A piezoelectric linear actuator formed from a multitude of bimorphs

Sensors and Actuators A 109 (2004) 242–251

A piezoelectric linear actuator formed from a multitude of bimorphs James Friend∗ , Akira Umeshima, Takaaki Ishii, Kentaro Nakamura, Sadayuki Ueha Precision and Intelligence Laboratory, Tokyo Institute of Technology, 4259 Nagatsuta, Midori-ku, Yokohama 226-8503, Japan Received 2 June 2003; received in revised form 24 October 2003; accepted 24 October 2003

Abstract This paper describes the design, construction, and testing of a linear actuator using a multitude of cantilever bimorphs. Each beam is vibrated both in the fundamental axial mode and a high-order flexural mode using piezoelectric material with a specific electroding pattern to obtain both axial and flexural motion from each bimorph. Bidirectional linear sliding velocities and forces reached 27.7 cm/s and 0.09 N, respectively, from a single bimorph, while up to 17.6 cm/s velocity and 0.15 N in either direction, at approximately 1% efficiency, were obtained from an actuator composed of four closely-matched bimorphs. © 2003 Elsevier B.V. All rights reserved. PACS: 85.50.-n; 43.38.Fx Keywords: Piezoelectric actuator; Linear actuator; Bimorph; Ultrasonic actuator

1. Introduction Linear actuators are used throughout industry and in commercial applications, and as research and manufacturing move toward smaller and smaller products, even beyond microelectromechanical systems (MEMS), the actuators used in these industries must also become smaller. Piezoelectric linear actuators, in particular, promise high-force linear actuation with no backlash, extremely low minimum speeds, positioning accuracy measured in nanometers, and offer inherent braking when power is removed [1]. Despite these potential benefits, most piezoelectric actuator research has focused on rotary motion. There are a few exceptions. Alps Corporation, in particular, developed a small linear motor system [2–4] using either one or two multilayer piezoelectric actuators (MLPAs), producing up to 3.6 N sliding force or 30 cm/s sliding velocity. Using a single MLPA, Friend et al. [5] constructed a linear actuator giving maxima of 1.86 N force and 16.5 cm/s velocity at 18.9% efficiency. Kurosawa et al. [6] used surface acoustic waves transmitted along the surface of a LiNbO3 wafer to form an extraordinarily powerful actuator, with sliding forces and velocities reaching 3.5 N and 1.1 m/s, respectively.

∗ Corresponding author. Tel.: +81-45-924-5052; fax: +81-45-924-5091. E-mail address: [email protected] (J. Friend).

0924-4247/$ – see front matter © 2003 Elsevier B.V. All rights reserved. doi:10.1016/j.sna.2003.10.040

This paper describes the design, testing, and construction of a linear actuator using a collection of simple bimorph cantilever beams driven by piezoelectric material. When placed together in a set using suitable electrical driving circuitry, the beams can provide linear motion through the generation of axial vibration along their main axis combined with flexural vibration. Bimorphs and their application have been studied for over 60 years [7] by many researchers. Their relatively large bending displacement has proved useful in many devices, particularly in actuators for robotics [8], optics [9–11], and other applications [12–14]. To the authors’ knowledge, only recently have researchers begun to consider using axial excitation of the bimorph in conjunction with flexural excitation [15–19]; by matching the bimorph’s axial resonance to a flexural resonance, it is possible to obtain elliptical motion from the bimorph tip. In the following sections, a description of the basic actuation concept is followed by the refinement of the design using finite element analysis and an experimental study to determine the actuator’s performance characteristics.

2. Concept The basic idea of the actuator is to combine a series of bimorphs together in a column, with the output tips of each actuator aligned as shown in Fig. 1. With this configuration, motion can be imparted to a slider.

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Table 1 Characteristics of the NA piezoelectric material used in this study from Taiheyo Cement

Fig. 1. Combining several bimorphs together provides a method for generating sliding motion along the tips of the bimorphs.

