Feb 1, 1990 - accordion fashion provides a linear displace- ... ence on Robotics and Automation, Scottsdale, Ar- izona .... separated by a center shim or vane.
Design and Characterization of a Linear Motion Piezoelectric Micropositioner Zhixiao Wang, Musa K. Jouaneh, and David A. Dornfeld ABSTRACT: Small, lightweight, fast micropositioners provide fast, precise motion for applications in precision manufacturing. This paper describes the design and characterization of a precise linear micropositioner. The device is constructed of a number of piezoelectric plates placed on both sides of a sliding mass. Actuation of the plates in an accordion fashion provides a linear displacement of k0.36 mm. In addition, the device has a force capability of over 80 g, as well as position and force sensing capabilities. The dynamic performance of the device is characterized, and initial tests show that it has a second-order system behavior with nonlinear response. The device resolution is affected by the performance of the capacitive-type position sensor and varies from 3 pm to less than 0.6 pm.
Introduction There are many operations requiring precise motion of a tool or end-effector. One such operation is determining the surface roughness characteristics of machined components. Although the mechanical stylus has been used widely to obtain profiles of surfaces, the contact nature of the device, the difficulty of on-line measurements, and the slow tracking speed make it undesirable in automated manufacturing. As a result, a variety of other techniques have been developed in the past three decades to access surface roughness [l]. Of these techniques, optical profilometers seem to provide a promising approach for measurement of surfaces. Although different configurations of optical profilometers have been designed, a common design is to record the position variation of an objective lens that focuses a spot of light on the surface as a measure of height variations on the surface [ 2 ] , [3]. The position of the objective lens is adjusted in response to a feedback signal from a phoPresented at the 1989 IEEE International Conference on Robotics and Automation, Scottsdale, Arizona, May 15-19, 1989. The authors are with the University of California, Department of Mechanical Engineering, Berkeley, CA 94720.
todetector that monitors the intensity of the image so that it is kept in the best focal POsition as the profilometer traces the surface. This paper discusses the design and characterization of a compact micropositioner that is suitable for precision positioning applications, such as positioning the objective lens in certain optical profilometers. Interest in small, lightweight, fast manipulators for precise and fine motion applications has been receiving much attention in the last few years. In addition to providing a means for handling and moving delicate objects, they can improve the response speed and positioning accuracy of current industrial robots. Umetani and Suzuki [4] describe the use of micromanipulators such as microstabbers and micropipettes in the field of cellular biology and microsurgery. Fujita and Omodaka [5] describe the need for micromanipulators in handling of delicate objects for microchip manufacturing. Sharon et al. [6] have shown that a fast positioner attached to a general-purpose robot can compensate for the servo settling time of the robot, as well as for the tracking errors encountered when following a path. Several microactuators, based on various technologies, are reported in the literature. They include electromagnetic drive actuators [7], magnetically levitated micromachines [8], electrostatic actuators [5], [9], and shape memory alloy actuators [lo], [ l l ] . A microactuator driven by piezoelectric ceramic plates was chosen since piezoelectric materials have the characteristics of low mass, low heat generation, nonmagnetic, and low cost. A low-mass driver source gives a faster response, whereas low heat generation implies that high accuracy can be maintained since the positioner will not suffer a measurable loss of accuracy due to thermal expansion. A nonmagnetic material implies that the operation of the device will not be subjected to magnetic disturbances. Several piezoelectric microactuator designs have been reported in the literature. Twenty-six years ago, Ellis [12] described the design of several micromanipulators as fine toolholders for microdissection experiments under high magnification in biological research. The 0272 17089002000010 $01 00
10
c
micromanipulators, which had a maximum travel range of 0.6 mm, were driven by three piezoelectric plates through a linkage mechanism. Moriyama et al. [13] describe the design of a two-axis DC-motor-driven stage where an XYB piezo-driven fine stage is mounted on the first stage. The fine stage, driven by three piezoelectric actuators, has an 8-pm range of linear motion, a 160-prad range of rotational motion, and a positioning accuracy of k 0 . 0 5 pm. Each actuator is made of a single piezoelectric pipe. Umetani and Suzuki [4] describe the use of piezoelectric film in several microdevices, including a two-degree-of-freedom arm. The arm is composed of two pieces of bimorph plates connected together longitudinally with the both ends so as to bend perpendicularly toward each other. The arm, which is designed for microscopic operation, has a maximum range of travel of 0.350-0.500 rnm for a 300-V input. The micropositioner presented here has the unique features of a large motion range for a piezoelectric-driven device [ 141, [15], the versatility for trading off motion range against resolution, incorporation of position and force sensing in a compact fashion, and a new scheme of driving the piston. The micropositioner is constructed of two sets of piezoelectric bimorph plates stacked freely on both sides of a sliding mass. Excitation of the plates in an alternate fashion provides a linear motion of the mass. The microactuator is provided with a variablecapacitance transducer that is used as a position sensor. In addition, a set of two strain gauges mounted on the actuator is used as a force sensor. Both sensors were fabricated in our lab, thus trading the higher accuracy and better performance available in commercially developed sensors for expeditious availability. This paper is organized so that the next section gives a brief review of piezoelectricity, and then the mechanical design of the linear piezoelectric micropositioner is presented. Next, several tests to characterize its dynamic behavior are described, and the results are discussed. Finally, plans for future work and experimentation are proposed.
