Soft Polymer-Actuated Compliant Microgripper with Adaptive Vibration-Controlled Grasp and Release

Abstract

Microgrippers that incorporate soft actuators are appropriate for micromanipulation or microsurgery owing to their ability to grasp objects without causing damage. However, developing a microgripper with a large gripping range that can produce a large force with high speed remains challenging in soft actuation mechanisms. Herein, we introduce a compliant microgripper driven by a soft dielectric elastomer actuator (DEA) called a spiral flexure cone DEA (SFCDEA). The submillimeter-scale SFCDEA exhibited a controllable linear displacement over a high bandwidth and the capability of lifting 100.9 g, which was 670 times higher than its mass. Subsequently, we developed a compliant microgripper based on the SFCDEA using smart composite microstructure technology to fabricate three-dimensional gripper linkages. We demonstrated that the microgripper was able to grasp various millimeter-scale objects with different shapes, sizes, and weights without a complex feedback control owing to its compliance. We proved the versatility of our gripper in robotic manipulation by demonstrating adaptive grasping and releasing of small objects using vibrations owing to its high bandwidth.

Introduction

The ability to grasp and manipulate objects is a key aspect of robotic systems. The demand for miniaturized grippers has increased, revealing significant potential for micromanipulation,^1,2 microsurgery,³ and microassembly.⁴ Although microelectromechanical system structures have been successfully used for microgripper mechanisms,^1,5,6 the microgrippers are designed as two-dimensional (2D) parallel jaw grippers, which hinder their grasping performance.⁷ Grippers with three-dimensional (3D) structures, such as three-finger grippers, offer a more delicate and powerful gripping performance compared with 2D grippers. By leveraging smart composite microstructure (SCM) fabrication techniques, complex 3D structures with zero-backlash flexure joints can be created.^8,9 Recent studies have developed miniaturized grippers actuated by a shape memory alloy (SMA) spring using SCM technologies.^10,11 However, these grippers implement a 2D linkage mechanism to generate adaptive grasping, which results in limited grasping performance.

Another consideration of the microgripper is the actuator selection. To grasp various objects with different sizes and weights, the microgrippers require small-sized actuators with a large displacement and high output force. Moreover, a high actuator bandwidth is an important design aspect for the microgripper, as it allows the microgripper to release objects rapidly and actively using vibration.^12,13 It has been reported that the interactive forces between the gripper surface and micro/millimeter-scale objects are dominant, which hinder the release process.¹² The vibration-based releasing methods take advantage of inertial effects of object and end effector to overcome the interactive adhesion force.¹³

The most commonly encountered actuation principles for microgrippers are piezoelectric driven,^12,14,15 electrostatic driven,^13,16 electrothermally driven,^17,18 and SMA wires.^19,20 Piezoelectric devices are typically used to actuate microgrippers because of their compact size, high power density, high resolution, and high bandwidth. However, the small displacement of the piezoelectric actuator (<30 μm) limits the opening width of the grippers.^14,15 Electrothermal and SMA grippers are limited by their high operating temperature, high power consumption, low operational frequency, and hysteresis. Moreover, the rigidity of these actuators hinders gripping and manipulating objects without causing damage; thus, they require additional complex compliant mechanisms to solve this issue.^14–16,19

The use of soft actuators enables the grippers to provide safe interactions with the environment. In addition, compliance enables the grippers to grasp various unknown objects regardless of their size and shape without the need for precise feedback control.^21,22 Many microgrippers have been developed utilizing various soft actuators, such as SMA springs,^10,11 thermally activated polymers,^21,22 electromagnetic actuators,^23,24 ionic polymer metal composite (IPMC),²⁵ and light responsive actuators.²⁶ However, SMA springs and thermally activated polymers are limited by their high operating temperature and operating speed,^10,11,21,22 electromagnetic actuators and IPMCs are limited by their small gripping force,^23–25 and light-responsive actuators are limited by their operating speed and small gripping force.²⁶

Dielectric elastomer actuators (DEAs) are promising for microgrippers because they produce a large area strain (>100%), fast response time (in the order of milliseconds), high bandwidth (>100 Hz), high energy density (3.4 MJ⸱m⁻³), and high energy efficiency (90%).^27–29 DEAs consist of a DE membrane and compliant electrodes, which generate an area expansion owing to electrostatic pressure, known as the Maxwell stress. Although several impressive compliant grippers have been introduced using bending DEAs,^30–32 their design and size were not appropriate for micromanipulations, and their actuation speed was slow.