The sines and cosines refer to the relative temporal phase of the applied driving signal between each actuator placed in parallel. By using this arrangement, the motion along the tips is akin to bacteria cilia or millipede legs. Changing the direction of motion is accomplished by changing the direction of elliptical motion within each actuator through reversing the sign of the phase α, illustrated in Fig. 2(a). Even though the relative phase between each actuator remains the same, the sign change of the phase alpha will cause the motion of the slider placed against the tips of the bimorphs to reverse. The phase α is independent of the phase separating each bimorph in the completed actuator; the latter remains fixed at one-quarter of a wavelength for the purposes of this study. Each bimorph uses a thin plate of phosphor bronze with lead zirconium titanate (PZT) placed on either side of the plate, forming a bimorph. The top surface of each PZT plate is divided into two electroded areas, with the opposite surface electroded over its entire area. The polarization of each piezoelectric plate is indicated by the large arrows in Fig. 2(b). A so-called “hard” piezoelectric material was

Property

Value

Dielectric constant 11 33

1750 1460

Coupling coefficient kr k31 k33 kt k15

0.58 0.33 0.69 0.55 0.73

Piezoelectric strain coefficient d31 (×10−12 m/V) d33 (×10−12 m/V) d15 (×10−12 m/V)

−120 315 580

Piezoelectric voltage coefficient g31 (×10−12 V/N) g33 (×10−12 V/N) g15 (×10−12 V/N)

−10.8 25.0 37.5

Elastic constant E (×10−12 m2 /N) s11 E (×10−12 m2 /N) s33

12.1 16.6

E (×10−12 m2 /N) s66

31.5

Loss tangent, tan δ (%) Quality factor, Qm Density, ρ (×103 kg/m3 )

0.5 2000 7.6

used in this work, Taiheyo cement type NA; its properties are provided in Table 1. Phosphor bronze was chosen because the material has a low magnetic permeability; at about 1.1–1.2µ0 , the entire actuator is neither magnetic nor affected by strong magnetic fields. With the application of electric voltage as indicated on electrodes 2 and 4 in Fig. 2(a), the beam is bent transversely (along the z-axis) due to the opposing in-plane expansion and contraction of the two PZT plates. Electrodes 1 and 3, however, are driven in-phase, and so their effect is to deform the bronze plate axially (along the x-axis). A sinusoidal signal is used for all the electrodes, with a phase difference, α, between the 1–3 and 2–4 electrodes to permit the generation of elliptical motion along the beam, particularly at

Electrode 1 (sin ωt) Electrode 2 ( sin (ωt-α) )

PZT

Holder

Phosphor Bronze Beam

Ground 0

z

y x

Electrodes

PZT

Electrode 4 (-sin (ωt-α)) Electrode 3 (sin ωt ) (a)

(b)

Fig. 2. A bimorph for this application: (a) as assembled (note ground 0, and electrodes 1-4) and (b) its components.

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3. Design refinement via finite element analysis

Fig. 3. An illustration of the vibration of a free–free bimorph with PZT poled as shown (a), using (b) a single sinusoidal input, (c) a pair of opposing sinusoidal inputs, and (d) both similar and opposing sinusoidal signals on different portions of the bimorph, with a temporal phase difference, α.

the right end, indicated with an asterisk in Fig. 2(a). The idea is illustrated in Fig. 3 using a free–free bimorph for clarity. Elliptical motion as suggested in Fig. 3(d) is possible when the phase between the axial excitation and flexural excitation of the bimorph’s motion is one-quarter of a wavelength out of phase, as described in a variety of previous works on actuators using elliptical motion [20–22] and traveling-wave motion [23,24]. By fixing one end of the actuator, the actuator may be held, and output motion on the remaining free end is still elliptical in shape.

Fig. 4. Definition of the (a) geometry used for finite element analysis with a (b) example mesh used in the analysis.

The finite element analysis software ANSYS was used to refine the design of the actuator and obtain the largest possible motion from the output tip. While certain aspects of the geometry were fixed due to constraints with the piezoelectric material, its poling, or the like, much of the design remained variable. In particular, the electrode gap, piezoelectric element thickness, and alignment gap were all fixed at 0.5 mm. Those dimensions and others are shown in Fig. 4. The phosphor bronze plate has a width of W and thickness of t, while its length is the sum of its extension into the holder, the alignment gap, length of the PZT, lPZT , and tip length, l. The electrodes on the PZT may have different lengths lE1 and lE2 . This permits different ratios of axial-to-flexural excitation, based on the amount of PZT driven for each kind of motion, and changes the interaction between the generated moment distribution and driven vibration shape within the bimorph. Later, the motion of the tip is studied at three

Fig. 5. The (a) bending and axial modal displacement distributions, the (b) forward and reverse bending/axial mode shapes at 63.0 kHz, and (c) the calculated displacement distributions for the axial and bending directions, respectively; the displacement units are arbitrary.

