DSpace at VNU: Electrothermal Microgripper With Large Jaw Displacement and Integrated Force Sensors

DSpace at VNU: Electrothermal Microgripper With Large Jaw Displacement and Integrated Force Sensors tài liệu, giáo án, b...

1546 JOURNAL OF MICROELECTROMECHANICAL SYSTEMS, VOL 17, NO 6, DECEMBER 2008 Electrothermal Microgripper With Large Jaw Displacement and Integrated Force Sensors Trinh Chu Duc, Gih-Keong Lau, J Fredrik Creemer, Member, IEEE, and Pasqualina M Sarro, Fellow, IEEE Abstract—The novel design of a sensing microgripper based on silicon-polymer electrothermal actuators and piezoresistive force-sensing cantilever beams is presented The actuator consists of a silicon comb structure with an aluminum heater on top and filled polymer in between the comb fingers The sensor consists of a silicon cantilever with sensing piezoresistors on top A microgripper jaw displacement up to 32 µm at a 4.5-V applied voltage is measured The maximum average temperature change is 176 ◦ C The output voltage of the piezoresistive sensing cantilever is up to 49 mV at the maximum jaw displacement The measured force sensitivity is up to 1.7 V/N with a corresponding displacement sensitivity of 1.5 kV/m Minimum detectable displacement of nm and minimum detectable force of 770 nN are estimated This sensing microgripper can potentially be used in automatic manipulation systems in microassembly and microrobotics [2008-0064] Index Terms—Electrothermal actuator, microgripper, piezoresistive sensor, polymeric actuator, sensing microgripper I I NTRODUCTION W HEN manipulating micro-objects, the dexterity, accuracy, and speed are considerably improved when the force on the objects can be sensed and controlled in real time [1] The development of such miniaturized manipulators is of great interest for operating on living cells, minimally invasive surgery, microrobotics, and microassembly The manipulation of micro-objects with traditional microgrippers without a built-in force sensor normally requires a camera to obtain visual feedback This approach results in a 2-D image The depth perception of the contact between the manipulating tool and the object being manipulated is lost, making it difficult to identify the position of the tool [1] Moreover, only displacements and not force can be detected A microgripper with a built-in force sensor can address this limitation and, thus, is suitable for holding objects firmly, while avoiding any squeezing of delicate objects Manuscript received March 12, 2008; revised July 22, 2008 Current version published December 4, 2008 Subject Editor C.-J Kim T Chu Duc is with the Faculty of Electronics and Telecommunication, College of Technology, Vietnam National University, Hanoi, Vietnam (e-mail: trinhcd@vnu.edu.vn) G.-K Lau is with the School of Mechanical and Aerospace Engineering, Nanyang Technological University, Singapore 639798 (e-mail: mgkLau@ ntu.edu.sg) J F Creemer and P M Sarro are with the Electronic Components, Technology and Materials Laboratory, Delft Institute of Microsystems and Nanoelectronics, Delft University of Technology, 2628 CT Delft, The Netherlands (e-mail: j.f.Creemer@tudelft.nl; p.m.Sarro@tudelft.nl) 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.2008.2007268 In recent years, several designs of microgrippers with force feedback have been demonstrated A force-sensing microgripper for minimally invasive surgery application [2] employs piezoelectric actuation with strain gauge sensors on the side wall of the structure It is capable of actuating at high frequency (hundreds of hertz) with very high driving voltage In [3], a similar design device was presented It uses electromagnetic actuation and piezoelectric force sensing It generates large displacements at low voltage and a linear sensing output However, the main limitations of the aforementioned devices are the incompatibility with CMOS technology and rather large dimensions Electrothermal actuators with builtin piezoresistive force sensors were presented in [4] and [5] The jaw displacements and output sensing voltages are rather small, limiting their application An electrostatic microgripper with an integrated capacitive force sensor is presented in [6] This device is capable of motion up to 100 μm with a force sensitivity of 4.41 kV/m and a corresponding 70-nN forcesensing resolution However, the limitations of this device are its large size and complicated electronic circuit required by the electrostatic