Among various DEA configurations, conical DEAs are characterized by linear actuation at a high output force and high bandwidth.²⁷ Conical DEAs use various mechanisms to preload DEA membranes in a single direction, such as compressive springs,³³ flexure springs,³⁴ passive membranes with a spacer,^35,36 or magnetic repulsion³⁷ to convert the area expansion of the DEA to a linear actuation. However, these designs are difficult to reduce in size to the submillimeter scale because they have instability problems owing to buckling of the spring³³ and complex structures.³⁵ In addition, the viscoelasticity of passive membranes limits their actuation bandwidth.^35–37

This study aimed to develop a compliant microgripper to adaptively grasp various millimeter-scale objects. The microgripper is fabricated by integrating the spiral flexure cone DEA (SFCDEA) and gripper linkages based on the SCM technology (Fig. 1a). The fabricated microgripper enables the grasping of submillimeter-scaled objects (Fig. 1b). The microgripper was driven by a SFCDEA, which is one of a linear DEA (Fig. 1c). The optimal design of the SFCDEA was derived using an analytical dynamic model. Then, we designed the gripper linkage to convert the actuation of the SFCDEA into a gripping motion (Fig. 1d). The size of the fabricated microgripper was 10 × 10 × 18 mm³, and the mass was 0.5 g.

FIG. 1.

Design of the soft polymer-actuated compliant microgripper. (a) Schematic and image of a compliant microgripper. (b) Photograph of the compliant microgripper with a microsurgical robot. (c) Schematic of the operating principle of the SFCDEA, consisting of a spiral flexure spring, multilayered DEA, and connector. The SFCDEA produces a controllable actuation displacement by an applied input voltage V. (d) Image of the actuation of the microgripper. DEA, dielectric elastomer actuator; SFCDEA, spiral flexure cone DEA.

Table 1 presents comparison between our microgripper and other compliant two-finger microgrippers. Although other compliant microgrippers have been introduced, they do not satisfy all requirements for adaptive manipulation of millimeter-scaled objects, such as gripping size, force, and bandwidth. For example, SMA spring-based and thermally activated polymer-based grippers can produce high gripping force (>200 mN) and gripping size (>1.5 mm), but actuation speed is slow (>5 s). Electromagnetic and IPMC-based grippers are limited by their gripping force (<0.2 mN), and piezoelectric and electrostatic-based grippers are limited by their small gripping range (<0.1 mm). Our gripper can produce reasonable gripping force (20.5 mN) and gripping range (0–3.7 mm) for manipulation of millimeter-scaled objects with a high bandwidth (17 Hz). Owing to its benefits, our gripper is the only microgripper that can adaptively grasp densely piled objects and actively release them while operating at high frequencies without using complex control algorithms.

Table 1.

Comparison Between the Presented and Other Compliant Two-Finger Microgrippers, Capable of Grasping Millimeter-Sized Objects

Actuation principle	Size (mm)	Max. gripping size (mm) (normalized by gripper length)	Max. gripping force (mN)	Bandwidth (Hz)	Active release	Demonstration of pin-picking
DEA (This work)	18 × 10 × 10	3.7 (0.20)	20.5	17	Yes	Yes
SMA spring¹⁰	26 × 14 × -^a	12^a (0.46)	250^a	<0.2^a	No	No
Thermally activated polymer²¹	12 × 9 × 4	1.8 (0.15)	2000	<0.2^a	No	No
Electromagnetic²⁴	6 × 8 × 8^a	3.9 (0.6)	0.05^a	30^a	No	No
IPMC²⁵	22 × 12 × -^a	1.5 (0.07)	0.17	0.07^a	No	No
Piezoelectric¹²	62 × 36 × -^a	0.06 (0.97 × 10^–3)	30	100^a	Yes	No
Electrothermal+Electrostatic¹³	0.4 × 0.2 × -^a	2.26 × 10^–3 (5.67 × 10^–3)	12.9 × 10^–3	55 × 10^3,a	Yes	No

Approximation based on published data.