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discrete locations, referred by “left”, “center”, and “right” as shown in the figure for convenience. In ANSYS, standard brick elements were used to model the structure, with piezoelectric degrees of freedom and poled anisotropy used in the piezoelectric material components. The results of an analysis are given in Fig. 5; the axial and transverse modal displacement distributions along the x-axis are shown at 63 kHz using grounded electrodes. Here, the resonance frequencies of the fundamental axial mode and the third flexural mode were matched, giving a pair of combined axial–transverse vibration modes, each 180◦ out of phase with the other. The bending and flexural modes, and the actual combination of the two modes are shown in Fig. 5. As the bimorph extends axially, its tip bends in the minus Z-direction. The tip, for the other mode, bends in the positive Z-direction as the bimorph extends axially. Though bending vibrations and axial vibrations are often viewed as being separate, when one bending mode’s frequency is close to a flexural mode’s frequency, the motion of each mode takes on characteristics of the other mode, yet the modes remain orthogonal. Another example of this phenomenon was shown by Friend et al. [5], in another type of actuator. The displacement units are arbitrary, though the ratio of axial to transverse displacement is not; the transverse modal displacement is roughly three times the axial modal displacement at the tip in this configuration. The third flexural mode was chosen in this study, though other flexural modes might be useful depending on

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the design. Here, the third flexural mode exhibited a large strain in the piezoelectric elements, making it easy to generate, unlike the fundamental and second flexural modes, and the tip deflection was higher in the third mode than the fourth and higher modes. For this case, the tip length, l, was 1.5 mm, while the electrode lengths lE1 and lE2 were both 4.5 mm, the thickness of the bronze plate, t, was 0.2 mm, and the width of the actuator, W, was 4 mm. As a result, the length of the PZT was 10 mm. Using the same configuration, Fig. 6 shows the change in direction of the tip based on the value of the phase angle α. By changing the length of the tip, the response at the tip of the actuator can be tailored to form elliptical motion. For a short tip length of 1.5 mm, the motion of the left, right, and center portions of the tip are essentially identical regardless of the lengths of the electrodes. However, for longer tip lengths, the magnitude of transverse motion varies significantly along the width of the tip (y-axis). This, by itself, might merely cause excess wear during the actuator’s operation, but, unfortunately, the phase of the transverse motion also varies along the tip in these cases. The problem is illustrated in Fig. 7(a). The direction of transverse motion changes along the tip, which is clearly unacceptable. These unfortunate motions are being excited in the actuator because the structure has a finite width. Along y = ±W/2, the edges are free, and so the PZT causes bending along the y-axis as well. The phase of the transverse motion changes

Fig. 6. The effect of changing the phase angle α on the tip motion for an applied voltage of 1 V0−p ; the arrows on the circle beneath the values of α indicate the direction the tip travels around the ellipses. Note that the left, center, and right locations (L, C, and R in Fig. 4) are plotted together for each case, and are almost identical.

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output force as a function of the input voltage or the output vibration velocity as a function of the input current. The term, as used here, is taken from Ueha and Tomikawa [1]: A=

i F = , v V

(1)

where A is the force factor, i and V are the input current and voltage to the device, and v and F are the output vibration velocity magnitude and maximum force. The definition is over 60 years old, and serves as one estimate of the potential output force of the actuator if one only knows the output vibration velocity and input signal characteristics. In finite element analysis, the characteristics of the input signal and output vibration velocity are much easier to determine than the output force, due to the high computational cost of including contact mechanics in the analysis. As a result of the finite element design study, the actuator’s dimensions were fixed at l = 1.5 mm, lE1 = lE2 = 4.5 mm, t = 0.2 mm, and W = 4 mm. The electrode and alignment gaps were set at 0.5 mm, and the PZT length and thickness were fixed at 10 and 0.5 mm, respectively. The following sections describe the construction and testing of the actuators.