method used This paper presents a novel sensing microgripper based on silicon-polymer electrothermal actuators [7] and piezoresistive force-sensing cantilever beams [8] The proposed sensing microgripper is capable of providing a large jaw displacement and output sensing voltage This device is capable of monitoring the jaw displacement and resulting applied force The device is made on silicon-on-insulator (SOI) wafers with a fabrication process compatible with CMOS technology II D ESIGN In Fig 1, a schematic drawing of the sensing microgripper is shown The structure is based on the combination of siliconpolymer electrothermal microactuators and piezoresistive lateral force-sensing cantilever beams When the electrothermal actuator is activated, the microgripper’s arm and also the sensing cantilever are bent This causes a difference in the longitudinal stress on the opposite sides of the cantilever This changes the resistance values of the sensing piezoresistors on the cantilever The displacement of the microgripper jaws can be monitored by the output voltage of the Wheatstone bridge of the piezoresistive sensing cantilever beam The contact force between the microgripper jaws and clamped object is then determined from the displacement and stiffness of the microgripper arm [9] 1057-7157/$25.00 © 2008 IEEE CHU DUC et al.: MICROGRIPPER WITH LARGE JAW DISPLACEMENT AND INTEGRATED FORCE SENSORS 1547 TABLE I GEOMETRY OF THE SENSING MICROGRIPPER Fig Schematic drawing of the sensing microgripper [10]–[12] When the heater is activated, the generated heat is efficiently transferred to the surrounding polymer through the deep silicon comb finger structure that has a large interface area with the polymer layer The polymer layers expand along the lateral direction causing a bending displacement of the actuator arm As the polymer, we have selected SU8 2002 (Microchem Inc.) Its low viscosity (7.5 cSt) is specifically developed to produce thin layers (2–3 μm) [13] and is low enough for the void-free filling of the 3-μm-wide trenches The main properties of the materials used are summarized in Table II This electrothermal microgripper can be actuated with a low driving voltage, power consumption, and operating temperature Fig Front- and cross-side views of a sensing microgripper arm with geometry symbols and parameters B Thermomechanical Finite Element Modeling Fig shows the front- and cross-side views of the sensing microgripper design The geometrical parameters are given in Table I A Silicon-Polymer Electrothermal Microactuator The microgripper is designed in normally open operating mode Each actuator has a silicon comb finger structure with the aluminum metal heater on top A thin layer of silicon nitride is employed as the electrical isolation between the aluminum structure and the silicon substrate The gaps between the silicon comb fingers are filled with SU8 polymer (see Fig 1) Each actuator consists of 41 silicon comb fingers with SU8 polymer layers in between The silicon fingers are μm wide, 75 μm long, and 30 μm thick The SU8 polymer layers are μm wide The length/width (Lcomb /HSU8 ) and height/width (T /HSU8 ) ratios of the polymer layer are 25 and 10, respectively (see Table I) These values, being greater or equal to 10, satisfy the prerequisite for the maximum constraint effect To simulate the performance of the proposed sensing microgripper, a finite element modeling software COMSOL (Comsol Inc.) is used The related material properties (see Table II) are assumed to be temperature independent The 3-D thermomechanical model is used to determine the “steadystate” temperature distribution within the actuator and sensing cantilever structures The thermal expansion and resulting actuator displacement are computed based on the temperature results [12] The actuator is assumed to be immersed in air The silicon comb structure acts as heat source and the rest of the gripper arm as a heat sink The substrate is assumed to be thermally grounded, and therefore, the temperature of the device anchors is fixed and equal to the ambient temperature The heat dissipation through convection and radiation into the atmosphere can be ignored in comparison to the heat loss due to conduction in the actuator anchors when the working temperature is below 500 K [23]–[25] More details on the simulations can be found in [7] and [12] 1548 JOURNAL OF MICROELECTROMECHANICAL SYSTEMS, VOL 17, NO 6, DECEMBER 2008 TABLE II PROPERTIES OF