DEA, dielectric elastomer actuator; SMA, shape memory alloy.

Materials and Methods

Design concept and working principle of the SFCDEA

We designed an SFCDEA that consists of a multilayered DEA, spiral flexure spring, and connector. The basic components and operating principle of the proposed actuator are illustrated in Figure 1c. When an electric field is applied to the DEA, the planar area expands owing to the Maxwell stress. A spiral flexure spring was used to preload the DEA and convert the area expansion of the DEA to a linear actuation. The spiral flexure spring restricted the actuation of the DEA to the vertical direction, as it has a high radial/axial stiffness ratio.³⁸ A spiral flexure spring is composed of multiple spiral slots that can function as a linear spring.³⁹ Considering their thin (thickness = 0.1 mm), simple, and lightweight structure, spiral flexure springs can be fabricated with a small size. In addition, a spiral flexure spring can produce a large stroke without buckling.

The restoring force of the prestretched DE membrane was balanced with that of the spiral flexure spring under a zero-input voltage. When voltage was applied, the area expanded in the DE membrane owing to electrostatic forces. The area expansion of the DEA caused the connector to move downward until it reached another force balance. The proposed SFCDEA can produce large displacements and output forces because the spiral flexure spring can store and release the elastic energy.

SFCDEA fabrication procedure

The fabrication process of the SFCDEA is depicted in Supplementary Figure S1. Among various DE materials, 3M VHB 4905 film was chosen as the DE membrane because of its low Young's modulus of 0.4 MPa, high DE constant of 4.53, and high dielectric strength of 25 kV/m.³⁶ First, the DE membrane was biaxially prestretched and bonded to a frame with an inner diameter of 6 mm. The prestretch of the DE membrane can effectively improve the actuation performance of the acrylic-based DEA, since it simultaneously reduces the thickness of the DE membrane and enhances the electromechanical properties, such as electrical breakdown strength and pull-in stability.^40,41 However, the prestretch also causes a decrease in the DE constant.⁴² It has been reported that the circular acrylic-based DEA with a prestretch ratio of four produces the largest radial strain at 3500 V.⁴¹ Following the previous research, we choose the prestretch ratio of 4.

Both sides of the membrane were coated with the multiwalled carbon nanotubes (MWCNTs) as compliant electrodes. For the electrical connection, aluminum tape with a thickness of 30 μm was attached. The DE membranes with CNT coating became multilayered structures by stacking them thrice. The fabrication process was finalized by combining the multilayered DE membrane and prepared spiral flexure spring together from the center by two connectors at a space of 3 mm. A circular magnet was chosen as the connector to allow for easy center alignment and enhanced bonding force. The spiral flexure spring was made of a stainless steel (SUS304) sheet with a thickness of 0.1 mm and fabricated using chemical etching. The mass of the fabricated SFCDEA was 0.13 g (Fig. 2).

FIG. 2.

Overview of the spiral flexure spring as a biasing element in SFCDEA. (a) Predicted displacement response of the SFCDEA under different spring constants of the spiral flexure spring based on the analytic dynamic model. (b) Schematic design parameters of the spiral flexure spring with spiral gap g_spiral and slot width w_spiral. (c) Finite element analysis of spiral flexure spring. (d) Results of the finite element analysis. (e) Optical microscope images of the fabricated spiral flexure spring. (f) Restoring force of the spiral flexure spring as a function of deformation. FEM, finite element method.

Modeling and design optimization

We derived a dynamic model of the SFCDEA to optimize the dynamic performance. The dynamic model of the SFCDEA can be derived by combining the force balance equation, Kelvin–Voigt–Maxwell model, and Maxwell stress.^35–37 Details of the dynamic model of the SFCDEA are summarized in Supplementary Text S1 and Supplementary Figure S2. Then, we optimized the spring constant of the spiral flexure spring to obtain the maximum displacement of the SFCDEA using the derived dynamic model. Figure 2a shows the predicted actuation displacement of the SFCDEA in accordance with the spring constant of the spiral flexure spring using a model-based simulation at an actuation frequency of 0.1 Hz. Simulation results showed that the SFCDEA produced a maximum actuation displacement when the spring constant was 0.18 N/mm. The optimized design parameters of the SFCDEA are summarized in Supplementary Table S1.

