The output displacement of the traditional symmetrical microgripper is large, but its micro-components or parts are easily damaged due to the uneven force exerted on the left and right jaws of the gripper. The output force of the traditional asymmetric microgripper is stable. However, its output displacement is small, typically half the output displacement of the symmetric microgripper. To solve these problems, in this study, we designed a large-displacement asymmetric microgripper. First, we calculated the relationship between the theoretical input and output variables based on their geometric relationship. Then, we analyzed the performance of the microgripper using finite element software. Lastly, we used a piezoelectric actuator as the input driver of the microgripper. The errors associated with the theoretical and simulated output displacements were 7.05% and 9.24%, respectively. At 150 V of driving voltage, the maximum output displacement was 224 µm, and the actual magnification was 11.2 times. Microparts can be gripped in parallel and stably, which confirms the validity of the design.

  • A large-displacement asymmetric microgripper was designed.

  • The microgripper realizes a large amplification factor.

  • Microparts were grasped through the microgripper.

In recent years, with the rapid development of high technologies such as micro-electro-mechanical systems (MEMS), research on micro-operation has made great progress1 and is now widely used in the life sciences, automobile industry, information and communication fields, precision processing, and aerospace.2–6 The microgripper is a terminal manipulator that makes direct contact with a microobject, and its performance can determine the success of the micro-assembly task.

Currently, the key issues in microgripper research are the selection of the driving mode, increasing the displacement magnification, and improving clamping accuracy. Today, common driving modes of the microgripper include the piezoelectric drive,7 electrostatic drive,8 thermal drive,9 shape-memory alloy drive,10 and pneumatic drive.11 Compared with other driving modes, the piezoelectric drive has the advantages of fast response, high sensitivity, and large output force.12,13 Enlarging the gripping displacement is mainly accomplished by a micro-displacement amplifier. Commonly used micro-displacement amplifiers include the lever amplifier, bridge amplifier, and rhombic amplifier.14 The structure of the lever amplifier is simple and easy to produce, but its output displacement is small and the structure is not compact enough. The output displacements of the bridge and rhombic amplifiers are large and the piezoelectric ceramics are placed inside the mechanism, which makes the structures compact. Single lever, bridge, and rhombic amplifiers are categorized as single-stage amplifiers. Amplifiers composed of multiple single-stage amplifiers are called multi-stage amplifiers. A multi-stage amplifier can further expand the gripping stroke of the microgripper and increase the degree of magnification.15 Improvement in gripping accuracy is mainly achieved by its parallel and asymmetric gripping capability.

Currently, most research is focused on symmetrical microgrippers, with less attention paid to asymmetrical microgrippers. However, symmetrical microgrippers have some limitations in the scope of their application due to the uneven stress placed on their left and right jaws. Cui et al.16 designed a symmetrical microgripper based on the lever amplification principle to achieve parallel gripping; however, for an asymmetrical structure, the actual maximum gripping forces of the left and right jaws are 8.02 mN and 9.24 mN, respectively. Since the maximum gripping forces are different, it is easy to damage the thin-walled fragile grippers. Koo et al.17 designed an asymmetrical microgripper based on the lever amplification principle to realize stable gripping of the jaws. However, their microgripper could not grip in parallel. Xing et al.18 designed an asymmetrical microgripper based on the principle of lever amplification, which realized parallel gripping of the jaws, but the magnification was small. Zhao et al.19 designed a two-stage asymmetrical microgripper based on the principle of lever amplification that realized both the parallel gripping of jaws and high magnification.

In summary, the selection of driving mode performance, parallel gripping, stability, and high magnification characteristics have been considered comprehensively. Next, it is necessary to design an asymmetrical piezo-driven microgripper with a large jaw displacement.

The rest of the paper is organized as follows: In Section 2, we introduce the structural design of the microgripper. We analyze the microgripper characteristics by finite element analysis (FEA) in Section 3. In Section 4, we discuss the results of an experiment we conducted to verify the performance of the microgripper developed by our research institute. Section 5 concludes this paper.

Fig. 1 shows a diagram of the main view of a microgripper. The size of the mechanism is 46.50 mm×45.12 mm×5 mm. The microgripper mainly consists of a stack piezoelectric ceramic actuator (SPCA), a straight circular flexure hinge, a straight flexure hinge, a grasping jaw, a lever amplifier, a rhombic amplifier, fixing holes, and a preload bolt. To obtain a large amplification ratio, the microgripper is equipped with a two-stage amplifier.

Fig. 1.

Main view of the microgripper.

Fig. 1.

Main view of the microgripper.

Close modal

The right half of the microgripper served as the object of our analysis. Fig. 2 shows a pseudo-rigid-body model, in which the hinges have been replaced by torsional springs. Fin and xin represent the input force and displacement from the piezoelectric actuator, respectively, and Fout and xout indicate the output force and displacement, respectively. Table 1 shows the main dimensions of the mechanism.

Fig. 2.

Pseudo-rigid-body model of microgripper.

Fig. 2.

Pseudo-rigid-body model of microgripper.

Close modal
Table 1.

