A microcantilever system with slider-crank actuation mechanism

Sensors and Actuators A 226 (2015) 59–68

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Sensors and Actuators A: Physical journal homepage: www.elsevier.com/locate/sna

A microcantilever system with slider-crank actuation mechanism Xing Chen a , Dong-Weon Lee b,∗ a b

Department of Electrical and Computer Engineering, University of Utah, Salt Lake City, UT 84112, USA MEMS & Nanotechnology Laboratory, School of Mechanical Systems Engineering, Chonnam National University, Gwangju 500757, Republic of Korea

a r t i c l e

i n f o

Article history: Received 28 September 2014 Received in revised form 18 February 2015 Accepted 18 February 2015 Available online 25 February 2015 Keywords: Microcantilever Slider-crank mechanism Flexure hinge Big deflection Sharp tip Electrothermal actuator

a b s t r a c t A microcantilever system with new actuation principle is developed for enhancing the range of tip deflection. Differing from conventional springboard-like actuation principle, the proposed microcantilever system is based on a slider-crank mechanism. In this system, out-of-plane deflection of the microcantilever can be translated and amplified from in-plane motion of high-powered thermal actuators, which is located away from cantilever body to ensure both energetic driving force and good thermal isolation. Analytical models of both the microcantilever system and the electrothermal actuator are studied concerning the optimized geometry for desired amplification ratio and balanced force/displacement outputs. The prototype is fabricated by standard lithographic process, and an integrated sharp tip is easily machined by a hybrid wet/dry etching method. The fabricated cantilever is experimentally tested and shows that a 22 ␮m in-plane motion from actuators at an input electrical power of 1026 mW can drive the microcantilever to deflect up to 230 ␮m, demonstrating the expected actuation principle. The developed microcantilever system with integrated sharp tip shows the promise to work as a microprobe in the multifunctional analysis instrument. © 2015 Elsevier B.V. All rights reserved.

1. Introduction Microactuator is the basic element of microelectromechanical systems (MEMS), which introduces mechanical character into microelectronics and microstructures [1]. The invention and development of microactuator is considered as the most distinct symbol of MEMS making it differ from IC technology [2]. In general, microactuation implemented by microactuators can be categorized in light of their motion directions as, in-plane actuation (parallel to the substrate) and out-of-plane actuation (normal to the substrate) [3]. The out-of-plane actuation has particularly shown a broad range of applications, such as optical switches for communications [4], scanning confocal microscopes [5], oscillators [6], digital mirrors for projection displays [7], micro/nanopositioners [8], etc. Its another successful case is the actuation of microcantilevers working for scanning probe microscopy (SPM) family and bio/chemical detection [9], in which the microcantilevers integrated with microactuators can either deflect toward sample surfaces or mechanically resonant for scanning and sensing. This type of self-actuated microcantilevers have demonstrated extraordinary advantages beyond external actuation sources (e.g., piezoelectrical tube or stacked stages) for easing assembly,

∗ Corresponding author. Tel.: +82 62 530 1669; fax: +82 62 530 1689. E-mail address: [email protected] (D.-W. Lee). http://dx.doi.org/10.1016/j.sna.2015.02.029 0924-4247/© 2015 Elsevier B.V. All rights reserved.

accelerating scanning and response speed, and scaling system down to handheld size for on-site detection [10–12]. With continuing attraction of MEMS technology, it is also driving the evolution of self-actuated microcantilevers with new features toward some state-of-the-art applications. One of the most pioneering researches is to extend the powerful physical probing function of SPM with local chemical analysis [13,14], by means of combining it with another analysis instrument, such as mass spectrometer [15]. In this type of multifunctional system, a microprobe comprising a self-actuated microcantilever with big out-of-plane deflection is crucial to bridge two separate operating modes together and switch between each other. In addition, a sharp tip at the end of the microprobe is also important to steer the resolution of topographic imaging in the SPM mode and ionize samples for the chemical analysis mode [16,17]. However, most reported researches have revealed that the selfactuated microcantilever with simple springboard-like structure is inherently forbidden to achieve satisfactory big deflection because of its mechanics nature [18–20], and suggest to looking for an alternative out-of-plane actuation structure or principle. One smart effort to build such a microprobe is by applying the electrostatic pull-in effect and has demonstrated ultra-big deflection up to 400 ␮m, whereas the multi-stepped actuation can’t guarantee linear deflection increment and the involvement of special focused ion beam (FIB) machining keeps it from standard lithographic process [21]. Actually, some innovative designs by employing linkage