Two design parameters of the spiral flexure spring, spiral gap g_spiral and slot width w_spiral, must be controlled to fabricate a spiral flexure spring with a desired spring constant, as shown in Figure 2b. We used the finite element method (FEM) simulation (COMSOL software) to estimate the spring constant of the spiral flexure spring, as shown in Figure 2c. The estimated spring constant of the spiral flexure was calculated as the ratio between the applied force and deformation occurring at the center. Figure 2d shows the estimated spring constants in accordance with slot widths when the spiral gap is fixed at 6 mm. The spring constant and slot width have a negative linear relationship. The FEM analysis results showed that the spiral flexure spring had a spring constant of 0.18 N/mm when the spiral gap was 6 mm, and slot width was ∼0.25 mm. Figure 2e shows the fabricated spiral flexure spring with the specified design parameters.

The microscopic image shows that the spiral flexure spring can be fabricated with a high precision. The hysteresis curve of the spring was measured to evaluate the dynamic performance of the fabricated spiral flexure spring, as shown in Figure 2f. The fabricated spiral flexure spring has a desired spring constant with high linearity and small hysteresis error of 1.7% at 1 Hz. The repeatability test results showed that the spiral flexure spring can produce a constant restoring force over 1000 cycles at high frequencies (1, 20, 200 Hz), as shown in Supplementary Figure S3. In addition, the spiral flexure spring has a high radial/axial stiffness ratio of 46, as shown in Supplementary Figure S4. The high radial/axial stiffness can prevent the off-axial actuation of the SFCDEA (Fig. 3).

FIG. 3.

Actuation displacement test results of SFCDEA. (a) Optical images showing the actuation of SFCDEA under applied input voltage. (b) Actuation displacement of SFCDEA as a function of frequency of the input voltage at a fixed voltage of 4.5 kV. (c) Prediction and experimental actuation displacement of SFCDEA as a function of input voltage at 1 Hz. (d) Actuation displacement of SFCDEA and input voltage. (e) Hysteresis curve of actuation displacement at 1 Hz. (f) Repeatability test results of SFCDEA over 2000 cycles.

Results

Performance of SFCDEA

The actuation performance of the SFCDEA was evaluated by measuring the actuation displacement and blocked force, as shown in Supplementary Figure S5. The actuation displacement of the SFCDEA increased with the applied input voltage at various frequencies (Fig. 3a and Supplementary Movie S1). Figure 3b shows the measured actuation displacement as a function of the actuation frequency at a sinusoidal input voltage of 4.5 kV. We observed that the peak at 0.1 Hz presents the largest actuation displacement of 0.36 mm (12% of the actuation strain). In the frequency range of 0.1–100 Hz, the actuator can produce an actuation displacement larger than 0.13 mm. The calculated bandwidth is ∼10 Hz, as shown in Supplementary Figure S6.

Figure 3c illustrates the measured and model-based predicted actuation displacement as a function of the input voltage, where frequency of the input voltage is fixed at 1 Hz. The results indicate that the SFCDEA can produce a controllable actuation displacement by tuning the input voltage. In addition, the graph shows that the model-based prediction agrees well with experimental results. The SFCDEA has a fast response time of 32.5 ms at 1 Hz, as shown in Figure 3d. In addition, the actuator had a hysteresis error of 31.5% at an operating frequency of 1 Hz, as presented in Figure 3e, primarily due to the viscoelasticity of the DE membrane. The repeatability test results showed that the actuator could produce a fairly constant performance over 2000 cycles, as shown in Figure 3f. During the 2000 cycles, we have not observed significant off-axial actuation of the SFCDEA (Fig. 4).

FIG. 4.

Output force characteristics of the SFCDEA. (a) Blocked force as a function of frequency of the input voltage at a fixed voltage of 4.5 kV. (b) Output displacement characteristics of the SFCDEA under various external loads. (c) Single SFCDEA producing an actuation strain under external load of 100.9 g.