SILICON, ALUMINUM, AND SU8 Fig Steady-state thermal profile on actuator and cantilever Fig shows the simulated steady-state temperature profile along the line through the middle point of all comb fingers of the actuator and sensing cantilever when the microgripper jaw displacement is 25 μm at the applied voltage of 4.5 V The maximum temperature change of 195 ◦ C in the actuator occurs approximately at 300 μm from the anchor along its longitudinal axis The temperature in the cantilever changes linearly from ambient temperature at the anchor to 189 ◦ C at its tip The simulated temperature at the microgripper jaws is 190 ◦ C The average working temperature in the electrothermal actuator is estimated from the aforementioned simulated temperature at the middle point of all comb fingers Fig shows the simulated microgripper jaw displacement versus the average temperature change and also the maximum temperature change The maximum displacement of the two microgripper jaws djaws is 25 μm at the average temperature change of 150 ◦ C, corresponding to a maximum temperature change of 195 ◦ C (see Fig 3) The displacement of the sensing cantilever dcan is also simulated and shown in Fig The maximum sensing cantilever tip displacement is 9.3 μm when the microgripper jaw displacement is 25 μm (see Figs and 4) The initial gap between the two jaws of the microgripper is designed to be 40 μm Therefore, this proposed sensing microgripper is expected to be capable of gripping micro-objects with a diameter of 15–40 μm The simulated static lateral stiffness Kl of the sensing microgripper arms is 1.8 kN/m This value is obtained using a mechanical model with an external lateral load at the microgripper jaws The maximum output force of this microgripper is calcu- Fig Simulated microgripper jaw displacement and the cantilever tip displacement versus the average working temperature change and maximum temperature change lated through the maximum displacement of the microgripper arm and its stiffness of 22.5 mN C Piezoresistive Force-Sensing Cantilever Beam The force sensor design is based on the lateral force-sensing piezoresistive cantilever beam [8], [26] The four piezoresistors are located on the cantilever beam structure and connected to create a Wheatstone bridge (see Figs and 2) The piezoresistors are aligned along the [110] direction in the (001) crystal plane of the silicon wafer The resistor pair located on the cantilever are stress-sensing resistors When the electrothermal actuator is activated, the cantilever beam is bent parallel to the wafer surface Therefore, the differential change of resistance occurs on the two resistors RS1 and RS2 (see Fig 2) The resistance change of the piezoresistors depends on the displacement u of the tip of the cantilever beam and is given by [8] −πl Klcan ΔR = R Il L− Ls (1) where L is the length of the cantilever, Ls is the length of the piezoresistors, z is the distance from the resistor to the neutral plane of the cantilever, πl is the longitudinal piezoresistive coefficient of the resistors (in this paper, we assume the values of room-temperature first-order piezoresistive coeffi3 T is the lateral cients reported in [13]), and Il = (1/12)Wcan CHU DUC et al.: MICROGRIPPER WITH LARGE JAW DISPLACEMENT AND INTEGRATED FORCE SENSORS momentum of inertia of the cantilever Klcan is the lateral stiffness of the sensing cantilever given by [8], [27] Klcan = T ESi Wcan 4L (2) where ESi is the Young’s modulus of the silicon crystal, Wcan is the width of the cantilever, and T is the thickness of the cantilever The resistance change is estimated to be 12% when the tip of the sensing cantilever is bent 9.3 μm corresponding to a 25-μm displacement of the microgripper jaws (see Fig 4) The resistance of the piezoresistor also varies with the temperature The length of the piezoresistors is 68 μm (see Fig and Table I) Considering the simulated temperature distributions in the sensing cantilever (see Fig 3), the temperature in the sensing piezoresistors is changed from ambient temperature at the anchor to 60 ◦ C at the tip of the resistors Therefore, the temperature is, on average, changed by 20 ◦ C over the entire sensing piezoresistors when the microgripper jaw displacement is 25 μm The resistance change of the piezoresistor depends on the temperature change ΔTres , and it is given by ΔRT = αSi ΔTres R0 (3) where αSi = 1.3 × 10−3 is the temperature coefficient of resistance (TCR) of the p-type silicon [14] The