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Fig. 1. Schematics showing the structure of microcantilever system, the close-up of flexure hinge array, and actuation principle including states (a) before actuation and (b) after actuation.

mechanism have already been demonstrated in microsystems to enhance out-of-plane motion. For example, an out-of-plane rotational micromirror driven with in-plane electrostatic comb-drive actuators through surface-micromachined hinges has been developed to extend scanning angles [22]. Slider-crank together with other linkage system (such as Sarrus linkage or four/five-bar linkage) were applied in microrobotics to convert between circular and linear motions [23–25]. In another flying robotic insect work, a compliant mechanism based on slider-crank has exhibited flapping angle range up to 75◦ by amplifying small linear displacement from PZT actuators [26]. Inspired by the above work, a new actuation principle is introduced to microcantilever for enhancing its tip deflection range. Unlike conventional self-actuated microcantilevers dedicated to springboard structure, this new microcantilever system is based on a slider-crank mechanism, in which the out-of-plane deflection of microcantilever can be driven by in-plane locomotive microactuator. A microprobe integrating microcantilever, microactuators, and nanoscale sharp tip is fabricated with standard lithographic process without any assembly or involvement of special machining. This microprobe is targeted at the potential application in the multifunctional analysis instrument [13]. 2. System design and modeling 2.1. Working principle A slider-crank mechanism is so often used in mechanical machines to convert linear reciprocating motion into rotary motion or vice versa, which mainly consists of a rotary crank, a slider, a connection rod, and pivots. It is interesting to note that an elaborately designed slider-crank mechanism can have a controllable or even amplified transmission ratio between rotary angle and in-plane displacement. If this feature can be grafted into the self-actuated microcantilever, it would be helpful to enhance microcantilever deflection from limited motion of microactuators. Therefore, a mimetic system based on slider-crank mechanism is introduced in our work for microcantilever actuation.

As shown in Fig. 1(a), the proposed microcantilever system is designed to consist of equivalent components of a typical slider-crank mechanism, including a curled double-legged cantilever (working as crank), in-plane electrothermal actuator array (as slider), flexure hinge array (as pivots), a connection rod for mechanical link, and integrated sharp tip for potential scanning and manipulation applications [18]. The actuation principle is displayed in Fig. 1(b) that, when the in-plane electrothermal actuator array is actuated, it drives horizontally and drags the connection rod to move, together with the cantilever. According to force balance, the double-legged cantilever is subject to both the driving force from connection rod and the resultant counterforce at its fixed ends, which are equal in magnitude but opposite in direction. These two parallel forces exerting on different horizontal planes of the curled cantilever do not share the same line of action, so they constitute a force couple as the bending moment, responsible for deflecting cantilever. Based on this actuation principle, the deflection of cantilever y can be expressed by beam theory as y =

Mlc2 , 2EIc

(1)

where M is the bending moment (force couple) that is equal to the multiplication of driving force F by initial bending of the curled cantilever y0 when actuation starts, lc is the effective length of cantilever crank (from flexure hinge part of cantilever to its joint with connection rod), and EIc is the flexural rigidity of flexure hinges at cantilever’s fixed ends. It should be noted that the control of initial bending y0 by residual stress of deposited metal (will be discussed in detail later) is directly associated with the bending moment and deflection range of the cantilever. If the shape of the cantilever is straight without initial bending (y0 = 0), no matter how big the driving force F is the bending moment M keeps to be null, which condition is similar to the “dead point” in slider-crank mechanism. In order to demonstrate the functionality of this new microcantilever system, simplified solid models based on the structure in Fig. 1 are built and analyzed by finite element method (FEM) software ANSYS 9.0 with element type of SOLID92. As evident from one of the FEM results shown in Fig. 2(a), an input in-plane (x-axis) force

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tenable. From the geometry standpoint, the cantilever deflects in a circular motion with the flexure hinges as its center and lc as the effective radius, while two ends of the connection rod experience not only the same circular motion but also extra displacement of actuators, so lr must be bigger than lc for free actuation. If lr is smaller than or equal to lc , the connection rod suppresses the deflection and compresses the cantilever axially, which explains the buckling issue in Fig. 2(b). Therefore, the connection rod is kept longer than the cantilever in all the following designs (lr > lc ). The dependence of amplification ratio on device geometry can be then expressed by constructing partial derivatives of Eq. (4) with respect to each of geometry variables as below (a) The dependence on lr

 dy  dx

dlr

3

lc cos2  sinlr

= − − lc

2

2

+ lc sin2 

sin cos

lr 2 −lc 2 sin2 



2

2

lr − lc sin2 

 32

< 0.

