Pushing the Limits for Microactuators Based on ... - IEEE Xplore

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JOURNAL OF MICROELECTROMECHANICAL SYSTEMS, VOL. 21, NO. 3, JUNE 2012

Pushing the Limits for Microactuators Based on Electroactive Polymers Babita Gaihre, Gursel Alici, Geoffrey M. Spinks, and Julie M. Cairney

Abstract—We have previously reported on the fabrication and displacement output of electroactive polymer (EAP) microactuators less than 1 mm in length. The main limiting factor hindering their further miniaturization and their displacement output was the thickness of the commercially available polyvinylidene fluoride (PVDF) membrane used (∼ 110 μm). In this study, we have reduced the thickness of the PVDF layer using a spin-coating technique and then electrochemically deposited polypyrrole layers on both sides of this thin film to make ultrathin-film EAP substrates with a thickness of 48 μm. We then employed a laser ablation technique to fabricate microsized EAP actuators as small as 200 μm in length and 50 μm in width that can operate in both dry and aqueous media. This is the minimum-size EAP microactuator to be reported in the literature. Based on the operation principle of these actuators, we model them as a microcantilever beam under a uniformly distributed load. We then establish bending displacement and blocking force models to perform the following: 1) to estimate the actuation force, actuation moment, tip deflection, flexural rigidity, and strain energies per unit volume and mass for a set of microactuators as big as 850 μm × 250 μm × 126 μm and as small as 200 μm × 50 μm × 48 μm and 2) to evaluate their performance metrics. [2011-0250] Index Terms—Electroactive polymer (EAP) microactuators, microforce measurement, modeling and identification, performance evaluation and characterization, polyvinylidene fluoride (PVDF) thin film.

I. I NTRODUCTION

C

ONJUGATED or conducting polymers have recently been receiving significant attention as smart materials for actuators and sensors for novel microfabricated devices [1], [2]. These smart materials are distinguished by the presence of alternating single and double bonds between carbon atoms along the polymer backbone, which result in a bandgap, making them semiconducting. The p-doping can be accomplished electrochemically, and the charge on the polymer in the doped site is delocalized along the backbone because of the pi electron

Manuscript received August 23, 2011; revised December 22, 2011; accepted December 29, 2011. Date of publication February 6, 2012; date of current version May 28, 2012. This work was supported by Australian Research Council Discovery Project DP0878931. Subject Editor E. S. Kim. B. Gaihre, G. Alici (corresponding author), and G. M. Spinks are with the School of Mechanical, Materials and Mechatronic Engineering, University of Wollongong, Wollongong, N.S.W. 2522, Australia (e-mail: [email protected]; [email protected]; [email protected]). J. M. Cairney is with the Australian Centre for Microscopy and Microanalysis, The University of Sydney, Sydney, N.S.W. 2006, Australia (e-mail: [email protected]). 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/JMEMS.2012.2184084

system. Unlike in silicon, in conjugated polymers, the doping level is reversible and can be controlled. Among the several conducting polymers, polypyrrole (PPy) has gained particular attention due to its properties such as biocompatibility, biodegradability, ease of synthesis, low actuation power, ability to work in liquid and air environments, and stability with large volume changes [2], [3]. The latter is particularly important for electroactive actuators, whose operation principle is based on the volume expansion and/or contraction generated by the movement of ions during an electrochemical reaction [2]. Fabrication of bilayer microactuators based on conjugated polymers has been extensively studied by Smela [1], [2], Otero and Cortes [3], and Jager et al. [4], [5]. However, these bilayer actuators could only operate in specific aqueous media because the media provide the source of ions (electrolytes) required for actuation. Typically, the actuator is immersed in a liquid electrolyte along with reference and counter electrodes in a three-electrode operation. Electrochemically induced strain in the conjugated polymer film causes the bilayer to bend like a cantilever beam. Although bilayer actuators are the most frequently reported actuators in the literature, a new type of actuator has recently emerged, i.e., trilayer [6], [7]. This type of actuator is prepared by separating two electroactive polymer (EAP) films by an insulating soft and porous film. The middle insulating film acts as an ion reservoir. The two outer electroactive layers can be electrochemically oxidized and reduced. When one layer is oxidized, the other layer is reduced, and the oxidation and reduction of the layer can be done in a continuous and reversible way, which leads to a rocking-chair motion. The ions are shuttled between the two conjugated polymer films through a middle insulating layer during the electrochemical switching process to cause the differential volume change required to generate a bending motion from an actuator strip with one end fixed and the other end free to deflect. The most significant advantage of this type of actuator reported in this study over a bilayer type is that it can operate in both dry and wet environments, which widens the potential applications. Such dry systems have already been demonstrated on a macroscopic scale. Kaneto et al. demonstrated the first actuator with a rocking-chair motion that could operate for a short time in air [8]. Since then, significant research efforts have been dedicated to these actuators [6], [9]–[12]. However, until recently, none of them has been dedicated toward their miniaturization. Small-scale conducting-polymer actuators are useful in numerous applications, including the micromanipulation of living cells, bioanalytical nanosystems, data storage, laboratory-on-chip, microvalves, microswitches, microshutters,