4. Actuator construction and experimental setup

Fig. 7. The direction of vibration of the center (C) of the tip (a) opposes the tip’s corners (L, R) when the tip length is 4.5 mm, but (b) is the same when the tip is 1.5 mm long. The heavy arrows indicate the direction of travel; note the axes are asymmetric for clarity.

from positive (at L) to negative (at C) and back again (at R); the corners of the tip move in opposition to the center of the tip. The solution used in this case was to reduce the length of the tip to 1.5 mm, where the motion along the tip is not only in the same direction but also roughly the same magnitude, as shown in Fig. 7(b). Even though the phasing of portions of the output tip were found to be unaffected by the electrode configuration, the force factor is strongly dependent on the relative size of the electrodes, as shown in the calculated portion of Table 2. The equal-length electrode choice is a trade-off between the axial and transverse force factors, and is therefore likely the best candidate: the axial and transverse force factors are roughly the same. The term force factor refers to either the Table 2 The force factor for each bimorph type lE1 (mm)

1 4.5 8.5

lE2 (mm)

8.5 4.5 1

Force factor (N/V) Measured

Calculated

Error (%)

x-axis

z-axis

x-axis

z-axis

x-axis

z-axis

0.104 0.090 0.063

0.030 0.042 0.054

0.127 0.108 0.079

0.036 0.091 1.300

20 18 22.5

18 73.7 184

Three sets of actuators were fabricated, using combinations of the three electrode length sets {lE1 , lE2 } = {1, 8.5}, {4.5, 4.5}, and {8.5, 1}. Though the finite element analysis predicted that the equal-length electrode actuator would be superior to the other actuators, the alternatives were built to experimentally test this hypothesis. The phosphor bronze was polished flat to a 5 ␮m finish onto which the PZT was attached using epoxy adhesive. The output ends of the bimorphs were mounted in a special fixture and then polished smooth (also 5 ␮m) and parallel to the yz-plane of the actuator. The bimorphs were mounted in parallel slots cut into a 10-mm thick aluminum plate fastened onto a table. Depending on whether laser Doppler vibrometry of a single bimorph, performance testing of a single bimorph, or testing of the complete actuator was to be performed, the configuration of the experiment was slightly different, as shown in Fig. 8. The measurement of the vibration displacement distributions on single bimorphs were performed using a 2-D in-plane laser Doppler vibrometer (Polytek PI LDV-500/OFV-2802) and a standard single-point laser Doppler vibrometer (Polytek PI CLV-1000). Measurement of the performance of a single bimorph was accomplished by pressing a standard tool-steel rotary ball bearing against the output tip (Fig. 8(b)). The bearing’s diameter was 10 mm, making the subtended angle of contact with the bimorph tip 0.04◦ , implying the contact was parallel and flat to a good approximation. A sleeve was attached to the bearing to allow a mass to be suspended by a thin monofilament line wound around the sleeve. For the testing

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Fig. 8. The test fixture for (a) measuring the vibration velocity using a laser Doppler vibrometer (LDV), (b) measuring a single bimorph’s performance, and (c) measuring the actuator’s performance.

of a collection of bimorphs, the same bearing was used to press a 0.6-mm thick bakelite plate against the tips of four bimorphs, (Fig. 8(c)). Bakelite was chosen because of its flatness, rigidity, smooth surface finish, and compatibility with phosphor bronze as a contact material. The bimorphs were each separated by a distance of 3 mm. The steady-state condition of either the bimorph or the actuator’s operation was measured, allowing the moment of inertia of the bearing to be ignored in either case. The drag of the bearing was also ignored, being almost negligible at 0.13 mN along a tangent to the perimeter of the bearing. In all cases, a dual-output signal generator with controllable phase difference (Yokogawa FG-120) was used to a pair of amplifiers (NF HSA-4010) via two sinusoidal signals. Each amplifier was connected to one transformer. The output of one transformer was connected to electrodes 1 and 3 in Fig. 8(a). The other transformer was used for electrodes 2 and 4; the center tap of the output was grounded to the aluminum holder 0, and the high and low sides of the transformer output were used as the positive and negative outputs necessary for the sign change between electrodes 2 and 4. To measure the sliding speed, a digital video camera was used in conjunction with a computer to measure the motion of dots painted on the spinning part of the bearing and on the slider from frame to frame of the video. For the slider, the