(5)

(b) The dependence on lc

 dy  dx dlc

Fig. 2. (a) Actuation principle of the microcantilever system with a single flexure hinge proved by FEM simulation result, (b) one example of failed simulation model, and (c) structure of the single flexure hinge and its FEM simulation.

−

2.2. Optimization of the device geometry

 

y +

lc

2

− y2

+x

2

2

= lr ,

(2)

where y is the total cantilever deflection (y + y0 ), x is the total displacement of connection rod end from the hinge (in-plane displacement x + initial distance x0 ), and lr is the connection rod length, as denoted in Fig. 2(a). The equation can be parameterized by involving the cantilever slope angle  as y = lc sin, x=



2

2

lr − lc sin2  − lc cos.

(3)

Next, the amplified transmission ratio or amplification ratio between the out-of-plane cantilever deflection and the in-plane displacement of the microactuators is expressed by their derivative as lc cos dy = . 2 dx l sin cos c − + lc sin

−

2

2

⎡

lr − lc sin2 

⎢ ⎣

lc sin3 cos

+ lc sin

3

 2 3 2 lr − lc sin2  2 

−

−

2

2

2

2lc sin cos



lc sin cos lr − lc sin2 

⎤ 2

2

2

lr − lc sin 

⎥ ⎦

(6)

+ sin ⎥ > 0.

2 + lc sin2 

(c) The dependence on 

Through FEM simulations, the problem also appears that the transmission ratios between out-of-plane deflection and in-plane displacement are uncertain and vary with different device geometries, and the buckling of cantilever even occurs in some cases (see the FEM result in Fig. 2(b)). In order to fully understand the geometry effect to the transmission ratio, and take advantage of the slider-crank mechanism to amplify cantilever deflection from limited in-plane displacement, a pseudo-rigid-body model of the device is considered and its geometry is expressed as [27]



2

lc sin cos

lc cos ⎢−

acting at one end of the connection rod leads to the desired out-ofplane (y-axis) deflection of the cantilever, with the same way to a slider-crank mechanism (more details of this simulation will be given later).

2

cos

=

(4)

lr 2 −lc 2 sin2 

Eqs. (3) and (4) reveal a fact mathematically that the value of lr should be larger than that of lc ; otherwise the equation is not

dy dx

d

=−

⎡ ⎢

lc cost ⎣− −

lc sint 2

lc sin cos

−

4

2

2

+ lc sint 2

2

lc sin t cos2 t



lr

2

⎤

lr − lc sin 

3 2 2 − lc sin t 2

−

2

lc cos2  2

2

+ 2

lr − lc sin 





2

⎥

2

lc sin  2

2

2

lr − lc sin 

2

+lc cos ⎦ <0.

2

lc sin cos

−

2

2

+ lc sint 2

lr − lc sin 

(7)

The above calculations present three important design guidelines to optimize the structural geometry of this new cantilever system for gaining a desired amplification ratio: (1) With the increase of the connection rod length, the amplification ratio decreases. From this point, the length of the connection rod should be kept as small as possible, while it should be at least larger than the length of cantilever, based on the earlier discussion. (2) When the cantilever is designed to be longer, a higher amplification ratio can be obtained. It is obvious from theorem of similar triangles that a longer cantilever has bigger deflection at the same slope angle. (3) A smaller initial slope angle  0 or bending y0 of cantilever enables a higher amplification ratio, whereas it needs a bigger driving force for generating sufficient bending moment to start the actuation according to the Eq. (1). During the actuation process, as the slope angle goes up, the amplification ratio gets decreased.