Main dimensions of mechanism (mm).

a b c d
9.14  1.22  4.02  13.24 
a b c d
9.14  1.22  4.02  13.24 

The connecting rod mechanism ABC can be regarded as a crank-slider mechanism, as shown in Fig. 3. During movement of the crank-slider mechanism, the deformations of the Y- and X-axes are denoted as ΔBy and ΔAx, respectively. The following relationships hold:V

ΔByΔAx=By/dtAx/dt=vByvAx
(1)
vAx=wABlOA
(2)
vBy=vBcθ=wABlOBcosθ
(3)

where v is the velocity and w is the angular velocity.

Fig. 3.

Crank-slider mechanism.

Fig. 3.

Crank-slider mechanism.

Close modal

Because the lengths of connecting rods AB and BC are equal, we substitute Eqs. (2) and (3) for Eq. (1) to obtain the displacement amplification ratio of the crank-slider mechanism:

ΔByΔAx=lOBcosθlOA=a2b
(4)

The theoretical amplification ratio of the microgripper is as follows:(5)

Ramp=xoutxinad2bc
(5)

By applying a certain displacement at the input end, the rotational deformation of all the torsion springs can be obtained as follows:

θA=(π2θB)=θC=(π2θD)=xin2b
(6)
θE=θF=θG=θH=axin2bc
(7)

According to Howell20 and Paros and Weisbord,21 the rotational stiffness of a circular flexure hinge and flexure beam can be expressed, respectively, as follows:

{Krci=2EBti5/29πri1/2i=(A,,D)Krbi=2EBti5/29πri1/2i=(E,,H)
(8)

According to the principle of virtual work, the following relations hold:

Finxin+Foutxout=12(KrAφA2+KrBφB2+KrCφC2+KrDφD2+4KrEφE2)
(9)

Fig. 4 shows a diagram of the structure of the microgripper. The performance of the microgripper is mainly determined by the size of the hinges. By optimizing the size of the microgripper hinges, we can obtain the optimum geometric parameters of the microgripper, and achieve optimal performance. Table 2 lists the optimized dimensions of the microgripper parameters.

Fig. 4.

Diagram of microgripper structure.

Fig. 4.

Diagram of microgripper structure.

Close modal
Table 2.

Optimized dimensions of structural parameters.

Symbol Parameter Description Value
a   l × w   Dimension of straight flexure hinge  1.1 mm × 0.3 mm 
b   l × w   Dimension of straight flexure hinge  1.1 mm × 0.3 mm 
c   l × w   Dimension of rhomboid arm  8.1 mm × 1.76 mm 
d   l × w   Long arm lever  13.24 mm 
e   l × w   Short arm lever  4.02 mm 
f   l × w   Dimension of radius of straight circular flexure hinge  2.75 mm 
Symbol Parameter Description Value
a   l × w   Dimension of straight flexure hinge  1.1 mm × 0.3 mm 
b   l × w   Dimension of straight flexure hinge  1.1 mm × 0.3 mm 
c   l × w   Dimension of rhomboid arm  8.1 mm × 1.76 mm 
d   l × w   Long arm lever  13.24 mm 
e   l × w   Short arm lever  4.02 mm 
f   l × w   Dimension of radius of straight circular flexure hinge  2.75 mm 

The design parameters of the microgripper are as follows:

  1. The microgripper material and micro-parts include 7075 aluminum alloy, a modulus of elasticity E = 71 GPa, a Poisson's ratio v = 0.33, a yield strength σ = 455 MPa, and a density ρ = 2810 kg/m3.

  2. The thickness of the clamp is 5 mm. Fig. 5(a) and (b) show displacement and stress nephograms of the microgripper when the input displacement is applied at the input end of the microgripper and the micropart is not gripped. Fig. 5(a) shows that with an output displacement of the grasping jaw of 229.48 µm at an input displacement of 20 µm, and a simulation magnification of 11.474 times, parallel gripping can be realized. At the maximum output displacement, the maximum pressure on the weakest part of the microgripper is 258.13 MPa, which is less than the yield strength of the material. Therefore, the product can be used safely.

Fig. 5.

Finite element analysis diagram.

Fig. 5.

Finite element analysis diagram.

Close modal

Fig. 6 shows the relationship between the input displacement of a SPCA and the gripping force of the jaw when the microgripper has microaxes of 300 µm, 200 µm, and 100 µm, respectively. From the figure, we can see that there is a linear relationship between the input displacement of the SPCA, and the gripping force of the grasping jaw, which reveals that the microgripper has stable performance and achieves a smooth transition from the closure of the grasping jaw to the gripping of the micro-parts. The theoretical and simulated values of the output force are 5.45 N and 5.09 N, respectively, for an input displacement of 20 µm. The error of the theoretical value is 6.61% with respect to the simulated value, which is mainly attributed to the restraint of the two-stage amplifier on the single-stage amplifier.

Fig. 6.

Input displacement of SPCA versus the gripping force of grasping jaws.

Fig. 6.

Input displacement of SPCA versus the gripping force of grasping jaws.