preload force was supplied through the bearing, which was pressed against the slider using a spring. The deflection of the spring indicated the preload. A variety of sliding loads were applied to the bearing by winding a string around a sleeve attached to the bearing and placing different masses under gravity onto the string. The same string and masses were used to apply different sliding loads on the slider. 5. Results 5.1. Admittance characteristics of the actuator The admittance and phase characteristics of the actuator, with four bimorphs mounted in parallel as described earlier, are shown in Fig. 9. The first, second, third, and fourth electrodes of each bimorph, respectively, were placed in parallel with each other and connected to the transformers as described above. Ignoring the resonances, the admittance increases linearly rather quickly; since the actuators are connected in parallel the area of PZT material is large, forming a large shunt capacitor and causing a steady increase in admittance with respect to frequency. The presence of four very similar, but not identical, bimorphs generates four closely spaced, but distinct, resonances at each general resonance frequency. The spacing of the resonances depends

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the damping and the lack of the epoxy bond in the finite element analysis between the layers of the bimorph. 5.2. Displacement distribution within the bimorph

Fig. 9. The admittance characteristics with respect to frequency for the actuator, a set of four bimorphs electrically in parallel (see text).

on a number of factors, mostly the quality of construction and assembly, and has traditionally been a problem with using multiple piezoelectric devices together to form an actuator [25]. In this case, the resonances of interest at around 60–70 kHz are narrowly spaced between 69.35 and 70.23 kHz with an overall quality factor of 105 (i.e., computed by including all four resonances together). Precision assembly helps to make this achievement possible. Note that the experimentally measured resonance frequency is fairly close to the predicted frequency, with an error of about 9%. The error is believed to be due to difficulty in predicting

The vibration displacement distributions along the x-axis for the three different electrode configurations are shown in Fig. 10. From Fig. 10, the axial displacement at the tip drops almost to zero as electrode 1 becomes longer, while the flexural displacement magnitude is largely unaffected. Note the similarity between the predicted displacement distribution in Fig. 5 and the measured displacement distribution in Fig. 10(b). Knowing the displacement magnitudes at the tips of the actuators and the input current into the actuators permits the calculation of the force factor akin to the method used in the finite element analysis; the results of the measurement and calculation is provided in Table 2. The measured and calculated force factors are the average of the right, center, and left parts of the output tip from the experiment and finite element analysis, respectively. The measured force factor is consistently lower than the calculated values, likely for the same reasons that caused the frequency prediction to be different. The measured and calculated force factors, however, follow the same trends as the electrode lengths are changed. As electrode 1 becomes longer (lE1 ), the axial force factor decreases, while the flexural force factor increases. In particular, the choice judged superior in the finite element analysis, where the electrodes have the same

Fig. 10. Measured displacement distribution at 60 kHz of a single bimorph at 50 V0−p for (a) {lE1 , lE2 } = {1, 8.5}, (b) {lE1 , lE2 } = {4.5, 4.5}, and (c) {lE1 , lE2 } = {8.5, 1}.

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Fig. 11. Measured bearing velocity versus input phase for {lE1 , lE2 } = {4.5, 4.5}, {8.5, 1} mm, with a preload of 0.14 N, at 69.1 kHz.

length, is a trade-off between the decreasing axial force factor and increasing flexural force factor in the experiment. 5.3. Operating performance of the bimorph A single bimorph’s ability to generate sliding motion was tested as described earlier; the results are shown in Fig. 11 for {lE1 , lE2 } = {4.5, 4.5} mm and {8.5, 1} mm. The excitation frequency was 69.1 kHz. In this section, the applied voltage was 50 V0−p . The third configuration, {lE1 , lE2 } = {8.5, 1} mm, was unable to turn the bearing. For the equal-length electrode version, the sliding velocity was reversible and reached maximum speeds of 23.6 and 20.9 cm/s for input phases of −53◦ and 109◦ , respectively. Though, in theory, the maximum speeds should be reached at ±90◦ , the presence of the axial resonance shifted the true maxima and minima. With {lE1 , lE2 } = {1, 8.5} mm, the actuator turned the bearing over the phase range 60◦ to −150◦ (−150◦ = 210◦ ), but the direction of rotation did not change. Notably, the rotation stopped almost entirely over the phase range where the bearing turned in the opposite direction for the equal-length electrode configuration. Using a weight attached to a string wound around a sleeve, itself affixed to the bearing as described earlier, the efficiency and sliding force versus velocity were measured as shown in Fig. 12 with a preload force of 0.44 N. The velocity of the sliding surface achieved in these tests was 27.7 cm/s, with a maximum sliding force of 90 mN. Beyond 90 mN, the rotation of the bearing was unstable and prevented reasonable measurement of the sliding velocity. Overall, the efficiency was very low at approximately 1%.