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2.3. Flexure hinge design As can be seen from above pseudo-rigid-body models of this new microcantilever system, the flexure hinge plays an important role as a pivot to translate the in-plane motion into bending moment. Because of frictionless and backlashless transmission, and easy microfabrication, flexure hinge is more preferred for microdevices in contrast to the revolute joints [22]. Given the fact that flexure hinge is a thin membrane linking and rotating two adjacent rigid parts through flexing, we find out that there exists an optimum hinge length in the design. On one hand, when hinge deforms into arc with its length as perimeter, the effective radius of curvature should be larger than the depth of flexure hinge for adapting to 90◦ angled deformation, and a longer hinge also offers a higher flexibility to the microcantilever system. On the other hand, the slender hinge may easily experience buckling when it becomes longer, as the critical force of the instability is proportional to the square of length [28]. In a model following above guidelines, a single flexure hinge (25 ␮mlong, 10 ␮m-wide, and 1.5 ␮m-thick) is designed at each fixed end of a 520-␮m long, 60 ␮m-wide, and 3 ␮m-thick cantilever, with a 590 ␮m-long, 50 ␮m-wide, and 3 ␮m-thick connection rod, as illustrated in Fig. 2(a). Its FEM simulation shows that a 34 ␮m inplane motion at one end of the connection rod when exerted with 600 ␮N driving force, gives the cantilever an amplified deflection of 433 ␮m and deforms single hinge by a slope angle of 41◦ , which verifies the designed actuation principle. However, the local von Mises stress at the single hinge reaches up to 1.5 GPa (Fig. 2(c)), which is much beyond the safe value for single crystal silicon-based microdevices (about 600 MPa) considering the defects brought by practical fabrication. In order to solve the problem, serial flexure hinge array (like springs in series) is considered instead of a single one, as shown in Figs. 1 and 4. The total values of deflection yhinge and slope angle  hinge of this serial compliant mechanism can be expressed by Castigliano’s second theorem with each hinge li as a node Fig. 3. Plotted results of (a) deflections y of a cantilever (lc = 520 ␮m) with respect to different connection rod lengths lr when the driving displacement is constant, and (b) driving displacements x required to deflect the cantilever up to a slope angle of 90◦ with each increment of 5◦ .

yhinge =

i=l n

hinge = To validate these analytical deductions, a microcantilever system with different geometry dimensions is modeled as a pseudo-rigid-body and studied by graphic approach. Fig. 3(a) shows the simulated deflection results of a cantilever (lc = 520 ␮m) with respect to different connection rod lengths (lr = 590–890 ␮m, each with an increment of 50 ␮m). In spite of experiencing the same displacement (34 ␮m), the cantilever deflection drops from 433 to 352 ␮m with the increased connection rod length, i.e., decreased amplification ratio, which agrees with #1 guideline. In the second study, the displacements needed to deflect the cantilever up to 85◦ slope angle with each increment of 5◦ is investigated in a microcantilever system with constant dimensions (lc = 520 ␮m, lr = 590 ␮m, and  0 = 5◦ ). As shown in Fig. 3(b), the displacement increases nonlinearly with the slope angle. A small displacement of ∼1 ␮m can drive the cantilever to deflect from the slope angle of 5◦ –10◦ at the beginning, by contrast, the displacement required for the same increment of 5◦ but from 80◦ to 85◦ rises dramatically to be ∼36 ␮m, which presents a good agreement with #3 guideline. It also shows the value of amplification ratio decreases below 1 when the slope angle is over some point (65◦ in this structure); in other words, the input displacement becomes bigger than output deflection in that condition, which suggests that this actuation principle can work effectively only at a relatively moderate range of deflection, and the initial bending y0 should be controlled below this critical point.

 n 

 i=l

0

li

0 li

Mi ∂Mi dyi + EI ∂F

Mi ∂Mi dyi . EI ∂F

 0

li

Ni ∂Ni dyi EA ∂F



, (8)

Eq. (8) shows that the total displacement of serial compliant system equals to the sum of all corresponding deformations and rigid-body motions over each individual node [29]. In other words, when the flexure hinge array is subject to the same total amount of deformation, each hinge undertakes smaller mechanical strain energy in comparison to the single hinge case. Therefore, the microcantilever system is redesigned by replacing with an array of five hinges (20 ␮m-long, 40 ␮m-wide, and 1 ␮m-thick), and adding an extra 350 ␮m-long extensional beam at the free end of cantilever for increasing amplification ratio according to #2 guideline. Fig. 4 shows the FEM analysis of the new configured microcantilever with the same 600 ␮N force input, and the results reveal a larger cantilever deflection (∼500 ␮m) while decreased maximum von Mises stress (591 MPa) below the safe value compared to the previous single hinge design, and a satisfactory amplification ratio of ∼8.3 between out-of-plane deflection and in-plane displacement (∼60 ␮m). 2.4. Microactuator design As another important component, the electrothermal actuator is chosen as driving source because of its simple structure, big displacement and large driving force output [30–32], in contrast