resistance will change by 2.6% when the average temperature change in the sensing piezoresistors is 20 ◦ C (see Fig 3) The Wheatstone bridge reduces the temperature influence on the output voltage from a first- to second-order effect, because both sensing resistors on a beam undergo the same temperature shift The two additional resistors outside the sensing cantilever are not subjected to stress They form a matched reference pair that makes the sensor signal more insensitive to common-mode external error sources, such as variations of the environmental temperature (see Fig 2) Assuming that, when the actuator is activated, the resistance values of the sensing resistors RS1 and RS2 are R0 + ΔRT + ΔR and R0 + ΔRT − ΔR, respectively, the output voltage of the Wheatstone bridge is given by [8] Vout = 2VCC R0 ΔR (2R0 + ΔRT )2 − ΔR2 ΔR VCC R0 (4) where VCC is the bias voltage The output voltage is expected to change by 1.7 mV when the displacement of the sensing microgripper jaws is μm Combining (4) and the simulated lateral stiffness Kl , the sensitivity of this sensor is estimated to be 1.9 V/N For the large microgripper displacement of 25 μm and resulting average temperature change of 20 ◦ C in the sensing piezoresistors, the approximation of (4) is valid within 5.7% Another second-order effect that should be considered is the temperature sensitivity of the piezoresistive coefficient, which, according to (1), directly influences ΔR Our piezoresistive coefficient is dominated by the material coefficient π44 of p-type silicon In the range of 25 ◦ C–140 ◦ C, it has a temperature coefficient of 300–500 ppm/◦ C [28] For a temperature rise 1549 of 20 ◦ C on average, this yields a change in the output voltage of 0.8% This is, in most cases, negligible The thermal and 1/f noises are two dominant noise sources of the piezoresistive cantilever [8], [29], [30] The noise voltage of the Wheatstone bridge over the bandwidth of interest (fmin , fmax ) is given by [8] Vn = 4kB Tres R(fmax − fmin ) + αVB2 fmax ln ci Ls Ws Ts fmin 1/2 (5) where VB is the voltage across a resistor with a total number of carriers N , α is a dimensionless parameter that is between 3.2 × 10−6 and 5.7 × 10−6 in single crystal silicon [30], ci is the charge carrier concentration, Tres is the temperature in the resistors, and Ls , Ws , and Ts are the resistor length, width, and thickness, respectively (see Table I) The minimum detectable displacement (MDD) and minimum detectable force (MDF) of the force sensor depend on the minimum detectable signal which is determined by the noise of the cantilever The MDD and MDF corresponding to the calculated noise of the piezoresistors can be estimated by MDD = ujaw Vout /Vn MDF = Fjaw Vout /Vn (6) where ujaw is the sensing microgripper jaw displacement and Fjaw is the lateral force applied to the jaws of the sensing microgripper III F ABRICATION The realized sensing microgripper is shown in Fig The device is 490 μm long, 350 μm wide, 30 μm thick, and with a 40-μm gap between the two jaws The piezoresistive forcesensing cantilever is 390 μm long and 10 μm wide with four piezoresistors on the surface [see Fig 5(b)] Other parameters related to the geometry can be found in Table I The fabrication process (see Fig 6) is based on the Delft Institute of Microsystems and Nanoelectronics (DIMES) bipolar process [26], [31] and the silicon-polymer actuator process [7], [32] SOI wafers with 527-μm-thick silicon (p-type, 100 orientation), 400-nm-thick silicon buried dioxide layer, 30-μmthick single-crystal silicon layer (p-type, 100 orientation), and 1-μm-thick n-type epitaxial layer, with a resistivity of 0.5 Ω · cm, are used An additional 500-nm-thick p-type epitaxial layer with a resistivity of 3.75 × 10−2 Ω · cm is grown to form the piezoresistors By using epitaxial growth, a uniformly doped layer with an accurate thickness within 2%–3% of nominal value can be obtained, resulting in resistors of well-defined sizes The piezoresistors are defined using reactive ion etching (RIE) of silicon as shown in Fig 6(b) A 300-nm-thick low-pressure chemical-vapor-deposited silicon nitride layer is deposited as an electrical insulation layer on the front side On the back side, it serves as the masking layer during etching in KOH solution Then, on the wafer front side, contact windows are opened, and a 600-nm-thick aluminum layer is deposited The piezoresistor connections and electrothermal heaters are defined by using