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GAIHRE et al.: PUSHING THE LIMITS FOR MICROACTUATORS BASED ON ELECTROACTIVE POLYMERS

cantilever light modulators, micro-optical instrumentation, artificial muscles for macro-/microrobotics, and more [2], [5], [7] and [13]. With these applications in mind, there is an increasing need for microsized EAP actuators that can operate in both dry and wet media. Templating the shape of a conjugated polymer during synthesis is the most common approach to actuator fabrication in microelectromechanical systems (MEMS) microactuation technology. However, the use of this technique for the fabrication of trilayer actuators involves various tedious steps in depositing three layers one above another. The use of different chemicals for masking and demasking and for removing unwanted materials might cause reduction in performance of the actuators. Furthermore, exact alignment of these layers is very challenging for microscale devices. An alternative to this templating technique is micromachining, in which the trilayer is synthesized as a bulk material and then fabricated into the desired shape/structure by laser irradiation or dry plasma etching processes. Our group has used this technique for the miniaturization of the trilayer actuators with some promising results [7]. The main limiting factor for further miniaturization of those actuators has been the thickness of the polyvinylidene fluoride (PVDF) layer, which acted as the middle layer cell separator and ion reservoir. In previous work, we were successful in reducing the thickness of the PVDF layer by using a spincoating technique [14]. The trilayer actuator was then prepared by the deposition of PPy layers on both surfaces of the PVDF thin film. We had reduced the thickness of the PVDF layer to 32 μm and deposited 8-μm PPy layers on each side of the PVDF layer, providing an overall actuator thickness of 48 μm, which is 3.5 times thinner than our previously reported microactuators [7]. The actuators were fabricated to an 850 μm × 250 μm size using a laser ablation technique, and this previous study reports their displacement output results only [15]; no experimental blocking force results and no mathematical models estimating the blocking force and the deflection were reported before. Along with the tip deflection, the blocking force is the other important parameter to be measured to quantify the performance of these microsized EAP actuators. In this paper, we report on the blocking force and the tip displacement of three different types of microactuators of five different sizes. The three different types of actuators that are investigated utilize three different types of PVDF middle layers for the electrolyte storage. The actuator produced using the commercially available thick PVDF membrane is termed the “thick” actuator, and the actuators made in-house of thin porous and thin ionincorporated PVDF films are termed “thin porous” and “thin ion-incorporated” actuators, respectively. All of these actuators, with a length of 850 μm and a width of 250 μm, showed voltage-dependent force and displacement. While the blocking force was a maximum for the thick actuator, its tip displacement was the lowest. While the force decreased drastically for a porous actuator, its tip displacement was greater than that of the thick actuator. The force and displacement outputs were improved for the ion-incorporated actuator compared to a porous actuator. This study is the first time that the force output of a dry EAP microactuator operating in air has been reported.

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We modeled the behavior of these actuators as a cantilever beam under a uniformly distributed load and used the deflection and blocking force models to identify the flexural rigidity of the actuators, the actuation force, the bending moment, and the strain energy per unit volume, generating the mechanical outputs under a voltage input ranging from 0.1 to 0.6 V with steps of 0.1 V. It was found that the strain energies per unit volume and mass were comparable to those of mammalian muscles. It was also found that, the smaller the actuator volume, the higher the strain energy injected into the actuators for the mechanical output as a result of the electrochemical processes of oxidation and reduction. Compared to other microactuators such as piezoelectric microactuators [7], [23] and other microand nanocantilevers [24], [25], the speed of response and the blocking force of the EAP microactuators in the study are lower, but the tip deflection is relatively higher. However, the most notable feature of the EAP microactuators over other microactuators is their low-power operation and low footprint. The primary contribution of this study is to report on the fabrication, modeling, and performance evaluation of the first dry-type EAP microactuators operating in air. Laser ablation was used to fabricate the thick ion-incorporated actuators of five different sizes (1050 μm × 50 μm, 850 μm × 50 μm, 650 μm × 50 μm, 450 μm × 50 μm, and 200 μm × 50 μm). This technique involves the use of a high-powered pulsed laser to remove unwanted materials from a substrate. The force measurement was done using silicon-based microforce sensors. The voltage-dependent tip displacement of the smallest ionincorporated actuator (200 μm × 50 μm) was also measured, and its blocking force was estimated using an analytical force model and verified experimentally. This is the smallest or the minimum-volume dry-type ionic conducting-polymer microactuator reported so far. II. E XPERIMENTAL W ORK A. PVDF Thin-Film Preparation and Gold Sputter Coating The PVDF powder (Aldrich) films were prepared from a 10% PVDF solution in dimethylformamide [(DMF); Sigma-Aldrich] in the presence of different additives [1% salicylic acid, Ajax chemicals, or 0.025-M lithium triflouromethanesulfonimide (LiTFSi)]. The film prepared by adding salicylic acid is termed herein as the porous film, and the film prepared by adding LiTFSi is referred to as the ion-incorporated film. A PVDF film without any additive was also prepared for comparison, which is termed the nonporous film. The film preparation procedure involved stirring the PVDF solution (with or without additives) overnight followed by spin coating (500 r/min; 30 s) on a glass slide and drying at 50 ◦ C [14]. For the preparation of the porous film, the film was further heated to 200 ◦ C for 30 min to remove the salicylic acid. These porous and nonporous PVDF films were compared to a commercially available PVDF membrane (110 μm thick from Millipore). Gold was sputter coated (magnetron sputter coater SC 100MS) on both sides of the PVDF film/membrane at 2 × 10−3 mbar pressure and 30-mA current. The coated layers of gold serve to provide a conductive surface onto which the PPy