1990 IEEE IEEE Control Systems Magazine
Background Piezoelectricity (or “pressure electricity”) is a property of certain materials, such as quartz, Rochelle salt, barium titanate, and polyvinylidene fluoride polymer. If a force is applied to a solid crystalline dielectric, the crystal structure is distorted. The lattice distortion results in a net relative displacement between the positive and negative charges within the lattice, provided that the charge distribution within the crystal structure is asymmetrical. As a result of the internal charge displacement, equal and opposite external charges appear on opposite faces of the crystal. Conversely, if an electric field is established across a solid crystalline dielectric having an asymmetrical charge distribution in its lattice, the opposite of the preceding situation occurs. A displacement of these charges is induced, and a mechanical distortion of the crystal results. These two properties are the direct and converse piezoelectric effects first studied by Curie in 1880. Since then, piezoelectric materials have been used in a variety of acoustical and electromechanical transducers. For a detailed discussion on piezoelectric ceramics, see [ 161. Commercially, piezoelectric materials are available in a variety of forms, including rectangular or circular sandwichlike structures of two expander plates, in which case the unit is referred to as a bender bimorph [17]. The expander plates are made of polycrystalline piezoelectric ceramics and are separated by a center shim or vane. A bending-type response is obtained when an electric field is applied across the plates. The two plates can be connected either in series or parallel, as shown in Fig. 1. There is no structural difference between a series-type bender and its parallel equivalent. However, motion sensitivity in terms of deflection per unit of applied voltage is greater by a factor of 2 for parallel connections.
Piezoelectric Actuator The peak-to-valley height of a surface, which is a measure of the maximum variation in the profile of a surface, varies between I O and 20 pm for metal machined surfaces [I81 and up to several hundred micrometers for wood surfaces 1191. In addition, typical measurement resolution should be less than I pm for metal surfaces and around 5 pm for wood surfaces. These specifications were the primary guidelines for designing the micropositioner. To make the positioner usable for both metal and wood surfaces, a micropositioner with a motion range - up. to k0.36 mm and a structure for trading motion resolution against range was
February 1990
piston
c
plates
7
/
Series connection
%L Ball bearing
I+ I
11 ~
.
_
Fig. 3. Sectional drawing of micropositioner. 1
Parallel connection
Fig. 1. Electrical connections for a simply supported bimorph. designed and fabricated. An actual photograph of the micropositioner is shown in Fig. L.
Although the most common method for supporting piezoelectric plates in transducers is either in a cantilever beam fashion or a simply supported beam fashion (mounted freely on an edge on each end of the plate), the latter support method was selected in our case by choice, favoring larger displacements over force capacity. This follows since a number of piezoelectric bimorphs stacked serially and electrically driven to have alternate bending directions ideally can be thought of as a number of springhass units connected in series. Note that the deflection of a bimorph produced by a given signal is proportional to the product of the applied voltage and the capacitance of the bimorph.