Close modal

Fig. 7 shows a photograph of the physical model of the microgripper, in which the material used for the micro-clamp is 7075-T6 (SN) of aluminum alloy. The microgripper was processed using a LSWEDM (low speed wire electrical discharge machine). After processing, the microgripper was drilled and polished. The experimental equipment included an HPV-1 C 0300 A0300 piezoelectric ceramic driving power supply, a micro-sodium positioning worktable, a PZT (Suzhou Mat, Inc. SZBS150/5×5/20, open-loop travel 20 µm) driving micro-positioning stage, a high resolution capacitive displacement sensor (BJZD's MA-0.5), a data acquisition card (NI's PCI-6221), a 24-V DC-regulated power supply WP100-D-G, and a host computer and display. To eliminate external interference as much as possible, all the devices were installed on a high-performance vibration isolation platform.

Fig. 7.

Physical model of microgripper.

Fig. 7.

Physical model of microgripper.

Close modal

4.2.1. Analysis of microgripper performance

To validate the performance of the microgripper, we conducted a series of experiments. Fig. 8 shows the experimental device. Fig. 8(a) shows a theoretical working diagram of this device, and Fig. 8(b) shows a photographic diagram of the actual components of the experimental device.

Fig. 8.

Working diagram of the experimental device.

Fig. 8.

Working diagram of the experimental device.

Close modal

Fig. 9 shows the relationship between the driving voltage and the elongation of the SPCA. The results confirm the non-linear characteristics of the piezoelectric materials. Under open-loop control, the maximum elongation of the SPCA was 22.08 µm at a driving voltage of 150 V.

Fig. 9.

Input displacement versus drive voltage.

Fig. 9.

Input displacement versus drive voltage.

Close modal

Fig. 10 shows the variation in the tip displacements of the microgripper jaw with the input displacements obtained using different approaches. It can be seen from the graph that the theoretical magnification of the microgripper is 12.34 times and that of the simulation analysis is 11.47 times. This error is mainly due to the use of the pseudo-rigid-body method, which regards a flexible hinge as a torsional spring and a connecting rod as a rigid body, which means no deformation occurs during the movement of the mechanism. However, during actual movement, the rigid body will undergo minor deformation. The experimental magnification factor was 11.2 times. This error is mainly due to an accuracy error in the process of microgripper machining and measurement error in the process of experimental testing, as well as the vibration and noise of the equipment and other factors.

Fig. 10.

Tip displacement versus the input displacement of the SPCA.

Fig. 10.

Tip displacement versus the input displacement of the SPCA.

Close modal

To verify the parallel gripping characteristics and stability of the microgripper, we performed several grasping experiments. Fig. 11 shows a photograph of the microgripper grasping a metal wire and a plastic plate, respectively, which show that the microgripper exhibits parallel gripping characteristics and self-adaptability for grasping an object. Fig. 12 shows the microgripper approaching and grasping a 250-μm metal ball.

Fig. 11.

Grasping manipulation of microobjects of different shapes and sizes.

Fig. 11.

Grasping manipulation of microobjects of different shapes and sizes.

Close modal
Fig. 12.

(a–c) Approaching and (d) grasping a microobject (250 µm).

Fig. 12.

(a–c) Approaching and (d) grasping a microobject (250 µm).

Close modal

4.2.2. Performance comparison

As reported in references,18–21Table 3 lists the magnification of each microgripper, from which we can see that the magnification of the microgripper designed in this study can reach 11.2 times. Compared with similar types of microgrippers, the magnification has been greatly improved.

Table 3.

Comparison of parameters with similar microgrippers.

Microgripper Magnification
Reference17   4.35 
Reference18   4.16 
Reference 19   6.1 
Design of this paper  11.2 
Microgripper Magnification
Reference17   4.35 
Reference18   4.16 
Reference 19   6.1 
Design of this paper  11.2 

To address the shortcomings of traditional symmetrical and asymmetrical microgrippers, in this study, we designed a large-displacement asymmetrical microgripper that has both the large output displacement of the traditional symmetrical microgripper and the stable gripping performance of the traditional asymmetrical microgripper. First, we calculated the relationship between the theoretical input and output variables using the pseudo-rigid-body method. Then, we analyzed the performance of the microgripper by FEA. With a 300-μm microaxis, the gripping force was 5.09 N at a 20-μm input displacement. Lastly, we processed a solid model using a LSWEDM, and built an experimental platform to verify the results. The experimental results confirmed the correctness of the theoretical calculation and simulation analysis, and the microgripper successfully clamped a 250-μm metal ball in parallel and stably, which proves the correctness of the design.

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Xiaodong Chen received his B.Eng. degree in mechanical manufacturing and automation from the Northwest University of Technology, Xi'an, China, in 2017. He is currently working toward the M.Sc. degree in Liaoning University of Petroleum and Chemical Technology. His current research interests include flexure mechanisms and MEMS.

Zilong Deng received his B.Eng. degree in mechanical manufacturing and automation from the Liaoning University of Technology, Shenyang, China, in 1993, and the M.Sc. degree in mechanical manufacturing and automation from Jilin University, Changchun, in 2000. He became a professor at the School of Mechanical Engineering, Liaoning University of Petroleum and Chemical Technology in 2013. His current research interests include Mechatronics and flexure mechanisms.