The blocked force of the SFCDEA was measured at various frequencies with a sinusoidal input signal of 4.5 kV, as shown in Figure 4a. Two peaks were observed at frequencies of ∼62 and 124 Hz. The simplified Maxwell stress equation can be used to explain the appearance of these two peak values, in which the stress is proportional to the square of the applied voltage.³⁶ The maximum measured blocked force at 124 Hz was 1.2 N. The actuator was able to produce a blocked force of 0.23 N at a low operating frequency of 0.1 Hz, as shown in the inset of Figure 4a. The blocked force measured at the peak frequency of 124 Hz exhibited a consistent output, as shown in Supplementary Figure S7.

In addition, we measured the output displacement of the SFCDEA under various external loads. A sinusoidal input voltage signal of 4.5 kV was applied to the actuator at an operating frequency of 1 Hz. The measured load–displacement curves are shown in Figure 4b. The SFCDEA with a mass of 0.13 g produced a displacement of 0.14 mm (4.67% linear actuation strain) under a preload of 100.9 g, which was 670 times higher than its mass, as presented in Figure 4c and Supplementary Movie S1 (Fig. 5).

FIG. 5.

Design and operating principle of the SFCDEA based compliant microgripper. (a) Schematic of the basic structure of the microgripper, consisting of an actuator frame, a spacer, and a gripper linkage. (b) Schematic of the operating principle of the microgripper. The gripper linkages convert and amplify the actuation of the SFCDEA to gripping motion. (c) Kinematic design of the microgripper with controllable opening width. (d) Estimated maximum opening width (s_open) with various design parameters. (e) Estimated force transmission ratio (G_F) with various design parameters.

Design concept of the microgripper

We used the proposed SFCDEA and designed a compliant microgripper consisting of a gripper linkage, spacer, and actuator frame, as shown in Figure 5a. Figure 5b shows the operating principle of the gripper. The gripper linkage converts the actuation displacement of the SFCDEA into a gripping motion. The gripper is designed to be in a closed state initially. The microgripper is opened by actuation of the SFCDEA, which is connected to the inner linkage. The gripper uses the passive restoring force of the SFCDEA to grasp objects. Therefore, this mechanism consumes minimal power.

The kinematic design of the microgripper is shown in Figure 5c and Supplementary Figure S8, where a_g is the inner gap, b_g is the height gap, s_closed is the minimum opening width, and s_open is the maximum opening width of the gripper. The relationship between the maximum opening width of the gripper and actuation displacement of the SFCDEA can be derived using the kinematic equations, as presented in Supplementary Text S2. Using the kinematic solution, the microgripper can be designed to enable a specific target opening width, as shown in Figure 5d. The results show that the gripper linkage design converts and amplifies an actuation displacement of 0.36 mm of the SFCDEA to an opening width larger than 4 mm.

In addition, the gripping force capability of the microgripper is a key factor. The force transmission ratio (G_F) is defined as the ratio between the restoring force of the SFCDEA and the grasping force, as described in Supplementary Text S2. Figure 5e shows the calculated force transmission ratios for various design parameters. We observed a trade-off between the maximum opening width and force transmission ratio of the microgripper, as shown in Supplementary Figure S9. The design parameters of the microgripper were selected to satisfy the target opening width (s_open) with the maximum force transmission ratio (G_F), as presented in Supplementary Table S2 (Fig. 6).

FIG. 6.

Fabrication process of microgripper. (a) Manufacturing process of gripper linkages and top frame based on SCM technology. (b) Components and manufacturing process of the SFCDEA frame using a locking mechanism. (c) Schematic of the structure of the microgripper with components labeled. (d) Photograph of the fabricated compliant microgripper. SCM, smart composite microstructure.