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Fig. 1. Photomicrographs of the (a) thick, (b) porous, and (c) 0.025-M LiTFSi films.

layers can be electrodeposited. The coating times for the thin film and PVDF membrane were 2 and 20 min, respectively. B. Trilayer Actuator Preparation The actuators were synthesized using a galvanostatic polymerization method, which involves the electrodeposition of PPy layers on both surfaces of the gold-coated films (acting as a working electrode) by passing a constant current through a solution containing 0.1-M pyrrole (Merch) dissolved in 0.1-M LiTFSi/propylene carbonate [(PC); Aldrich] (with 1% water). The polymerization was conducted for 4 h at −33 ◦ C and a current density of 0.1 mA · cm−2 . Stainless steel mesh was used for two counter electrodes positioned on both sides and parallel to the gold-coated PVDF film/membrane. A potentiostat/galvanostat (EG&G Princeton Applied Research Model 363) was used to generate the constant current. Before the start of the polymerization, the solution was degassed by passing nitrogen gas for 30 min followed by cooling for 30 min. After polymerization, the actuators were rinsed with acetone and soaked with 0.1-M LiTFSi/PC solution. C. Fabrication of the Microactuators The trilayer actuators were fabricated to the required size by using a laser cutting system [JDSU Q-series Q-201HD (Q-switched doubled Nd:YAG laser system)]. The ablation conditions for the fabrication of thick actuators were as follows: wavelength = 266 nm (UV), repetition f requency = 600 Hz, pulse duration = 15 ns, average power = 20 mW, pulse energy = 33 μJ, spot size = 6 μm, f luence = 120 J/cm2 , and seven passes at 10 mm/min. The conditions were set so that the laser cannot damage the membrane. The ablation conditions for the fabrication of thin (made of thin PVDF membrane) actuators were comparably milder than

the conditions set for the thick actuators. The repetition rate, fluence, and average power were 500 Hz, 70 J/cm2 , and 5–10 mW, respectively. The rest were the same as those of the thick actuators. The microactuators were cut to the desired size leaving an area of 2 mm × 1 mm at one end of the cantilever actuator for an electrical connection. D. Evaluation of the Electromechanical Performance of the Microactuators The voltage- and frequency-dependent tip displacements of the microactuators were measured using a noncontact laser displacement sensor (Micro-Epsilon NCDT-1700-10). The force measurement was done using silicon-based microforce sensors (FemtoTools; FT-S270 and FT-S540). III. E XPERIMENTAL R ESULTS AND D ISCUSSION A. Surface Morphology of the Films Fig. 1 shows the photomicrographs of the films. The PVDF films were prepared by spin coating of 10% PVDF solution in DMF. In order to make the film porous, 1% salicylic acid was added to the spinning solution. Decomposition of salicylic acid at 183 ◦ C and evaporation of the resulting degradation product at 200 ◦ C have been previously found to result in a porous structure in the films [14], [15] [Fig. 1(b)]. Alternatively, in order to incorporate a salt in the film, LiTFSi was added to make 0.025-M LiTFSi in the 10% PVDF solution [15]. The resulting solution was homogeneous, and since the film was heated only to 50 ◦ C, it is reasonable to think that, with the addition of LiTFSi in the spinning solution, the salt was entrapped in the film. Microscopic examination of the films [Fig. 1(c)] shows that the surface of the ion-incorporated film was rougher than that of the thick film, possibly due to the salt ions trapped and


Fig. 2.

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(a) Optical table and (b) machining table used for the fabrication.

microactuator. There are three main layers in the structure: two outer black PPy layers, which are the electroactive layers, and an inner porous separator of the PVDF film that holds the liquid electrolyte. The gold particles are not visible in this image. When an opposite voltage is applied to the two outer layers, one layer contracts, and the other layer expands; moreover, mechanical constraints arising due to the middle electroinactive PVDF layer result in an overall bending of the actuator to one side. C. Force and Displacement Measurements Fig. 3.

(a) Surface and (b) cross-sectional images of a thin microactuator.

retained in the films. The film was more opaque than the thick and porous films. B. Laser Ablation The films and membrane were then sputter coated with gold on both surfaces, followed by electrochemical deposition of PPy layers over the gold layers for 4 h. This resulted in the deposition of 8 μm of PPy film on each side, making an overall thickness of 48 and 126 μm for the thin and thick actuators, respectively. Laser ablation was used to fabricate the actuators to the desired shape. A direct-write process was used to cut the actuator, which involved using a programmable, high-resolution, and high-stability translation stage on which the substrate rests. The laser system was externally frequency doubled on an optical table [Fig. 2(a)] to produce 266-nmwavelength (UV) light. This was performed with a single pass in a BBO crystal. The laser beam is delivered onto a set of XY Z stages, which are Aerotech ATS100 models having 200 mm of travel with ∼ 0.5-μm resolution [Fig. 2(b)]. The low pulse-repetition rate and moderate powers were used to reduce heat accumulation in the material. Fig. 3(a) shows a 200 μm × 50 μm × 48 μm microactuator ablated with 5-mW laser power and 300-Hz repetition frequency. No burning and peeling out of the layers occurred at this power. Fig. 3(b) shows a cross-sectional image of the thin