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Fig. 4. The flexure hinge array structure and its employment in a microcantilever with simulated deflection and stress distribution at a driving force of 600 ␮N.

to other types of microactuators such as electrostatic comb-drive actuator. A V-shaped electrothermal actuator, with two angled thin legs, can transfer the compressive forces generated by thermal expansion into linear motion to drive the cantilever in this study. It was reported experimentally that the output force and displacement of the V-shaped electrothermal actuator are strongly dependant on its geometry [33], we thus conduct a theoretical study on that for understanding the fundamentals and providing a guideline to serve for our design. Fig. 5(a) shows a sketched single V-shaped electrothermal actuator with two symmetrical legs la at a tilt angle ˇ with respect to central line. During actuation, the total forces toward x and z axes are expressed as ˙Fz = 0, ˙Fdriving = 2Fbeam cos(ˇ − ˛)

[33]. The V-shaped electrothermal actuator in our microcantilever system, therefore, is designed with long legs and big tilt angle to ensure a big displacement; meanwhile, a large number of actuators in an array is used to provide sufficient driving force, compensating for the decreased driving force from big angled actuator. Details of the designed parameters can be found in Table 1.

(9)

where Fdriving is the driving force from a single V-shaped electrothermal actuator toward x-axis, Fbeam is the joule heating force per leg, and ˛ is the variable angle with actuation. The actuator geometry in Fig. 5(b) can be expressed as la = z/sinˇ,

z = xtanˇ

(10)

Eq. (10) can be further processed by derivatives of the parametric equations as dx 1 = −z , dˇ sin2 ˇ dla cosˇ = −z . dˇ sin2 ˇ

(11)

By the chain rule, Eq. (11) gives the below equation to describe the relationship between the displacement of actuator at x-axis dx and the thermal expansion increment of leg dla as dx 1 = dla cosˇ

(12)

Eq. (9) depicts that when Fbeam is constant, Fdriving decreases with bigger tilt angle ˇ. The driving force nonlinearly increases with the declining ˇ in the actuation process. While, it presents an opposite effect of the tilt angle to the actuation displacement as indicated in Eq. (12). Our analytical results show good agreements with the experimental findings reported by other research

Fig. 5. Sketch of a V-shaped electrothermal actuator with its geometry before and after actuation describing (a) driving force and (b) deformation.

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Table 1 Structural parameters of the device. Parameters Cantilever part Length of cantilever Width of cantilever Thickness of cantilever Length of extensional beam Width of extensional beam Thickness of extensional beam Length of connection rod Width of connection rod Thickness of connection rod Length of hinge Width of hinge Thickness of hinge Number of hinges in an array Actuator part Tilt angle Length of leg Width of leg Thickness of leg Number of actuator in an array