Fabrication of the microgripper

The fabrication process of the microgripper is illustrated in Figure 6. The microgripper was fabricated using SCM technology, as shown in Figure 6a. Five layers, including a polyimide film of 12.5 μm, two fiberglass layers of 100 μm, and two 50 μm sheets of acrylic adhesive (3M F9460) were individually laser micromachined and prepared. The five layers were precisely aligned using dowel pins and laminated into a composite structure. The resulting composite structure was cut along the cutting line to release it from the surrounding material. Once released, the structures were folded and glued to maintain their configuration. Subsequently, the inner linkage, outer linkage, and top frame were created. Figure 6b shows the SFCDEA frame fabrication process, which was simplified using a locking mechanism with a locker.⁴³ The space between the spiral flexure spring and DEA can be easily maintained by folding and rotating the lockers, as shown in Figure 6b. The fabrication of the microgripper was finalized by combining all components as shown in Figure 6c.

The size of the fabricated microgripper was 10 × 10 × 18 mm³, and the mass was 0.5 g, as shown in Figure 6d and Figure 7.

FIG. 7.

Experimental evaluation of the microgripper. (a) Optical images of the gripping motion of two-finger gripper at various input voltages. (b) Opening width of two-finger gripper in accordance with various DC input voltages. (c) Optical images of the gripping motion of three-finger gripper at various input voltages. (d) Opening width of three-finger gripper in accordance with various DC input voltages. (e) Gripping force of two-finger gripper in accordance with various opening widths. (f) Measured gripping force of two-finger gripper at various input voltages.

Gripping actuation test

The opening width of the gripper could be controlled by modulating the input voltage applied to the SFCDEA, as illustrated in Figure 7a. The actuation of the two-finger microgripper is shown in Supplementary Movie S2. The gripper was able to actuate at a high speed considering the high bandwidth of the SFCDEA. As the polyimide film-based flexure hinges are essentially frictionless and contribute minimal damping and stiffness, the SCM-based hinge can operate at high frequency with minimal energy dissipation.^8–11 To check the mechanical stability of the polyimide flexure hinge under high frequency operation, we conducted the repeatability test, as shown in Supplementary Figure S10. The results showed that the behavior of the polyimide flexure hinge does not change across 10,000-folding cycles at a frequency of 100 Hz.

Figure 7b shows the opening width as a function of the input DC voltage. The experimental results show a reasonable consistency with the predicted results based on the kinematic model with a small variation. The difference between the experimental and model-based results was calculated as 0.3 mm at the maximum input voltage. This is likely due to fabrication errors of the gripper linkage and SFCDEA owing to its small size. The calculated bandwidth of the two-finger microgripper is ∼17 Hz, as shown in Supplementary Figure S11. Compared with the bandwidth of the SFCDEA (10 Hz), the bandwidth shift in the gripper is likely owing to the small gaps in the flexure hinge regions. In addition, a three-finger microgripper was fabricated, which exhibited a delicate and powerful gripping performance, as illustrated in Figure 7c. The actuation of the fabricated three-finger microgripper is shown in Supplementary Movie S2. The opening width of the three-finger microgripper was obtained considering the input voltage change and compared with the predicted results based on the kinematic model, as shown in Figure 7d.

Gripping force tests

The gripping force of the two-finger microgripper was measured with opening widths in the range of 0–3.7 mm. The experimental setup for measuring the gripping force of the microgripper is shown in Supplementary Figure S12. Results showed that the gripper produces a maximum gripping force of 20.5 mN, as shown in Figure 7e. The gripping force of the two-finger gripper can be also estimated using the derived kinematic model of the gripper, as shown in Supplementary Text S2. The measured gripping force of the gripper was consistent with the estimated gripping force.

Furthermore, we measured the gripping force according to the input voltage, with an opening width of 3.7 mm. The results indicate that the gripping force can be controlled by modulating the input voltage, as shown in Figure 7f. To further evaluate the gripping performance, we measured the holding force of the gripper with objects of the same size and different surfaces, as shown in Supplementary Figure S13. Under the same gripping force, results showed that the holding force is highly dependent on the friction and adhesion between the gripper and the object. The gripper was able to produce the highest holding force of 78 mN (Fig. 8).

FIG. 8.

Performance of the SFCDEA based microgripper. (a) Pick-and-place demonstration of the microgripper. (b) Compliant microgripper grips and lifts various objects, including a weight (2 g), metal ball (0.5 g), M3 bolt and nut (1.1 g), capacitor chip (0.4 g), and polymer tube (0.1 g). (c) Microgrippers were able to operate with a microsurgical robot. (d) Microgripper adaptively grasping densely-piled object without damage using vibrations. (e) Microgripper actively releasing object using vibrations.