Actuator Design A sectional drawing of the micropositioner is shown in Fig. 3. It consists of two compartments of rectangular cross section that are situated adjacent to a sliding mass. The mass slides on a set of ball bearings placed in a groove in the wall of the motor housing on each side of the mass. The groove is de-
signed to allow adjustment of the clearance between the balls and the sliding mass. Each compartment has a stack of 14 piezoelectric rectangular expander plates (bimorphs) arranged in pairs and freely stacked in the compartment. Currently, two different size bimorphs, 2 x 0.75 X 0.024-in. and 1.5 X 0.75 X 0.024-in. rectangular bimorphs made of PZT-5 ceramics (Piezoelectric Products, Inc., Metuchen, New Jersey), were used due to the desirability of having a large displacement. Electrical connections to the bimorphs were made in a parallel fashion. The bimorphs were placed in each compartment, such that each bimorph has its electrical connections on the side opposite to the connection of its adjacent neighbor. To ensure that the bimorphs will bend in a simply supported fashion, two round metal rods were soldered on either end of each plate; on the other side, a rod was soldered at the center. Initially, a voltage level, half of the maximum voltage range (100 V) used, is applied to each bimorph. This results in bending of the bimorphs such that each is bent in a direction opposite to its neighbor. Each pair will form a bowlike shape. Although not yet fully implemented, a knob switch selects a variable number of bimorph pairs to be activated so that different motion ranges can be selected. Since an equal number of bimorph pairs are activated in each compartment, the piston will be centered initially at the midpoint of the range of travel. If n , and n2 denote the number of small and large bimorph plates that are activated, respectively, and K , and K2 are the corresponding displacement coefficients per volt of applied voltage, then the ideal free-motion range of travel is given by
D,,,
Fig. 2. Photograph of micropositioner.
=
(nlKI + n2K2)Vm,,
(1)
The above equation gives only a rough estimate of the free-motion range since in the micropositioner built, the K‘s are dependent on the applied voltage level. Since we would like the bimorphs to be in contact with the sliding mass at all times, to be able to maintain a specified desired force, the actual range
11
of motion is smaller than without this constraint. To move the piston in either direction, a voltage level of +AV is added to the activated plates in one compartment, and -AV is added to the other, resulting in the motion of the piston in the direction of high to low voltage levels. The bimorph plates slide on the surface of insulating plates placed on the bottom and sides of the rectangular box. Since regardless of the number of bimorph pairs activated, the same voltage range is applied to the plates, this results in motion resolution that increases with a fewer number of plates activated. Thus the device has an inherent capability of trading motion range against resolution. In what follows, results will be shown only for the case of the all-bimorph plates being activated.
Position Sensor As a result of the micropositioner's small
i
range of travel, a variable-capacitance transducer with one side fixed to the frame and the other side attached to the moving piston is used as a position sensor. Figure 4 shows the circuit used to measure the sensor capacitance. If R L is the load resistance, C, the sensor capacitance, w the driving frequency of the square input voltage signal applied to the sensor plates, and C, the stray capacitance in the wires, then the voltage output is given by
The sensor is made of two 2 X 1-cm copper plates glued to an insulating holder. The fixed plate is mounted on a manually adjustable small table equipped with a micrometer. This arrangement is used for calibrating the sensor, where the sensor output is recorded as the table position is changed. The sensor is designed to operate in the displacement mode, where the displacement is measured relative to the initial position, since the zero point, which is dependent on the capacitor's plate spacing, is difficult to establish. The resolution of this sensor is not constant, and is dependent upon the spacing of the capacitor plates, being higher when the plates are closer together. The resolution varies from
0.6 to 3 pm for the range of travel. This nonlinear behavior is caused by having a large-size capacitor compared with the stray capacitance. This behavior is expected since, for a parallel-plate capacitor, with dielectric constant E, capacitor area A, and plate separation distance d, the capacitance is given by
the system for different step sizes. The figure shows a peak time of less than 15 msec for a 100-g step input in force. It also shows a nonlinear behavior that is dependent on the step size. The reasons for this behavior are discussed in the next section.