Symbols

Values

lc wc tc le we te lr wr tr lh wh th Nh

520 ␮m 60 ␮m 5 ␮m 350 ␮m 250 ␮m 5 ␮m 590 ␮m 60 ␮m 5 ␮m 20 ␮m 40 ␮m 1 ␮m 6

ˇ la wa ta Na

88.94◦ 1202 ␮m 8 ␮m 5 ␮m 50

3. Device fabrication Following the analytical study of the pseudo-rigid-body model for enhanced deflection/displacement amplification ratio and FEM simulation for structural optimization of the microcantilever system and particularly the flexure hinge, the device dimensions were finalized as described in Table 1. The starting material for fabrication was a silicon-on-insulator (SOI) wafer with 5 ␮m-thick (1 0 0) oriented and 0.025 -cm highly-doped device layer, 2 ␮m-thick buried oxide (BOX) layer and 350 ␮m-thick handling layer, respectively, and five photomasks were used in the fabrication. Flexure hinge in MEMS is usually fabricated by reducing local thickness with lowered stiffness, and the same way was used in this work to carve 126◦ inclined groove structure as the flexure hinges by wet etching [34]. With a 200 nm-thick silicon dioxide layer as etching mask, the SOI wafer was operated into tetramethylammonium hydroxide (TMAH) bath at 80 ◦ C with etching rate of 0.36 ␮m/min to etch away around 3.5–3.6 ␮m thick top silicon layer (designed to be 4 ␮m), as shown in Fig. 6(a). Based on the anisotropic etching on different crystal planes between (1 1 1) and (1 0 0), the etching pattern was precisely oriented to the <1 1 0> direction on the (1 0 0) device layer, forming the V-shaped groove hinges. After removing the first SiO2 etching mask, another thicker SiO2 (1 ␮m) was prepared by wet oxidation at 1000 ◦ C for 4.5 h, which was used as the etching mask for the following dry etching and to trim the thickness of hinges into the designed value as shown in Fig. 7(a). Since there was a big step height between the V-shaped groove hinges and other unetched parts, 6 ␮m-thick photoresist AZ4620 was used to cover the entire surface thoroughly, and was then lithographically patterned. It was emphasized earlier about the importance of fabricating a finely featured tip at the end of the microcantilever for its expected functionality. Unfortunately, the size of tip apex is hindered by the limitation of UV photolithography, for example, the minimum spatial pattern from our photolithography system can merely reach as small as 1–2 ␮m. A hybrid technology was thus introduced in for the sharp tip fabrication [35], as shown in Fig. 6(b and c). In this method illustrated detaily in Fig. 8, after hardbaking the AZ4620 pattern at 130 ◦ C for 30 min on a hotplate, the SiO2 layer beneath photoresist was over-etched in the buffered hydrofluoric acid (BHF). Due to the isotropic etching nature of SiO2 in the BHF, the SiO2 neck was laterally etched and became narrow over time. Finally, the neck was etched through, leaving a sharp cone-like SiO2 pattern. The condition of over-etching was optically observed under the optical

Fig. 6. Main steps of the fabrication process and the corresponding mask layers (solid purple patterns are protected area and openings are substrate to be etched).

microscopy, a distinct color difference allows for real-time vision of the SiO2 pattern beneath photoresist pattern, which is clearly shown in Fig. 8(Step 1). After photoresist removal from patterned SiO2 as etching mask (Fig. 8(Step 2)), the reactive ion etching (RIE) using SF6 gas was followed to over-etch the top Si device layer up to the BOX layer for shaping the Si cantilever body and tip (Figs. 6(c) and 7(b)). Due to the isotropic etching of the RIE, the Si tip underneath the cone-like SiO2 pattern was further sharpened, presenting a nanometer-scale apex whose shape and aspect ratio were controllable by adjusting the recipes in both wet and dry etchings, as illustrated in Fig. 8(Step 3). Another low temperature oxidation (at 950 ◦ C) method for sharpening an even fine tip was optional after this RIE process [36]. To fabricate electrical interconnections and generate an initial bending that is crucial for the actuation principle in this new system (see Eq. (1)), a thin metal film was introduced on the Si substrate for forming a bimorph bent beam [37]. It was reported that the initial bending degree of such a Si/metal bimorph cantilever is dependent on the thermal expansion coefficient (TEC) difference between metal and silicon as well as temperature change during deposition, and one with a constant silicon layer thickness can be maximized

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65

Fig. 7. SEM images illustrating the results of main fabrication process.

by optimizing metal layer thickness [38]. Since the temperature controlling was regrettably unavailable for our deposition facility, choice of suitable metal and its optimized thickness became important. Considering that, aluminum was chosen due to the biggest difference in TEC between it (23.1 × 10−6 /◦ C) and Si (2.6 × 10−6 /◦ C). The optimum Al/Si thickness ratio of 0.6 (i.e., a 600 nm-thick Al on ∼1 ␮m-thick Si hinge) for getting the maximum initial deflection was calculated with respect to width ratio (Al/Si: 18 ␮m/40 ␮m) and Young’s modulus ratio (Al/Si: 69 GPa/160 GPa) [38]. The Al film was deposited by e-beam evaporation and then patterned by wet

etching in the Al etchant at room temperature for a precise etching control, as shown in Figs. 6(d) and 7(c). A deep reactive ion etching (DRIE) process was followed to shape V-shaped electrothermal actuator array with uniform Si thickness identical to that of the device layer on the SOI wafer (Fig. 6(e)). The Bosch DRIE technique was applied to fabricate the steep-sided actuators that move merely with horizontal motion (Fig. 7(d)). Finally, the device was released by backside wafer-through etching. The same Bosch DRIE process was used to etch the Si handling layer up to the BOX layer as etching barrier, followed by

Fig. 8. Detailed description of hybrid technique to fabricate sharp silicon tip.