Manipulation of various objects

The performance of the microgripper was evaluated by demonstrating the pick-and-place behavior of various small objects, as shown in Figure 8a. The gripper successfully demonstrated the grasping and manipulation of a rigid weight (diameter = 5 mm, mass = 2 g), metal ball (diameter = 5 mm, mass = 0.5 g), M3 bolt and nut (diameter = 5.4 mm, mass = 1.1 g), capacitor (diameter = 5.1 mm, mass = 0.4 g), and thin polymer tube (diameter = 4 mm, mass = 0.1 g), as shown in Figure 8b and Supplementary Movie S3. The microgripper enables the successful grasping of various objects without complex sensory feedback owing to the softness of the SFCDEA.

In addition, we further fabricated compliant two-finger microgripper with a smaller size (size = 6 × 6 × 12 mm³, mass = 0.16 g) using miniaturized SFCDEA (diameter = 6 mm). We evaluated the performance of the microgripper by demonstrating the pick-and-place of a knurled nut (diameter = 3.5 mm, mass = 0.1 g), as shown in Figure 8c and Supplementary Movie S4. During the demonstration, the microgripper was attached to surgical instruments with three degrees of freedom (3-DOF). This demonstration reveals the feasibility of implementation with various microrobots.

Adaptive grasping and active release using vibration

The microgripper can operate at a high frequency because of the high bandwidth of the SFCDEA. The fabricated compliant microgripper can adaptively grasp densely-piled small-sized objects using vibrations without complicated algorithms, as shown in Figure 8d and Supplementary Movie S5. The gripper can be effectively inserted into the gaps between the target and surrounding objects by generating a vibration at the end of the gripping tip.

In addition, the adhesion force between the polymer pads and objects hinders their release. The vibration of the compliant microgripper allows the active detachment of the objects, as shown in Figure 8e and Supplementary Movie S6. The microgripper used the vibration and exhibited a rapid active release of various objects. This is due to the inertial effect of the object and the end effector. The release acceleration is the key of the active detachment, and it has been reported that the minimum acceleration of the end effector to release the object decreases with the larger object size.⁴⁴ In this research, 10 Hz vibration was sufficient to release a polymer sphere with a diameter of 4 mm.

Conclusion

In this study, we developed a compliant microgripper driven by an SFCDEA. SFCDEAs provide compliant linear actuation with a high output force, high bandwidth, rapid response, and lightweight design. We used the characteristics of the SFCDEA and fabricated a compliant microgripper. The microgripper can be fabricated on a millimeter scale using SCM technology. The fabricated gripper design consumes minimal power while grasping objects, as it uses the passive restoring force of SFCDEA to grasp objects. We demonstrated the pick-and-place of various millimeter-scale objects, to prove the feasibility of our microgripper for micromanipulation. The proposed microgripper could grasp various objects regardless of their size and shape without using a feedback control, owing to its compliance. Our microgripper can operate at a high frequency, producing both grasping motion and vibration. The fabricated microgripper can use this high bandwidth characteristic and adaptively grasp densely piled objects and use vibration to actively release objects.

The primary challenge of our compliant microgripper is the requirement for a high input voltage. Future studies will attempt to decrease the operating voltage of the SFCDEA. To this end, the thickness of the DE membrane must be reduced. In addition, further miniaturization and improvements to increase the gripping force will be explored in future studies. The gripping force of the gripper depended on the restoring force of the SFCDEA. Therefore, the gripping force could be increased by increasing the number of DE layers.

Footnotes

Acknowledgment

The IEEE conference article (DOI: 10.1109/ICRA48891.2023.10160797) contributes to develop this work.

Author Disclosure Statement

No competing financial interests exist.

Funding Information

This work was supported by the R&D Program (No. 2022R1A2B5B02002074) of the National Research Foundation (NRF) of Korea and the Field-Oriented Technology Development Project for Customs Administration (2022M3IA1095075) of the National Research Foundation (NRF) of Korea funded by the Ministry of Science & ICT and Korea Customs Service.

Supplementary Material

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