RIE [see Fig 6(c)] 1550 JOURNAL OF MICROELECTROMECHANICAL SYSTEMS, VOL 17, NO 6, DECEMBER 2008 Fig SEM pictures of (a) sensing microgripper and close-ups of (b) piezoresistors, (c) jaws, and (d) section of the thermal actuator IV M EASUREMENT S ETUPS Fig Schematic view of the sensing microgripper fabrication process The top silicon layer is subsequently etched by deep RIE to define the silicon frame until the buried oxide layer is reached [Fig 6(d)] Negative photosensitive SU8 2002 polymer is applied and patterned [see Fig 6(e)] A special prebake and postbake procedure is followed to ensure the void-free filling of the high aspect ratio structures More details can be found in [7] Finally, the bulk silicon is etched from the back side in a 33-wt% KOH solution at 85 ◦ C until the buried silicon dioxide layer is reached The front side of the wafer is protected during the etching in KOH by a vacuum holder The last step is the release of the structure by dry etching the buried silicon dioxide layer from the back side [see Fig 6(f)] For the electrical characterization of the microgripper, dc voltages are applied by using an HP4155A semiconductor parameter analyzer (Agilent Technologies, Inc.) The voltage is ramped from to 4.5 V The displacement is monitored by the charge-coupled device camera on the top of the probe station The static displacement of the microgripper at any actuating voltage is then obtained by enlarging the picture and comparing it with the initial picture External mechanical vibrations cause a blur on the static picture which determines the accuracy of the measurement This inaccuracy is about ±1.5 μm At the same time, a bias voltage VCC with an amplitude of V is applied to the Wheatstone bridge The Wheatstone bridge output is also monitored by the semiconductor parameter analyzer The thermal behavior of the microgripper is investigated by using a Cascade probe station with a heated wafer chuck (Cascade Microtech, Inc.) The investigated temperature range is from 20 ◦ C to 200 ◦ C (the highest temperature of this measurement system) with 10 ◦ C stages and an accuracy of ±0.1 ◦ C In order to get a stable temperature on the device, the measurement is performed after the chuck temperature has reached the setting point to allow sufficient stabilization This externally supplied thermal energy causes expansion in the constrained polymer layer and the resulting actuation A DSP lock-in amplifier SR850 (Stanford Research Systems, Inc.) is used to characterize the frequency behavior of this sensing microgripper A sine signal with amplitude of VPP = V, offset of V, and frequency in the range from 0.1 to 500 Hz is applied to the actuator The corresponding output signal of the Wheatstone bridge is recorded CHU DUC et al.: MICROGRIPPER WITH LARGE JAW DISPLACEMENT AND INTEGRATED FORCE SENSORS 1551 Fig Device operation: (a) Initial position of the sensing microgripper jaws, (b) when 4.5 V is applied to both arms Fig Sensing microgripper jaw displacement versus power consumption Fig Simulated and measured sensing microgripper jaw displacement versus applied voltage The maximum measured displacement is 32 µm at 4.5 V V M EASUREMENT R ESULTS A Electrothermal Actuator Characteristics Fig shows the images of several typical positions of the microgripper jaws In Fig 7(a), the initial position is where the gap between the two jaws which is 40 μm can be seen The distance between the two jaws is close to μm when applying a voltage of 4.5 V to both arms [see Fig 7(b)] Fig shows the displacement response of the microgripper jaws in air when a dc voltage is applied to the electrothermal actuator This measured movement is the total change between the two microgripper jaw positions when both arms are activated The measured results are within 7.5% of the simulated value for all data points A maximum movement of 32 μm is measured at an applied voltage of 4.5 V Therefore, this presented microgripper is capable of manipulating a microobject with a diameter from to 40 μm The power consumption is calculated by the applied voltage and the corresponding current on the electrothermal microactuators Fig shows the measured with linear fitted and simulated values of the jaw displacement versus power consumption On average, the device needs around mW for a 1-μm displacement of the microgripper jaws Fig 10 Sensing microgripper jaw displacement versus average working temperature The average increasing temperature in the