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Fig. 9. Fabricated devices. (a) Photograph from the backside of a batch of devices. (b) SEM image of the V-shaped grooves acting as flexure hinge array, and the inset shows the aluminum interconnections over the hinges. (c) SEM image of a slightly pre-bent cantilever. (d) SEM image showing the configuration of sharp tip at the end of the cantilever and the actuator array.

BHF etching to strip the entire exposed SiO2 layer, as shown in Fig. 6(f).

4. Results and discussion Fig. 9 illustrates the fabricated devices. Due to some fabrication variations for each device from different fabrication batches, the fabricated microcantilevers presented different degrees of initial bending; while only few of them showed satisfactory initial bending, which were sorted out and wired by silver conductive epoxy (CW2400, Chemtronics, USA) for the subsequent test. As our research interest for the microprobe was its static deflection, only the DC response of the microcantilever system was tested. In the experimental setup, a sourcemeter (KEITHLEY 2400, Keithley Instruments, Inc., USA) was used to supply DC current to the V-shaped electrothermal actuators and record the feedback voltage, in which way the electrical power consumed for joule heating can be also calculated. An optical microscope (ECLIPSE L150, Nikon, Japan) equipped with a charge-coupled device (CCD) camera and image-processing program was applied to observe the in-plane displacement of the V-shaped electrothermal actuators, capture image data, and read the values. The out-of-plane deflection of the microcantilever was measured by both the laser displacement sensor (KEYENCE, LK series, Japan) and single-point laser vibrometer (OVF-534, Polytec Inc., USA). The extensional beam at the end of the microcantilever system was designed to have a sufficient space (350 ␮m long and 250 ␮m wide) for locating the laser spot from deflection measurement. Fig. 10 shows the optical images of the V-shaped electrothermal actuator array and microcantilever parts captured before and after actuation. By comparison of two states, it can be clearly seen that when getting heated by DC current input, the actuator array moved with translational motion and its legs tilted by an angle of ˛, whose actuation drove the microcantilever to deflect upwards.

This experiment demonstrates the proof-of-concept for the actuation principle expected from this new microcantilever system, and verifies that the in-plane motion of actuators can drive the out-ofplane deflection of the cantilever. The corresponding deflections and displacements as a function of power consumptions were characterized and plotted in Fig. 11. The maximum microcantilever deflection was approximately 230 ␮m when the displacement of actuator was 22 ␮m, at the input DC current of 50 mA and a resultant electrical power of 1026 mW. Afterwards, the further motion of actuator by increased input current was found to cause fracture of the hinges, and the device hence failed. When the input

Fig. 10. Optical images capturing the V-shaped electrothermal actuator array and microcantilever before and after actuation. One broken leg acts as a reference to show the displacement of the actuator array. The legibility at a fixed microscopy focusing suggests the degree of deflection.

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Fig. 11. Experimental results showing the relationship between input power with displacement of V-shaped electrothermal actuator array and cantilever deflection, and comparison with FEM simulation results.