C, = EAId
To identify the dynamic behavior of the micropositioner, several tests were performed. These include a step-response test and a frequency-response test. A block diagram of the components of the open-loop system is shown in Fig. 6. A frequency-response test of the power amplifier shows a constant gain of 17 dB followed by a downward slope of -20 dBldec. The comer frequency is dependent on the amplitude of the input signal and varies from 400 Hz at a 1-V peak-to-peak input signal to 80 Hz for a 5-V peak-to-peak signal. There is a corresponding phase drop above the comer frequency. For low frequencies ( < 100 Hz), the amplifier response can be considered as a constant static gain. As discussed earlier, the calibration curve for the position sensor, shown in Fig. 7, shows a nonlinear relationship between piston displacement and sensor output. The calibration data were found to be best fitted by several exponential functions. The fitted functions were then used to generate a table of piston displacement versus the readings of the 12-bit analog-to-digital (AID) converter employed. The calibration data were not used directly to generate the table lookup due to the coarse resolution of the data. Since the nonlinearity is of a static type, it has a nonlinear effect only on the gain frequency response of the system. The steady-state relationship between input voltage to the amplifier and output displacement of the micropositioner is shown in Fig. 8. The figure shows a nonlinear dependence of the displacement on input voltage, with the displacement being dependent on the voltage magnitude applied to the microactuator since the last change of the direction of motion. The displacement-voltage relationship is assumed to be comprised of the sum of two parts: ( I ) a nonlinear displacement-voltage part with hysteresis (The linear portion of this curve has a slope equal to the inverse of the resultant stiffness of the 14 bimorph plates, each considered as a simply supported beam); (2) a Coulomb friction part that is dependent on the direction of motion. These two parts are shown in Fig. 9. Note that the manufacturer's literature [20] indicates that a linear relationship exists between input voltage to the bimorph plates
(3)
A linear relationship between the percent de-
flection signal AV,lV, and the change in plate separation distance is obtained only when plate capacitance is small compared with stray capacitance. This can be seen from Eq. (2), where C, >> C,.
Force Sensor
To make this micropositioner suitable for a variety of tasks, a sensor for measuring forces is incorporated into the design. A force sensor made of two strain gauges was constructed. The sensor is composed of two vertical, rectangular, thin metal plates placed parallel to each other. The plates are connected to form a box through two square, thick plates attached to the top and bottom sides of each plate. The top plate is designed to hold a small tool, whereas the lower plate is used to fix the sensor to the top of the sliding piston through screws. The strain gauges are attached to the inner side of the vertical plates. Two strain gauges were used to eliminate variation due to temperature and to increase sensitivity. The sensor has a force range of 200 g, which is over the 80-g design force capacity of the micropositioner. It has a linear resolution that can be set to well below 1 g. To illustrate the response speed of the actuatorforce sensor system for step change in force, Fig. 5 shows a normalized step response of
1.6 e,
c 0
E1.2 2
-0 g 0.8 1.0
U)
0
U
E
N ._ m 0.4
Z
0 0
Fig. 4. Electrical model of capacitance sensor.
72
10
20 30 Time, rnsec
40
50
Fig. 5. Force step response of posirionersensor system.
Dynamic Characterization
IEEE Control Systems Magazine
---
(5) becomes
Position sensor
Actuator
M,X
--U
+ 2kx
= kc(v,
- ~ 2 )
(6)
The input voltage to the elements (bimorph plates) is of the following form:
Fig. 6. Components of open-loop system.
+ U , sin (wt + 4) v2 = 6 - U , sin (wt + 4) U, =
6
Then Eq. (6) becomes
Me;
0
0.25
0.50 0.75 1.00 Displacement, mm
1.25
Fig. 7. Calibration curve for position sensor. 1
0.401
+ 2kx = 2kcv, sin (wt + 4)
(7)
This equation is of standard form representing the steady-state, undamped, forced harmonic motion. For this second-order system, the natural frequency is given by w, = (k& m,)1’2with k, = 2k. The experimentally observed resonance frequency as a function of input voltage was compared with the one computed from the preceding model assuming a linear ideal behavior for each element of the system (22 Hz). In this case, each bimorph plate stiffness is computed as that of a simply supported beam with a center load. The bimorph plates are considered as springs in series. The effective system mass is the sum of the actual mass of the piston and ball bearings and the effective mass of the plates. For a calculation of the effective mass of the system, see the appendix. For this linear behavior, k, = 2k,,, where k,, is the resultant stiffness of the 14 plates on each side of the piston. The resonance frequency results are displayed in the Table, and Fig. 11 shows a typical measured frequency response of the system with a 2-V peak-to-peak amplitude signal. The data indicate that as the amplitude of the voltage signal increases, the apparent “stiffness” of the bimorph plates decreases. Moreover, the Table shows also that as the amplitude level increases, the resonance frequency tends to level off to a value of 25 Hz. The resonance frequency at high amplitude ( - 2 5 Hz) agrees with the one com-
7-r(b) 0.4l
l
x
-
Fig. 9. Characteristics of steady-state model of the micropositioner. 0.0 1.6 3.2 4.0 6.4 Voltage input from A/D
-6.4 - 4 . 8 - 3 . 2 - 1 . 6
Fig. 8. Experimental steady-state displacement versus voltage relationship. and its displacement. However, the experimental data show a nonlinear relationship that is a function of the input voltage applied since the last change of the direction of motion. The cause of this nonlinear behavior is not yet fully understood and is under investigation. Whereas the actual system exhibits significant nonlinearities, as a first step, a simplified linear model for the micropositioner will be derived as a reference mark. For modeling, the system is essentially a single-degreeof-freedom system. It can be represented as an effective mass connected to a spring on either side. Figure 10 shows a diagram of the model. This simple model is intended to explain the resonance behavior of the piston; hence, damping is assumed to be small. Since the nonlinear force-voltage interaction of the piezoelectric plates is neglected, this model cannot be used for control purposes without further modification. For a voltage excitation of U , and v2 applied to the first and second elements, respectively, and with
February 1990
Fig. 10. Simplijed model of positioner.