electrothermal actuator ΔTave can be estimated by monitoring the change of the resistance of the aluminum heater It is given by ΔTave = Ract (ΔTave ) − Ract (ΔT0 ) Ract (T0 ) αAl (7) where αAl is the TCR of aluminum film (see Table II), Ract (T0 ) is the resistance of the electrothermal actuator (205 Ω at room temperature of −20◦ C), and Ract (ΔTres ) is the resistance of the actuator when the average temperature on the actuator is changed by ΔTave degrees The maximum resistance change is 72% at the applied voltage of 4.5 V, resulting in a maximum average temperature change of 176 ◦ C Fig 10 shows the jaw displacement versus the average working temperature The experimental values come within 7% of the simulated ones The results of the thermal characterization are also shown in Fig 10 The values obtained with the external heat mode come within 7% and 5% of the electrical and simulated ones, respectively It indicates that the aluminum depositing process behaves as expected, and the average working temperature of 1552 JOURNAL OF MICROELECTROMECHANICAL SYSTEMS, VOL 17, NO 6, DECEMBER 2008 Fig 11 also shows the output voltage of the piezoresistive force-sensing cantilever when the microgripper grips a 23-μmdiameter object with the inset clamped object image The sensing microgripper jaws close gradually until it grips the object The contact force between the microgripper jaws and the clamped object can be estimated by the jaw displacement in Fig 11, considering the simulated gripper arm stiffness of 1.8 kN/m (see Table III) The contact force between gripper jaws and object at the applied voltage V is then calculated as FContact = Kl ∗ (d (V ) − d(V )) Fig 11 Output voltage of the force-sensing cantilever versus the applied voltage on the electrothermal microactuator The inset shows the microgripper jaws with the clamped object the actuator can be well estimated from the resistance change of the aluminum heater However, the physical properties of a polymer material such as the volume coefficient of expansion, Young’s modulus, and so on are greatly changed in pseudosecond order at the glass transition temperature Tg where the material properties change from the glassy region to the rubbery plateau region [33] The glass transition temperature of a polymer varies widely with parameters such as the fabrication process and the microscopic structure [17], [33], [34] The Tg of SU8 is nearly the baking temperature when it is below 220 ◦ C for a baking time of 20 [17] However, the Tg can increase gradually up to the steadystate temperature of 118 ◦ C when the material is baked for a longer time (60 min) at a constant temperature of 95 ◦ C The effect of the glass transition temperature is apparent in the measurements of Fig 10, where two different working ranges can be distinguished The data points lay along straight fitting lines, which intersect each other just above 120 ◦ C This is fairly close to the steady-state SU8 glass transition temperature of 118 ◦ C reported in [17] It indicates that the proposed postbake process of this device is sufficient in this context Furthermore, it explains the nonlinear characteristic of the displacements due to the power consumption and also the working temperature (see Figs and 10) B Force-Sensing Cantilever Beam Characteristics Fig 11 shows the measured output signal of the Wheatstone bridge versus the voltage applied on the electrothermal microactuator The zero-stress resistance value of the piezoresistors at room temperature is 39 kΩ The bias voltage is V dc The maximum output voltage of the sensor bridge is 49 mV when the voltage applied to the actuator is 4.5 V The relation between the output voltage and the sensing microgripper jaw displacement is also shown in Fig 11 The sensitivity of the sensing microgripper derived from this curve is 1.5 kV/m This curve is linear within 2% The experimental results come within 10% of the calculated ones obtained from (4), indicating that the epitaxial growth, etching process, and resistor contacts behave as expected (8) where Kl is the lateral stiffness of the sensing microgripper arm, d (V ) is the displacement of microgripper jaws at applied voltage V without the clamped object in between the two jaws (dashed line in Fig 11), and d(V ) is the displacement of microgripper jaws at applied voltage V with the clamped object in between the two jaws (solid line in Fig 11) Fig 12 shows the calculated contact force of this proposed microgripper