The free displacement and the blocking force are two important parameters for the evaluation of electromechanical performance of an actuator. The free displacement without any mechanical load on the tip was measured by a laser displacement sensor. An actuator was clamped on a stand, and the laser beam was focused on the tip of the actuator. When opposite voltages were applied to the two outer PPy layers through electrical contacts, the tip of the actuator bent to one side, as shown in Fig. 4(a). For the force measurement experiment, the tip of the force sensor was placed in contact with the end of the trilayer actuator, as shown in Fig. 4(b). The actuator bent and pressed on the force sensor, and the output was an analog voltage. The measured force was then calculated from F = sensitivity × V , where F is the force axially applied to the sensor probe and V is the output voltage. MEMS fabricated single-crystalline silicon acted as the sensing element of the microforce sensor, which uses the principle of capacitive deflection measurement. It was not possible to record the blocking force while a microactuator was in motion due to the physical constraints associated with the force and deflection measurement setups. We were able to measure the blocking force where the deflection was zero and the maximum deflection where the blocking force was zero. Fig. 4(b) shows a schematic diagram of the force and displacement measurement setups. An external function generator (FG 601) was used to generate the input voltage signal with a specified frequency. The input was amplified using an eDAQ potentiostat (eDAQ

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Fig. 4. Schematic representations of the (a) free displacement and (b) blocking force measurement setups. (c) Photograph of the measurement setup. The labels CE, RE, and WE represent the counter, reference, and working electrodes, respectively.

EA161) operating in a two-electrode mode. A datalogger (ecorder, model ED821), which is an interface unit between the computer, the force sensor, and the potentiostat, recorded the voltage signal applied to the trilayer, the associated current drawn, and the voltage signal from the force sensor. The force and displacement experiments were carried out for the three types of actuators. Fig. 5(a) shows the variation of the blocking force with the voltage for the thick, porous, and ionincorporated actuators with the dimensions of 850 μm in length, 250 μm in width, and 126, 48, and 48 μm in thickness, respectively. In all cases, the blocking force was found to be proportional to the input voltage. A voltage-dependent force increase was also observed in our previous study on macrosized EAP actuators [17]. Furthermore, a dramatic decrease in the blocking force was observed with a reduction in thickness from 126 to 48 μm. For comparison, the thick actuator, the porous actuator, and the LiTFSi-incorporated actuator generated blocking forces of 468, 33, and 124 μN, respectively, under an input voltage of 0.6 V. It must be noted that, although the porous actuator and the ion-incorporated actuator had the same thickness, the ion-incorporated actuator generated approximately four times the blocking force. This increased force can be explained by calculating the charge stored in each actuator under the same input voltage. The amount of charge that passed through the actuator during force measurement at various voltages was also calculated, and the result is shown in Fig. 5(b). These

data indicate that the amount of charge that passed is linearly proportional to the voltage difference. The charge that passed is directly related to the movement of ions in and out of the active polymer layers, leading to the contraction and expansion of the relevant polymer layers [16], [17]. A higher voltage leads to a higher charge that passed through the actuators, resulting in more volume change and, hence, a higher blocking force. The large force observed in the thick actuator can be correlated to a higher flexural rigidity and a higher charge that passed to the actuator. With the reduction of the thickness, the modulus of rigidity of the film also decreased; hence, the force of the thin actuators also decreased compared to the thick one. Between the two films, the one with incorporated ions showed a higher blocking force compared to the porous one. In our previous study, we have shown that the ion-incorporated film has greater conductivity than the porous one [15]. The results indicate that the movement of ions in the film contributes to the higher force in the ion-incorporated actuator. The blocking force also changes with the frequency of the square-wave input, as shown in Fig. 5(c). As explained earlier, the force depends on the movement of ions in and out of the PPy layers during the redox process. When the frequency of the input signal was increased, the ions are stopped in the middle and forced back to change the direction of movement without giving them sufficient time to reach deep into the PPy layer to generate more force [17]. Fig. 5(d) shows a typical step response recorded for


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Fig. 5. (a) Voltage-dependent blocking force at 0.01 Hz. (b) Charge at 0.01 Hz. (c) Frequency-dependent blocking force at 0.6 V of the microactuators (850 μm × 250 μm). (d) Typical step response chart during blocking force measurement at 0.01 Hz for the ion-incorporated actuator.