F, and F2 denoting the forces exerted on the piston by the bimorph plates, Newton’s second law of motion gives
F I ( ~ ,-)
F 2 ( ~ 2 )=
(4)
Let k , and k2 be the stiffness of elements 1 and 2, respectively; cI and c2 the constants relating input voltage to zero-load displacement; 1, and l2 the actual instantaneous displacements of the elements; and F, and F2 given by F , ( u , ) = k , ( c , v , - I , ) and F2(vz) = kz(c2v2 - 12). Assuming k , = k2 = k and c , = c2 = c , and using the preceding relations, Eq. (4) becomes
k(c(v1
+ 12
- ~ 2 )
- 11) = M,X
(5)
If x is the equilibrium position displacement of the elements, and 1, = lo x and l2 = lo - x, then l2 - 1, = -2x. Using the preceding relation and the moving terms, Eq.
+
Table Resonance Frequency as Function of Amplitude Amplitude, V 0.5 1.o 1.5 2.0 2.5 3.0
4.0 5.0
Resonance Frequency, Hz 46.2 36.3 31.6 29.5 27.8 26.2 25.2 25.0
13
X = 36.307 HZ Y, = -6.6143dB FREQ RESP Y, = -77.578 deg FREQ RESP 0.0 180 dB
Phase Deg -180 -80.0 F,,Y 999.99 rn
Log, Hz
999.99
Fig. 11. Frequency-response plot.
puted from linear beam theory (22 Hz) using the preceding model. The 25-Hz value lies within the upper and lower bounds on the natural frequency estimation using linear theory (28 and 17 Hz, respectively). The upper bound is obtained by considering only the mass of the piston and bearings, whereas the lower bound is obtained by considering the total mass of the piston and plates as an effective mass. This quasilinear behavior at high amplitudes is expected since the effect of nonlineanties diminishes. The nonlinear behavior was also shown in the step-response results of the system. Figure 12 shows a normalized response for two different amplitudes in both the positive and negative directions. The delay at the start of the response is dependent on the amplitude level and decreases with increasing amplitude. The response shows a second-order underdamped behavior as expected.
pendent, nonlinear behavior. It also shows that the resonance frequency of the micropositioner for large-amplitude input signals is a function of the ideal stiffness of the bimorph plates. Since the stiffness and the effective mass are a function of the number and size of the plates, the designer has to choose between high bandwidth and a small motion range (but high resolution) that is obtained from few activated plates, or smaller bandwidth with a larger range of motion (with low resolution) for many plates. For practical use of the micropositioner, a closed-loop controller is needed. The dynamics of the micropositioner indicate the need for a nonlinear compensator. The hysteresis shown in Fig. 8 can be represented by adding another state variable, z , to our simplified model and by considering the input force to the positioner to be a function of the position, velocity, and the new state variable. A form for the dynamic behavior of z is given in [21]. As a result, a stable controller design and its implementation using this extended model are under way. Future experiments will include position and force tracking experiments, as well as using the positioner for measurement of surface roughness.
Appendix: Calculation of the Effective Mass The effective mass of the system is computed using the following formula, where M,,,,,, is the total mass of the piston and ball the total mass of the bibearings, Mspnng5 morph plates considered as springs in this case, and r the ratio of effective to total mass of the springs.
