in Section IV-B. Fig. 5 illustrates W \u2032 as seen on the surface of the sphere. Obviously, a vector n4 is inside W \u2032 when either it is inside \u03a8(F1 \u222a F2), inside \u03a8(F2 \u222a F3), or inside \u03a8(F3 \u222a F1), or, finally, inside the pyramid defined by the axis of the three cones" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002565_j.tws.2008.08.010-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002565_j.tws.2008.08.010-Figure3-1.png", "caption": "Fig. 3. Cushion disc: (a) front view of the studied cushion disc", "texts": [ " In this work, we investigate the behavior of the clutch disc (Fig. 2). It allows a soft gradual re-engagement of torque transmission. This progressive re-engagement obtained by the clutch disc characteristics in the axial direction preserves the driver\u2019s comfort and avoids mechanical shocks. It also plays the role of a damper through the springs disposed around the hub. They enable the clutch disc to filter the torque variations of the combustion engine (Fig. 2). The axial elastic stiffness of the clutch disc is obtained by a cushion disc (Fig. 3(a) and (b)) which is a thin waved sheet, located between the two facings and fixed by rivets. It acts like a spring allowing a soft gradual re-engagement. This nonlinear axial stiffness is obtained by cutting the cushion disc into paddles and forming them to get the wavy shape (Fig. 4). The nonlinear axial elastic stiffness of the cushion disc is described by the cushion curve (Fig. 5). This load\u2013deflection curve gives the axial load versus axial displacement obtained by compressing a cushion disc between two flat pressure plates" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000718_1.1829068-Figure6-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000718_1.1829068-Figure6-1.png", "caption": "Fig. 6 Sun-ring-planet epicyclic spur-gear train", "texts": [ " To illustrate this and emphasize the previously discussed principle of kinematic equivalence and similarity, the kinematics of four mechanisms are compared: the simple sun-ring-planet gear train, the Humpage bevel-gear train, the harmonic drive, and the wobble gear. A line-by-line comparison of their kinematics may be the best way to demonstrate similarity and equivalence and lend credibility to the idea that they are members of a common mechanism class. Sun-Ring-Planet Planar Gear Train. The simple sun-ringplanet gear set is commonly accepted as the most basic of epicyclic gear trains. A typical representation of this train is shown in Fig. 6. The standard tabular method, or any other kinematic analysis method, gives the reduction-ratio results in Table 1, where Ni is the number of teeth on link i. Using the tabular method with the table considered as a matrix having entries Ai j , link i is considered fixed, and the speed ratio for any other two links j and k is simply Ai j /Aik . Humpage Epicyclic Bevel-Gear Train. The Humpage reduction set @11# shown in Fig. 7 is slightly more complex than the simple sun-ring-planet set in that it has a compound planet and an additional link ~besides being a non-planar mechanism" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003430_j.jelechem.2011.05.012-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003430_j.jelechem.2011.05.012-Figure2-1.png", "caption": "Fig. 2. The whole device used for the reverse iontophoretic experiments. The cross section of the biosensor was represented in the dash square. (1) Low chamber, (2) naked mouse skin, (3) biosensor, (4) inlet, (5) precise micro liter pipette, (6) lead line, (7) magnetic rotor, (8) CHI 660A electrochemical workstation, (9) insulating layer, (10) hydrogel membrane, (11) Os complex mediator film, (12) Ag|AgCl layer, (13) Au layer, (14) PC substrate.", "texts": [ " During the amperometric experiments, a potential of 0.1 V vs. the integrated Ag|AgCl reference electrode was employed. The in vitro experiments were performed in three-compartment vertical diffusion cells designed for iontophoresis experiments [4,25]. But in these experiments, both electrode chambers located on the surface side of skin were replaced by the hydrogel membrane on the biosensors. The hydrogel membrane served as the biosensor electrolyte as well as the collected reservoir of the extracted glucose (Fig. 2). A couple of biosensors were fixed by a device to make hydrogel membrane attached tightly to epidermis esis and the reaction sequence on the biosensor. of the naked mouse skin. The lower chamber of the diffusion cell fully contained the physiological buffer solution to simulate the change of blood glucose of animals at the desired level. The two biosensors were electrically connected to the CHI 660A electrochemical workstation and then a constant current of 0.5 mA cm 2 was applied via Ag|AgCl electrodes, which served as anodic and cathodic electrode respectively, to complete the reverse iontophoretic circuit" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0003573_s12239-011-0008-x-Figure12-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0003573_s12239-011-0008-x-Figure12-1.png", "caption": "Figure 12. Triaxial-transmission-error measurements using a rotary encoder.", "texts": [ " Therefore, in this study, the revolution angles of the three axes on the axle were measured to make up for the aforementioned disadvantages of the proposed method. The transmission errors in the input torque section of the axle, which has noise sections in vehicles, were measured using the triaxial transmission error evaluation. This method of evaluating the triaxial transmission error of the axle may also be applied to FF manual transmission and its axle. Figure 11 shows the pro- cesses of configuring and computing the triaxial transmission error, and Figure 12 is a diagram of the actual triaxial transmission error measurement experiments. Three ring encoders (ERM280) of HEIDENHAIN with three Rotec EDR counter boards were used on the input, LH output, and RH output of the axle. Each encoder rotation signal was calculated with RAS software by Rotec. Because a revolutionary angle meter has an