Fig. 6. Blocking force versus free displacement of the microactuators (850 μm × 250 μm) at 0.005 Hz. 1, 2, 3, 4, and 5 represent the results obtained at 0.2, 0.3, 0.4, 0.5, and 0.6 V, respectively.

the ion-incorporated actuator with the dimensions of 850 μm × 250 μm × 48 μm. The current that passed through, the input voltage, and the blocking force of the actuator are shown in the first, second, and third rows, respectively. Fig. 6 shows the force-versus-displacement curves for the thick, porous, and ion-incorporated actuators at various input voltages. Every actuator showed a voltage-dependent force as

well as displacement. The thick actuator had a higher force but a lower displacement compared to the porous and ionincorporated actuators. With the inclusion of LiTFSi ions in the film, an improvement in both the force and the displacement was observed in the thin actuators. The larger tip displacement in the thin porous and ion-incorporated actuators is expected because the bending stress for the thin actuator is much higher compared to its thicker counterpart due to a smaller area moment of inertia (3.80 × 10−6 mm4 ) of the thin actuator compared to the thicker one (53.60 × 10−6 mm4 ). A detailed explanation on the calculation of bending stress and area moment of inertia for the actuators can be found in our previous work [14], [15]. With the inclusion of LiTFSi ions in the film, the ionic conductivity of the thin film was increased [15]. Since the bending stress remains the same for all of the thin actuators, it appears that the ionic conductivity has played a primary role in the displacement of the ion-incorporated actuator. Fig. 7(a) shows a typical step response recorded for the ionincorporated actuator with dimensions of 200 μm × 50 μm × 48 μm. The input voltage, the current that passed through in and out, and the tip-to-tip displacement of the actuator are shown in the first, second, and third rows in Fig. 7(a). Fig. 7(b) shows the voltage-dependent tip displacement of the actuator. The actuator showed a maximum tip-to-tip displacement of 74 μm at ±0.5 V and 0.005 Hz. The tip displacement decreased

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Fig. 7. (a) Step response, (b) voltage-dependent displacement, and (c) charge at 0.005 Hz and (d) frequency-dependent displacement at 0.6 V of the ion-incorporated actuator of size 200 μm × 50 μm × 48 μm.

with decreasing voltage. Fig. 7(c) shows the charge that passed through the actuators during the displacement measurement. As discussed earlier, the shrinkage and expansion of the PPy layer depend on the movement of ions in and out of the polymer layers during the redox process. Higher voltage indicates a higher charge that passed, i.e., larger number of ions passing in and out of the polymer layers, causing more shrinkage and expansion of the PPy layers, resulting in a higher displacement. The variation of the bending displacement of this actuator with the frequency of the input signal is shown in Fig. 7(d). The explanation for this behavior has been presented before for the larger microactuators. The blocking force for this microactuator could not be measured using the same setup used for the larger microactuators as the actuator and the force sensor were too short to make contact for the force measurement. However, as will be presented in the next section, we have employed an experimentally verified blocking force model to estimate the blocking force. IV. P ERFORMANCE Q UANTIFICATION OF M ICROACTUATORS The microactuators are cantilever beams with one end fixed and the other end free under a uniformly distributed load, as shown in Fig. 8.

Fig. 8.

Schematic of an analogous beam model for the microactuators.

The blocking force FB of such a beam is given by [18], [19] FB =

3qL 8

(1)

where q is the internal force per unit length. It is known that the steady-state blocking force is proportional to the input voltage [17] FB = αV.

(2)

The proportionality constant α is identified experimentally from the blocking-force-versus-input-voltage measurements. The blocking force and the proportionality constant α were determined for the microactuator with dimensions of 850 μm × 250 μm × 126 μm (the thick actuator) and are shown in the top plot in Fig. 9.


Fig. 9. (Top) Variation of the blocking force with the input voltage, (middle) the predicted deflection, and (bottom) the predicted blocking force.

Combining (1) and (2) results in q=8

αV . 3L

(3)

The tip deflection δ of the microactuator is expressed by δ=

qL4 8EI

(4)

where EI is the flexural rigidity of the microactuators. It is not straightforward to calculate EI for the composite EAP actuators as the moduli of elasticity of the PPy and PVDF layers change with the size and type of ions, solvent, input voltage, and size of the actuators [17], [19]–[21]. It is therefore necessary to determine the flexural rigidity experimentally when possible. Using the experimental data obtained for the microactuator with dimensions of 850 μm × 250 μm × 126 μm, the internal force per unit length and the flexural rigidity are identified and are reported in Table I. The maximum internal bending moment Minternal and bending force Finternal due to the uniformly distributed load of q are calculated from Minternal = q

L2 2

Finternal = qL.

(5)

The work done on the actuator is converted into strain or potential energy to deflect the actuator or generate the actuator stroke. After doing necessary mathematical manipulations, the strain energy generating this mechanical output is given by [22] U=

q2 L5 . 40EI

(6)

The strain energies per unit volume and unit mass (m = 26.99 μg for the microactuator with the dimensions of 850 μm × 250 μm × 126 μm) are calculated and provided in Table I.