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Fig. 12. Measured sliding velocity and efficiency versus sliding force at 69.1 kHz. The line represents a linear curve-fit of the sliding velocity.

age was measured. The operating frequency of the actuator was 69.1 kHz. A video of the actuator’s operation is provided on http://www.ueha.pi.titech.ac.jp/BimorphActuator. wmv. The velocity of the slider versus input phase α is shown in Fig. 13 for a preload force of 0.44 N. Similar to the results for a single bimorph (see Fig. 11), the direction of sliding reverses as the phase is changed. The maximum forward velocity is 8.57 cm/s, while the maximum velocity in the opposite direction is 11.1 cm/s. For this measurement, the applied sliding force was 0.1 mN, essentially the drag of the bearing. The sliding velocity is almost linearly dependent on the applied voltage above 20 V0−p ; Fig. 14 illustrates the relationship when varying either the axial, transverse, or both electrode voltages. When only one voltage was to be changed, the remaining voltage was held at 77 V0−p . The input phase α was set at 60◦ . During these measurements, the axial velocities reached 17.6 cm/s but an applied sliding force of only 0.1 mN. The actuator’s sliding force and efficiency versus velocity are plotted in Fig. 15; the sliding velocity is the steady-state velocity. Though only one direction of motion is shown in the figure, the actuator was bidirectional, with a negligible difference in performance between the two directions. Notice that while the efficiency remains below 2% throughout,

5.4. Operating performance of the actuator After assessing the performance of a single bimorph, a collection of four equal-length electroded bimorphs were placed together in the aluminum plate described earlier. Using a bakelite plate as a slider, with the rotary bearing pressed against the plate to provide a preload, the velocity of the slider versus input phase, sliding load, preload, and volt-

Fig. 13. Measured sliding velocity versus input phase at 69.1 kHz for input voltages of 53, 63, and 78 V0−p .

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the sliding force and velocity generally increase with an increase in preload to maxima of 0.15 N and 18.1 cm/s, respectively. The performance of the actuator decreased dramatically when using a preload greater than 1.1 N, because the bimorphs tended to buckle under heavier axial loads.

6. Conclusions

Fig. 14. Measured sliding velocity versus input voltage, zero-to-peak, on either the axial electrodes (1 and 2), the transverse electrodes (3 and 4), or all electrodes together. The fixed-voltage electrodes are set at 77 V0−p ; the lines represent logistic least-square fits of the data.

The design, construction, and testing of a linear actuator using a collection of bimorphs was demonstrated. For this study, finite element analysis of the structure proved useful for both designing and predicting the experimental characteristics of the bimorphs. Elliptical motion was obtained from the output tips of the bimorphs, and when combined together, sliding motion of up to 18.1 cm/s velocity and 150 mN force could be obtained. The best configuration according to finite element analysis, {lE1 , lE2 } = {4.5, 4.5} mm and a tip length l = 1.5 mm, was also the best configuration of the three tested experimentally. Unfortunately, the efficiency was low throughout the operating range of the slider, probably due to the extremely low apparent contact area between the thin bimorph and the slider. In the future, increasing the number of bimorphs in the actuator, and placing them on both sides of the slider will be explored. Further, modifications to the tip of the bimorph by folding the tip or using a plate with a thicker tip would increase the tip’s width along the axis of sliding and may increase its lateral stiffness. Those changes could improve the actuator’s efficiency while avoiding the phase problems along the tip.