angle revolution power of 20480 per revolution, it was possible to measure the transmission error with a high precision: 200 or more times the number of gear teeth (30~80) to be measured" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0002896_rnc.1682-Figure5-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0002896_rnc.1682-Figure5-1.png", "caption": "Figure 5. Axisymmetric rigid body with two controls.", "texts": [ " With the help of the generalized nonlinear H\u221e synthesis condition, we will be able to tackle the spacecraft command tracking problem directly by including the tracking error as one of output variables. The design process of such an output feedback nonlinear H\u221e control law is outlined below. Copyright 2010 John Wiley & Sons, Ltd. Int. J. Robust. Nonlinear Control 2011; 21:2079\u20132100 DOI: 10.1002/rnc The modeling of axisymmetric spacecraft has been fully addressed in [30, 31] and the references therein. It is assumed that the spacecraft is rotating around the symmetric axis b\u03023 on a two-dimensional plane as illustrated by Figure 5. Two control torques T1,T2 are directed along the asymmetric axes b\u03021, b\u03022 and they are perpendicular to each other and its symmetry axis. Use 1, 2, I1, I2 to denote the angular velocities and moment of inertia with respect to b\u03021, b\u03022. Two kinematic parameters w1,w2 are introduced through the stereographic projection [31]. They are used to describe the relative orientation of the body frame and inertial frame through two perpendicular rotations. By defining [x1 x2 x3 x4]T = [ 1 2 w1 w2]T, we obtain a fourth-order spacecraft state\u2013space model as \u23a1\u23a2\u23a2\u23a2\u23a2\u23a3 x\u03071 x\u03072 x\u03073 x\u03074 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a6= \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 0 am 0 0 \u2212am 0 0 0 1+x2 3\u2212x2 4 2 x3x4 0 m x3x4 1+x2 4\u2212x2 3 2 \u2212m 0 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u23a1\u23a2\u23a2\u23a2\u23a2\u23a3 x1 x2 x3 x4 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a6+ \u23a1\u23a2\u23a2\u23a2\u23a2\u23a3 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0000774_1.7143-Figure2-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0000774_1.7143-Figure2-1.png", "caption": "Fig. 2 Wrinkled membrane.", "texts": [ " The Cauchy stress \u03c3 of the membrane is given, by definition, as \u03c3 = (1/ det F)F \u00b7 S \u00b7 FT (11) Strains and stresses in taut regions of the membrane can be evaluated by Eqs. (7), (8), and (11). In wrinkled regions, however, these equations need to be modified in order to account for the wrinkling effects. Roddeman et al.9,10 have proposed a wrinkling model in which the deformation gradient tensor in wrinkled regions is modified so that the stress field in these regions is consistent with the TF theory. This subsection outlines the wrinkling model proposed by Roddeman. Consider a small part of a wrinkled membrane shown in Fig. 2. In the figure, a wrinkled membrane is represented by a surface with dotted lines. A pseudo mean surface, resulting from removing the wrinkles, is also indicated by a rectangular plane with solid lines. Let t and w denote unit vectors along and transverse to the wrinkles. From a basic assumption in the TF theory, the vectors t and w coincide with the principal axes of the Cauchy stress of the membrane. It is also supposed that the Cauchy stress in the t direction is positive and that in the w direction equals zero", " Alternatively, we attempt to approximate the deformation of the wrinkled membrane by the pseudo mean surface. Consequently, the deformation gradient F to be considered becomes that corresponding to the pseudo mean surface. However, the Cauchy stress \u03c3 calculated from this deformation gradient does not satisfy the uniaxial tension conditions just stated. This is because the deformation gradient F ignores the out-of-plane deformation of the wrinkled membrane and thus underestimates the length (1 + \u03b2)l of the wrinkled membrane in the w direction (Fig. 2). To recover the uniaxial tension conditions, the deformation gradient F has to be modified according to F\u2032 = (I + \u03b2w \u2297 w) \u00b7 F (12) where I denotes the identity tensor and \u03b2 represents a measure of the amount of wrinkliness. A physical interpretation of the term (I + \u03b2w \u2297 w) is that it stretches the wrinkled membrane surface along to w until its wrinkles just vanish, as shown in Fig. 3. This stretching involves rigid-body movements only because membranes are supposed to possess no bending stiffness in the TF theory" ], "surrounding_texts": [] }, { "image_filename": "designv11_20_0001196_1.2346690-Figure3-1.png", "original_path": "designv11-20/openalex_figure/designv11_20_0001196_1.2346690-Figure3-1.png", "caption": "Fig. 3 Same-sign CCS components", "texts": [ " Note that points 1, 3, 5, and 7 show purely tangential forces and points 2, 4, 6, and 8 are purely radial forces. The tangential forces can be considered as forces that affect the energy input into or out of the whirl orbit whereas the radial forces are restoring forces like direct stiffness. Based on the force diagram there are two issues to be considered: i radial distortion force on the orbit, and ii tangential forces responsible for rotordynamic stability. First, consider the radial forces acting on the circular orbit. Figure 3 shows a particular same-sign stiffness cross-coupling resultant force Fr at an angle between points 1 and 2, which decomposes into a tangential component Ft and a radial component FR. The radial component FR is a maximum at points 2, 4, 6, and 8 and zero at points 1, 3, 5, and 7. Using the radial force components from same sign-stiffness cross-coupling coefficients, a force distribution is generated as shown in Fig. 4. The dotted line force distribution represents positive valued stiffness cross-coupling coefficients and the solid line force distribution illustrates negative cross coupling: \u201ea\u2026 radial distortion forces ess cal orbits FEBRUARY 2007, Vol" ], "surrounding_texts": [] } ]