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As shown in Table I, the actuation force per unit length and the flexural rigidity (due to change in the modulus of elasticity) are functions of the input voltage. Another set of experiments was conducted to verify q and EI. The new set of experimental deflection and blocking force data is depicted in Table II. The strain energy per unit volume and the strain energy per unit mass have been calculated and are depicted in Table II for comparison with the data in Table I. Equations (1)–(4) are employed to estimate the deflection and blocking force of another thick microactuator with the same dimensions. The estimated results and experimental results are shown in the bottom two plots in Fig. 9. To further verify the blocking force and deflection models, the deflection and blocking force data were obtained for an ionincorporated actuator with dimensions of 850 μm × 250 μm × 48 μm. These data and the actuator parameters identified and calculated are presented in Table III. The strain energies per unit volume and unit mass (m = 16.15 μg for the microactuator with the dimensions of 850 μm × 250 μm × 48 μm) are calculated and are also provided in Table III. Another set of experiments was conducted to verify q and EI for this thinner microactuator. The new set of experimental deflection and blocking force data is shown in Table IV. The strain energy per unit volume and the strain energy per unit mass have been calculated and are also depicted in Table IV for comparison with the data in Table III. Equations (1)–(4) are employed to estimate the deflection and blocking force of another thick microactuator with the same dimensions. The predicted results and experimental results are shown in the middle and bottom plots in Fig. 10. It was not possible to measure the blocking force for the microactuators with the dimensions of 200 μm × 50 μm × 48 μm due to physical constraints associated with the force measurement system used. Assuming that the moduli of elasticity of the actuators with the same thickness are approximately the same, we can calculate the flexural rigidity and q for the equivalent cross section of this smallest microsized actuator. Using the EI data in Table III, EI is scaled down by the width ratio (1/5). The internal bending force is proportional to the cross section of each microactuator, and under the same applied voltage (Finternal )1 A1 = (Finternal )2 A2 ⇒ q2 = q1

L1 w1 t1 L2 w2 t2 (7)

where Li , wi , and ti are the length, width, and thickness of the actuators, respectively. For the microactuators with dimensions of 850 μm × 250 μm × 48 μm and 200 μm × 50 μm × 48 μm, (7) reduces to q2 = 5q1

L1 . L2

(8)

Using the deflection data for the microactuator with the dimensions of 200 μm × 50 μm × 48 μm, the blocking force and the strain energies per unit volume and unit mass (m = 0.76 μg for the microactuator with the dimensions of 200 μm × 50 μm × 48 μm) are calculated. The estimated force and deflection data are shown in Fig. 11. These results show that the blocking force and deflection models and the way we estimate

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TABLE I E XPERIMENTAL T IP D EFLECTION AND B LOCKING F ORCE DATA U SED TO I DENTIFY THE PARAMETERS IN THE D EFLECTION AND B LOCKING F ORCE M ODELS FOR THE M ICROACTUATOR W ITH THE D IMENSIONS OF 850 μm × 250 μm × 126 μm

TABLE II S TRAIN E NERGIES C ALCULATED FOR THE M ICROACTUATOR W ITH THE D IMENSIONS OF 850 μm × 250 μm × 126 μm

TABLE III E XPERIMENTAL T IP D EFLECTION AND B LOCKING F ORCE DATA U SED TO I DENTIFY THE PARAMETERS IN THE D EFLECTION AND B LOCKING F ORCE M ODELS FOR THE M ICROACTUATOR W ITH THE D IMENSIONS OF 850 μm × 250 μm × 48 μm


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TABLE IV S TRAIN E NERGIES C ALCULATED FOR THE M ICROACTUATOR W ITH THE D IMENSIONS OF 850 μm × 250 μm × 48 μm

250 μm × 48 μm. It follows that, the smaller the EAP microactuators are, the higher the strain energies are per unit volume and mass. V. C ONCLUSION AND F UTURE W ORK

Fig. 10. (Top) Variation of the blocking force with the input voltage, (middle) the predicted deflection, and (bottom) the predicted blocking force.

Fig. 11. (Top) Variation of the predicted blocking force with the input voltage and (bottom) the predicted and experimental deflection data.

the modulus of elasticity and the internal force per unit length q are accurate enough to predict the bending performance of these microactuators, the smallest EAP microactuators reported so far. With reference to the data in Table V, the strain energies per unit volume and unit mass of this actuator are also calculated. These values are clearly much higher than that of the microactuators with the dimensions of 850 μm ×

Thin-film microactuators were fabricated, and their performance has been evaluated in terms of their displacement and force outputs. A lower laser power gave a good cut for the thin actuator while a higher laser power was required to cut the thicker actuator, which was more than three times as thick as and more conductive than the thin microactuator. The tip displacement of a porous actuator was larger than that of the thick actuator under identical conditions, although the conductivity of the thicker one was higher. However, the force decreased drastically in the porous actuator. Inclusion of LiTFSi salts in the film increased not only the displacement but also the force of the resulting actuators, and these data were dependent on both the voltage and the frequency. We have modeled the microactuators as cantilever beams under a uniformly distributed load, and their bending displacement and blocking force models are employed to estimate the actuation force, actuation moment, tip deflection, flexural rigidity, strain energy per unit volume, and strain energy per unit mass, generating the mechanical output. The models are further verified using a new set of experimental data. For comparison, the actuation performances of the three microactuators with sizes of 850 μm × 250 μm × 126 μm, 850 μm × 250 μm × 48 μm, and 200 μm × 50 μm × 48 μm are evaluated in terms of the same performance metrics. The results show that, the smaller the actuator volume, the higher the strain energy injected into the actuators for the mechanical output as a result of the electrochemical processes of oxidation and reduction. This is the first time, to the best of the authors’ knowledge, that EAP microactuators as small as 200 μm × 50 μm × 48 μm have been fabricated. Their actuation performance has been quantified, and their deflection and blocking force are modeled and accurately predicted. We conclude that, the smaller the EAP microactuators are, the higher the strain energies per unit volume and mass and their blocking force under the same applied voltage are. Future work involves further characterization of the ionincorporated microactuators including their lifetime over many experiments, expanding the modeling approach such that it more accurately describes the relationship between the motion at the molecular level (i.e., ion transport) and the macrolevel mechanical motion.