Future Work and Experimentation
Acknowledgments The authors would like to thank Professor Ron Fearing for providing the circuitry for the position sensor and P. Hsu for his help in designing the power amplifier circuitry. Thanks also go to Sami Bayyuk for useful discussions and suggestions on modeling of the microactuator and on computation of the effective mass, and to Professor Masayoshi Tomizuka as well as the reviewers for their helpful comments. Funding for this work was from a consortium of industries supporting the Laboratory for Manufacturing Automation in the Department of Mechanical Engineering at the University of California at Berkeley.
References [I]
[2]
[3]
141
[5]
[6]
The dynamic characterization of the micropositioner shows a strong, amplitude-de1.5 1
1 1.
-0
10
20 30 Time, msec
2.5 V
40
Fig. 12. Position step response of positioner.
14
I
50
Here r has been computed from an energy method analysis [22], assuming that the dynamic shape of each plate is identical to the static deflection shape. The effective mass of the end plates is 17/35 m, where rn is the mass of one pair of plates. The plates next to them have an effective mass of 17/35 m plus an effective mass due to the motion of the end plates. The next pair of plates has an end motion equal to the deflections of the preceding two pairs of plates and superimposed on that is their effective mass of 171 35 m. Summation of all the effective masses and division by the total mass yields r = 0.354. This value of r is plausible by comparison with the value for springs ( r s 1/3) and for simply supported beams ( r = 171 35).
171
[8]
[9]
[IO]
I . Sherrington and E. Smith, “Modem Measurement Techniques in Surface Metrology: Part 11; Optical Instruments.” Wear, vol. 125, no. 3 , pp. 289-308, 1988. 0. Dupuy, “High Precision Optical Profilometer for the Study of Micro-Geometrical Surface Defects,” Proc. Inst. Mech. E n g . . vol. 182, pp. 255-259, 1967-1968. F. Arechi, D. Bertani. and S . Ciliberto, “A Fast Versatile Optical Profilometer,” Opt. Cornmun., vol. 31. no. 3, pp. 263-266, 1979. Y . Umetani and H. Suzuki. “Piezo-Electric Micro-Manipulator in Multi-Degrees of Freedom with Tactile Sensibility ,” Proc. 10th Intl. Synp. Ind. Robots, pp. 571-579, 1980. H. Fujita and A. Omodaka, “Electrostatic Actuators for Micromechantronics,” 1987 IEEE Micro Robots and Teleoperators Workshop, pp. 83-92, Nov. 1987. A. Sharon, N. Hogan, and D. Hardt, “More Analysis and Experimentation on a Macro/ Micro Manipulator System,” Proc. Symp. Modeling and Control of Robotic Manipulators and Manufacturing Processes, ASME Winter Annual Meeting, Boston, MA, pp. 417-422, Dec. 1987. R. Hollis, “Design for a Planar X Y Robotic Fine Positioning Device.” Proc. Sjmp. Robotics und Manufacturing Automation, ASME Winter Annual Meeting, pp. 291298, Miami, FL, Nov. 1985. R. Pelrine and 1. Busch-Vishniac. “Magnetically Levitated Micro-Machines.” 1987 IEEE Micro Robots and Teleoperators Workshop, pp. 113-117, Nov. 1987. R. Mahadevan, “Capacitance Calculations for a Single-Stator, Single-Rotor Electrostatic Motor,” 1987 IEEE Micro Robots and Teleoperutors Workshop, pp. 93- 100, Nov. 1987. K . Kuribayashi, “A New Compact Robot Hand Using Shape Memory Alloy Actuator, ’’ Proc. Japan- USA Symp. Flexible Automation, pp. 393-400, Osaka, Japan, July 1986.
IEEE Control Systems Magazine
J. Walker. "A Small Rotary Actuator Based on Torsionally Strained SMA," 1987 IEEE Micro Robots aiid Teleoperurors Workshop, pp. 106-108. Nov. 1987. G . Ellis. "Piezoelectric Micromanipulators," Sciriiw, vol. 138. pp. 84-91, Oct. 12, 1962. S. Moriyama. T . Harada. and A. Takanashi, "Precision X-Y Stage with a Piezodriven Fine-Table." B d / . Jcipciii Soc. Precisioii Eiig.. vol. 22. no. I . pp. 13-17. Mar.