DC current reached ∼96 mA, the V-shaped electrothermal actuator array was observed to burn, indicating a reference temperature of ∼1000 ◦ C at that moment; while there was no sign of any thermal damage to other regions, which suggested a good thermal isolation between high-powered actuation source and other components in this device. By comparing the experimental results with the FEM ones in Fig. 11, we can find out that the experimental deflections were smaller than the simulated one when actuated at the same displacements, and the deflection/displacement amplification ratio at this deflection range (∼10.5) was smaller than ∼16 from the simulated results. One possible reason is that the actual thickness of hinges was bigger than the designed one. Also, the actual maximum deflection until fracture was smaller than half of the simulated value (∼500 ␮m), and the deflection increment was not smooth. The discrepancy could be attributed to the fabrication process in practice. In the simulation, the flexure hinges can stand the larger cantilever deflection considering von Mises stress limit of ideal structure made of silicon; while the fabricated flexure hinges suffered from many practical defects such as pin-holes after experiencing wet etching, and particularly, the stress concentration at the sharp corner of groove could substantially deteriorate the mechanical properties of structure [39], making hinges vulnerable to big deformation. This issue could be improved by graving round shaped hinges with isotropic wet etching. Moreover, polymer hinges could be introduced into this microcantilever system, which has much lower Young’s modulus (e.g., 3.5 GPa for SU8 [40,41], and 1 MPa for PDMS [42]) than that of Si (160 GPa) for efficiently increasing the deflection range of cantilever without plastic deformation or fracture. Another issue is that the fabricated devices have undesired small initial bending, which inevitably increases the driving force to start actuation. In order to enhance the degree of initial bending with a controllable process, the RF magnetron sputtering technique with temperature adjustable base should be considered in the future work. 5. Conclusions This study has proposed a microcantilever system with new actuation method, and has investigated its theoretical models with experimental validation. Based on a slider-crank mechanism,

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the new microcantilever system enabled out-of-plane motion by in-plane actuation, making it different from other conventional actuation principles working in springboard-like microcantilevers. It presented two main advantages: first, the actuation source is away from the microcantilever, which allows one to equip high-powered actuators without concerning any thermal influence to the microcantilever body or samples; second, the feature of amplified transmission ratio between the out-of-plane deflection and in-plane displacement is greatly beneficial to enhance the deflection range of microcantilever from limited motion of microactuators. The proposed microcantilever system was designed with equivalent components of a typical slider-crank mechanism, whose pseudo-rigid-body model was studied to understand the fundamental principle and obtain the optimized structure geometry. Through the theoretical study, it revealed that the working effectiveness and amplification ratio of the system are highly dependent on the geometry parameters, such as the length of connection rod and cantilever crank as well as degree of initial bending, and this actuation principle can work most effectively at relatively moderate range of deflection. The designed system was optimized by FEM simulation, particularly concerning the elastic deformation of flexure hinges whose maximum von Mises stress was controlled under 600 MPa. Geometry of V-shaped electrothermal actuator was studied to balance the desired driving force and displacement. After the theoretical analysis of this system, a standard silicon fabrication process with five photomasks was used to machine the microprobe prototype. A hybrid wet/dry etching method was applied to fabricate the nanometer-scale sharp tip at the end of the microcantilever. The fabricated device was experimentally tested, and showed the maximum microcantilever deflection up to 230 ␮m under a 22 ␮m displacement of the Vshaped electrothermal actuator array with an amplification ratio of 10.5, which validates the conceived actuation principle and agrees with our theoretical model. The discrepancies between the simulated and experimental results were analyzed, and some suggestions based on the problems present in the current device were discussed to be improved in the future work.

Acknowledgments This study was supported by a grant of the Korean Health Technology R&D Project (HI13C1527), Ministry of Health & Welfare and the National Research Foundation (NRF) grant (No. 2012R1A2A2A01014711) funded by the Korea government.

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Biographies Xing Chen received the B.S. degree in mechanical engineering from Kunming University of Science and Technology, Yunnan, China, in 2004, and his M.S. and Ph.D. degrees in mechanical engineering from Chonnam National University, Gwangju, Korea, in 2008 and 2012, respectively. From 2012 to 2014, he was a Post-Doctoral Research Fellow with the Department of Electrical and Computer Engineering, University of British Columbia, Vancouver, Canada. He is currently a research associate with the Department of Electrical and Computer Engineering, University of Utah, Salt Lake City, USA. His research interests include microactuators, microsensors based on graphene and stainless steel, scanning probe microscopy, and implantable biomedical microdevices. Dong-Weon Lee received his Ph.D. degrees in Mechatronics engineering from Tohoku University, Sendai, Japan in 2001. He has been a Professor of Mechanical Engineering at Chonnam National University (CNU), South Korea since 2004. Previously, he was with the IBM Zurich Research Laboratory in Switzerland, working mainly on microcantilever devices for chemical AFM applications. At CNU, his research interests include smart cantilever devices, miniaturized energy harvester, smart structures & materials, and nanoscale transducers. He is a member of the technical program committee of IEEE Sensors Conference, Transducers, and Microprocesses and Nanotechnology Conference.