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TABLE V S TRAIN E NERGIES C ALCULATED FOR THE M ICROACTUATOR W ITH THE D IMENSIONS OF 200 μm × 50 μm × 48 μm

ACKNOWLEDGMENT This work was performed in part at the OptoFab Node of the Australian National Fabrication Facility at Macquarie University. R EFERENCES [1] E. Smela, “Microfabrication of PPy microactuators and other conjugated polymer devices,” J. Micromech. Microeng., vol. 9, no. 1, pp. 1–18, Mar. 1999. [2] E. Smela, “Conjugated polymer actuators for biomedical applications,” Adv. Mater., vol. 15, no. 6, pp. 481–494, Mar. 2003. [3] T. F. Otero and M. T. Cortes, “A sensing muscle,” Sens. Actuators B, Chem., vol. 96, no. 1, pp. 152–156, Nov. 2003. [4] E. W. H. Jager, E. Smela, O. Inganas, and I. Lundstrom, “Polypyrrole micro actuators,” Synth. Met., vol. 102, no. 1–3, pp. 1309–1310, 1999. [5] E. W. H. Jager, O. Inganas, and I. Lunstrom, “Microrobots for micrometersize objects in aqueous media: Potential tools for single cell manipulation,” Science, vol. 288, no. 5475, pp. 2335–2338, Jun. 2000. [6] J. M. Sansiñena, V. Olazábal, T. F. Otero, C. N. Polo da Fonseca, and M.-A. De Paoli, “A solid state artificial muscle based on polypyrrole and a solid polymeric electrolyte working in air,” Chem. Commun., no. 22, pp. 2217–2218, 1997. [7] G. Alici, V. Devaud, P. Renaud, and G. Spinks, “Conducting polymer microactuators operating in air,” J. Micromech. Microeng., vol. 19, no. 2, pp. 025017–025025, Feb. 2009. [8] K. Kaneto, M. Kaneko, Y. Min, and A. G. MacDiarmid, “Artificial muscle: Electromechanical actuators using polyaniline films,” Synth. Met., vol. 71, no. 1–3, pp. 2211–2212, Apr. 1995. [9] M. T. Cortés and J. C. Moreno, “Artificial muscles based on conducting polymers,” e-Polymers, vol. 41, pp. 1–42, 2003. [10] S. A. Wilson, R. P. J. Jourdain, Q. Zhang, R. A. Dorey, C. R. Bowen, M. Willander, Q. U. Wahab, M. Willander, S. M. Al-hilli, O. Nur, E. Quandt, C. Johansson, E. Pagounis, M. Kohl, J. Matovic, B. Samel, W. van der Wijngart, E. W. H. Jager, D. Carlsson, Z. Djinovic, M. Wegener, C. Moldovan, R. Iosub, E. Abad, M. Wendlandt, C. Rusu, and K. Persson, “New materials for μ-scale sensors and actuators: An engineering review,” Mater. Sci. Eng. R, Rep., vol. 56, no. 1–6, pp. 1–129, 2007. [11] F. Vidal, C. Plesse, G. Palaprat, A. Kheddar, J. Citerin, D. Teyssie, and C. Chevrot, “Conducting IPN actuators: From polymer chemistry to actuator with linear actuation,” Synth. Met., vol. 156, no. 21–24, pp. 1299– 1304, Dec. 2006. [12] A. S. Hutchison, T. W. Lewis, S. E. Moulton, G. M. Spinks, and G. G. Wallace, “Development of polypyrrole-based electromechanical actuators,” Synth. Met., vol. 113, no. 1, pp. 121–127, Jun. 2000. [13] Y. Fang, X. Tan, Y. Fang, and X. Tan, “A novel diaphragm micropump actuated by conjugated polymer petals: Fabrication, modeling, and experimental results,” Sens. Actuators A, Phys., vol. 158, no. 1, pp. 121–131, 2010. [14] B. Gaihre, G. Alici, G. M. Spinks, and J. M. Cairney, “Synthesis and performance evaluation of thin film PPy–PVDF multilayer electroactive polymer actuators,” Sens. Actuators A, Phys., vol. 165, no. 2, pp. 321– 328, Feb. 2011. [15] B. Gaihre, G. Alici, G. M. Spinks, and J. M. Cairney, “Effect of electrolyte storage layer on performance of PPy–PVDF–PPy microactuators,” Sens. Actuators B, Chem., vol. 155, no. 2, pp. 810–816, Jul. 2011. [16] T. W. Lewis, G. M. Spinks, G. G. Wallace, D. De Rossi, and M. Pachetti, “Development of an all polymer electromechanical actuator,” Polym. Prepr., vol. 38, no. 1–3, pp. 520–521, 1997.