1988. I141 H. Goto and T . Sasaoka. "Vertical Micro Positioning System Using PZT Actuators." B i t / / . Jupirii Soc. Prri.i.sion Erig. . vol. 22. no. 4. pp. 277-282. Dec. 1988. l l 5 l S. Moriyama, F. Uchida, and E. Seya. "Development of a Precision Diamond Turning Machine for Fabrication of Asymmetric Aspheric Mirrors." Opt. Eiig. . vol. 27, no. I I , pp. 1008-1012, Nov. 1988. B. Jaffe. "A Primer on Ferroelectricity and Piezoelectric Ceramics." Eng. Memo. 6014, Clevite Corp.. Electronic Re5earch Div., Dec. 1960. C. Germano. "Flexure Mode Piezoelectric Transducers." IEEE Tretris. Audio t r i i d Nrctroticoirstics. vol. AU-19. no. I . pp. 612, Mar. 1971. R. DeVor and S . Wu. "Surface Profile Characterization by Autoregressive Moving Average Models," Pro(,. ASME "AM. Paper 7 I-WAiProd-26. Washington. DC, 1971. C . Maxey. "Measuring Texture and Contact Area of End-Wood Surfaces." Mcrtrrinls Resecrrch criid Stmc1urd.s. vol. 4. no. 6. pp. 279-285. June 1964. Piezocrruriiic, B e d e r Elriiimts. Piezo Electric Products, Inc., Metuchen. NJ. Y. Wen, "Methods of Random Vibration for Inelastic Structures,'. Appl. Mrchcrri. Re\,., vol. 42, no. 2, pp. 39-52, Feb. 1989.
[22] S . Tinioshenko, D . Young. and W . Weaver. Vibrurioii Pro6Ieni.s in Engirieeririg. 4th Ed.. Wiley, New York, 1974.
mechanical design. and manufacturing automation.
Zhixiao Wang received the B.Sc. and M.Eng. degrees in mechanical engineering frorn the Harbin Institute of Technology. Harbin, China, in 1982 and 1985. respectively. From 1985 to 1987. Mr. Wang was a lecturer at Harbin Institute of Technology, and was a Visiting Research Engineer at the University of California at Berkeley in 1987 and 1988. Mr. Wang is currently a Ph.D. student i n the Mechanical Engineering Department at the University of California at Berkeley. His research involves manufacturing automation, robotics, and neural networks.
David A. Dornfeld received the B.S.. M.S., and Ph.D. degrees in mechanical engineering from the University of Wisconsin-Madison in the area of production engineering. His Ph.D. thesis concerned the stud) of the fundamentals of the mechanical pulping process (abrasive machining). He joined the faculty of the University of California at Berkeley in the Mechanical Engineering Department in 1977 and is presently Professor of Manufacturing Engineering and Director of the Engineering Systems Research Center in the College of Engineering. Dr. Domfeld's research activities are in several fields of manufacturing engineering and flexible automation: acoustic emission monitoring and analysis of nianufacturing processes, precision manufacturing. and intelligent sensors and signal processing for robotics and process automation. He has published more than 80 papers in these fields and is currently editing a series on computer integrated nianufacturing for Addison-Wesley. Professor Dornfeld is an active member of the ASME. both contributing to the technical programs and ,journals o f the society and as Senior Technical Editor. Trcrii.scic.riori.c of ASME. Jounicil of EngiiiecJriiig fix Iiidusrr\. as well as SME, AWS. Japan Society of Precision Engineering. and the U.S. Acoustic Emission Working Group. He is President-Elect of the Board of Directors, North American Manufacturing Research Institute.
Musa K. Jouaneh received the B.Sc. degree in mechanical engineering from the University of 1984. He obtained the M.Eng. and Ph.D. degrees in mechanical engineering from the University of California at Berkeley in 1986 and 1989, respectively. He is currently a Postdoctoral Research Associate at Berkeley. His research interests include robotics.
1990 E R A The 1990 IEEE International Conference on Robotics and Automation (ICRA) will be held on May 13-18, 1990, at the Hyatt Rcgency in Cincinnati, Ohio. The General Chairman is Professor R. A. Volz of Texas A&M University, and the Program Chairman is Professor A. J. Kolvo of Purdue University.
February
1990
Thc theme of this conference is "Intelligent Automation and Robotics" with emphasis on information technology for sensorbased systcms. Original basic and applied papers in all areas of automation and robotics are included. The conference hosts workshops and tours on Sunday, May 13, and Friday, May 18, 1990, and tutorials on Mon-
day, May 14. Conference sessions will be held on Tuesday, May 15, to Thursday, May 17, 1990. For further information, contact Harry Hayman at (305) 483-3037 or write to Robotics and Automation, P.O. Box 3216, Silver Springs, MD 20901.
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