Acknowledgements The authors appreciate the support of this work by Taiheyo cement and a grant-in-aid from the Japanese Ministry of Education, Culture, Sports, and Science.

References

Fig. 15. Measured sliding force and efficiency versus sliding velocity for three different values of applied preload force. The lines in the sliding force versus velocity plot are least square fits of the data for each preload.

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Biographies James Friend was born in Lubbock, Texas on 13 September 1970. He received the BS degree in aerospace engineering, and the MS and PhD degrees in mechanical engineering from the University of Missouri, Rolla in 1992, 1994, and 1998, respectively. He received the Best Paper award and Jefferson Goblet Student Paper Award from the ASME and AIAA for a paper delivered at the 97th Annual AIAA/ASME/AHS/ASC/ASCE Structural Dynamics and Mechanics Conference in 1996. In 1999, he joined the faculty at the newly-formed department of mechanical engineering at the University of Colorado, Colorado Springs. He is now at the Precision and Intelligence Laboratory, Tokyo Institute of Technology, as a research associate, with research interests in millimeter- and micrometer-scale piezoelectric actuators and their applications. He is a member of IEEE, ASME, and the Acoustical Society of Japan, and associate member of Sigma Xi. Akira Umeshima was born in Kanagawa Prefecture, Japan, on 10 October 1977. He received the BEng and the MEng from the Tokyo Institute of Technology, Tokyo, Japan, in 2000 and 2002, respectively. Mr. Umeshima is now at Kawasaki Heavy Industries’ Kobe Works in Kobe, Japan. Takaaki Ishii was born in Chiba, Japan, on December 7, 1964. He received the BSc degree in 1987, the MSc degree in 1990 both in physics from Sophia University, Tokyo, and the DEng degree in 2000 from the Tokyo Institute of Technology. In 1988, he was a Visiting Research Assistant of the Materials Research Laboratory at The Pennsylvania State University, USA, doing research on ultrasonic motors. From 1990 to 1993, he was an engineer for ALPS Electric Co., Ltd., in Niigata, Japan, engaged in research into piezoelectric ceramics and ultrasonic motors. He was a Research Associate of the Precision and Intelligence Laboratory at the Tokyo Institute of Technology from 1994 to 2002, working on ultrasonic motors, wear evaluation of friction materials, piezoelectric actuators and other ultrasonic devices. He has been a Research Associate of the Interdisciplinary Graduate School of Medicine and Engineering, University of Yamanashi since 2002. Dr. Ishii is a member of The Acoustical Society of Japan and the Japanese Society of Tribologists. Kentaro Nakamura was born in Tokyo, Japan, on 3 July 1963. He received the BEng, the MEng, and the DEng degrees from the Tokyo Institute of Technology, Tokyo, Japan, in 1987, 1989, and 1992, respectively. He has been an Associate Professor of the Precision and Intelligence Laboratory, Tokyo Institute of Technology, since 1996. His field of research is the application of ultrasonics and the measurement of vibration and sound using optical methods. He has received the Awaya Kiyoshi Award for encouragement of research from the Acoustical Society of Japan in 1996. Dr. Nakamura is a member of the Acoustical Society of Japan, the Japan Society of Applied Physics, the Institute of Electrical Engineers of Japan, and the Institute of Electronics, Information and Communication Engineers. Sadayuki Ueha was born in Kyoto Prefecture, Japan, on 28 February 1943. He received the BEng degree in electronic engineering from the Nagoya Institute of Technology in 1965 and the MEng degree in 1967, and the DEng degree in 1970, both in electric engineering, from the Tokyo Institute of Technology. He currently conducts research in high power ultrasonics. He has been a Professor of the Precision and Intelligence Laboratory, Tokyo Institute of Technology since 1992. He is a steering committee member of the World Congress on Ultrasonics and serves as the secretariat of WCU97. He received the Best Paper Award from The Japan Society of Applied Physics in 1975 and from the Acoustical Society of Japan in 1980, respectively. Dr. Ueha is a member of the Japan Society of Applied Physics, the Acoustical Society of Japan, the Institute of Electronics, Information and Communication Engineers and the Japan Society of Ultrasonics in Medicine. Currently, he is a member of the editorial board of the journal “Ultrasonics”.