[17] G. Alici and H. N. Nam, “Performance quantification of conducting polymer actuators for real applications: A microgripping system,” IEEE/ASME Trans. Mechatronics, vol. 12, no. 1, pp. 73–84, Feb. 2007. [18] R. J. Roark and W. C. Young, Formulas for Stress and Strain. New York: McGraw-Hill, 1989. [19] G. Alici, “An effective modelling approach to estimate nonlinear bending behaviour of cantilever type conducting polymer actuators,” Sens. Actuators B, Chem., vol. 141, no. 1, pp. 284–292, Aug. 2009. [20] G. M. Spinks, L. Liu, G. G. Wallace, and D. Zhou, “Strain response from polypyrrole actuators under load,” Adv. Funct. Mater., vol. 12, no. 6/7, pp. 437–440, Jun. 2002. [21] G. M. Spinks, B. Xi, D. Zhou, V. T. Truong, and G. G. Wallace, “Enhanced control and stability of polypyrrole electromechanical actuators,” Synth. Met., vol. 140, no. 2/3, pp. 273–280, Feb. 2004. [22] J. E. Shigley and L. D. Mitchell, Mechanical Engineering Design, 4th ed. New York: McGraw-Hill, 1984, pp. 132–134. [23] N. Jalili, P. X. Liu, G. Alici, and A. Ferreira, “Guest editorial: Introduction to the focused section on mechatronics for MEMS and NEMS,” IEEE/ASME Trans. Mechatronics, vol. 14, no. 4, pp. 397–404, Aug. 2009. [24] C. Tang and G. Alici, “Evaluation of length scale effects for mechanical behaviour of micro and nano cantilevers—Part 1: Experimental determination of length-scale factors,” J. Phys. D, Appl. Phys., vol. 44, no. 33, p. 335 501, Aug. 24, 2011. [25] C. Tang and G. Alici, “Evaluation of length scale effects for mechanical behaviour of micro and nano cantilevers—Part 2: Experimental verification of deflection models using atomic force microscopy,” J. Phys. D, Appl. Phys., vol. 44, no. 33, p. 335 502, Aug. 24, 2011.

Babita Gaihre received the M.S. degree in organic chemistry from Tribhuvan University, Kirtipur, Nepal, in 2001, and the Ph.D. degree from Chonbuk National University, Jeonju, Korea, in 2008. Her Ph.D. work involved synthesis and characterization of surface-modified nanoparticles and study of potential application of these surface-modified nanoparticles as a carrier system for targeted delivery of a drug. In 2009, she was appointed as Associate Research Fellow at the University of Wollongong, Wollongong, Australia, for a research project on the fabrication and performance study of microactuators made up of conducting polymers. After successful completion of this project, she is now working as a Research Fellow in the Australian Research Council Centre of Excellence for Electromaterials Science, University of Wollongong. Her current research involves patterning of conducting polymers for the fabrication of electrochromic and bionic devices.


Gursel Alici received the Ph.D. degree in robotics from the Department of Engineering Science, University of Oxford, Oxford, U.K., in 1994. He is currently a Professor at the University of Wollongong, Wollongong, Australia, where he is the Head of the School of Mechanical, Materials and Mechatronic Engineering. His current research interests include intelligent mechatronic systems involving mechanisms/serial–parallel robot manipulators; micro-/nanorobotic systems for medical applications; and modeling, analysis, characterization, and control of conducting polymers as macro-/micro-/nanosized actuators and sensors for robotic and bioinspired applications. He has produced over 190 refereed publications in his areas of research. Dr. Alici is a Technical Editor of the IEEE/ASME T RANSACTIONS ON M ECHATRONICS and a member of the National Panel on Mechatronics formed by the Institution of Engineers Australia. He was the recipient of the Outstanding Contribution to Teaching and Learning Award from the University of Wollongong in 2010.

Geoffrey M. Spinks received the Ph.D. degree from The University of Melbourne, Melbourne, Australia, in 1990 for his work on the fracture behavior of unsaturated polyesters. In 1990, he was appointed as Lecturer in the Department of Materials Engineering, University of Wollongong, Wollongong, Australia, where he is currently a Professor in the School of Mechanical, Materials and Mechatronic Engineering and a Chief Investigator with the Australian Research Council (ARC) Centre of Excellence for Electromaterials Science. He was a Visiting Researcher with BHP Research (1995) and Allied Signal (USA) (1999). He has published over 150 scientific journal articles and book chapters. His coauthored book Conductive Electroactive Polymers (CRC Press, 2009) is in its third edition. His main research interest is in the area of mechanical actuators based on organic materials, including conducting polymers, hydrogels, carbon nanotubes, and graphene. Prof. Spinks currently holds an ARC Professorial Fellowship.

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Julie M. Cairney received the Ph.D. degree in physical metallurgy from The University of New South Wales, Sydney, Australia, in 2002. She is currently an Associate Professor at The University of Sydney, Sydney. Her research interests focus on the relationship between microstructure and properties of materials, with a particular emphasis on the application and development of new characterization technologies in electron microscopy, focused ion beam technology, and atom probe tomography. Her current materials of interest include steel, nonferrous engineering alloys (such as Ni-based superalloys and Ti alloys), nanocrystalline metals, hard coatings (including nanocomposites and thermal barrier coatings), and thin films (including ferroelectrics).