The contact force is zero until the two gripper jaws reach the object at an applied voltage of about 3.75 V The contact force then increases up to 135 mN at the applied voltage of 4.5 V Combining the measured results in Fig 11 and the calculated force, the sensitivity of this built-in force sensing is estimated to be 1.7 V/N This sensing microgripper is capable of detecting the diameter of the clamped object and also the contact force between the microgripper jaws and the object This function is highly desirable for the closed-loop system needed in microassembly, microrobotics, minimally invasive surgery, and living cell surgery C Response Frequency of the Sensing Microgripper Fig 13 shows the measured voltage gain and phase shift as a function of frequency of this sensing microgripper using the lock-in amplifier The large-signal cutoff frequency of this sensing microgripper is measured as 29 Hz The transient response of the full range displacement of this sensing microgripper is also characterized The rise and settling times are measured to be 13 and 18 ms, respectively Combining (6) and the output signal from Fig 11 with the frequency bandwidth of the range 0.1–29 Hz, the MDD and the corresponding MDF of the sensing cantilever beam can be estimated to be about nm and 770 nN, respectively D Reliability The main failure mechanism observed during the test of the microgripper is the appearance of cracks in the aluminum heater and the silicon comb structure when the applied voltage is increased to about V and the working temperature of the actuator is too high There is no indication of the loss of adhesion between the SU8 and the silicon plates even at these temperatures To investigate the lifetime of the microgripper, it is repeatedly actuated in air with a 4-V amplitude (90% of its maximum displacement) and with a time period of s/sweep for 24 h (14 400 cycles) The same reliability testing process is CHU DUC et al.: MICROGRIPPER WITH LARGE JAW DISPLACEMENT AND INTEGRATED FORCE SENSORS 1553 TABLE III PERFORMANCE OF THE SENSING MICROGRIPPER Fig 12 Contact force between microgripper jaws and the objects versus the applied voltage Fig 14 Microgripper is bonded on the modified dual in-line package In this way, both mechanical manipulation and electrical connection of the sensing microgripper are possible modified socket chip (see Fig 14) Then, the chip is mounted on an xyz micromanipulator that allows moving the microgripper in three dimensions As testing objects, ∼30-μm-diameter glass balls placed on a silicon wafer surface are used The glass balls are rearranged to form the letter “L” as shown in Fig 15 The microgripper tip is moved to approach the ball [see Fig 15(a)] The microgripper closes to grasp the object The chip is then moved to the target position using the xyz manipulator [see Fig 15(b) and (c)] The microgripper finally opens to release the ball [see Fig 15(d)] When releasing the glass ball, we sometimes observe stiction between the microgripper jaw and the object However, we can get rid of this adhesion force by applying a small force between the glass ball and the silicon wafer surface before releasing the object Fig 13 Bode diagram of the sensing microgripper The sweep input voltage is applied to electrothermal actuator, and the output of the piezoresistive Wheatstone bridge is monitored The cutoff frequency is 29 Hz repeated after one week and then one month No degradation in performance is noticed E Object Manipulation The microparticle manipulating ability of this microgripper developed is investigated The microgripper is bonded on the VI C ONCLUSION A novel design of a sensing microgripper based on silicon-polymer electrothermal actuators and piezoresistive force-sensing cantilever beams is presented The sensing microgripper is 490 μm long, 350 μm wide, and 30 μm thick A microgripper jaw displacement up to 32 μm at an applied voltage of 4.5 V is measured The microgripper can be used to grasp an object with a diameter of 8–40 μm The maximum average 1554 JOURNAL OF MICROELECTROMECHANICAL SYSTEMS, VOL 17, NO 6, DECEMBER 2008 R EFERENCES Fig 15 Manipulating micro-glass balls to form letter L: (a) Initial position (b) Microgripper closes to grasp the ball (c) Microgripper is moved to the right position (d) Microgripper opens to release the ball working temperature change is 176 ◦ C at 4.5 V The output voltage of the piezoresistive sensing cantilever is up to 49 mV when the jaw displacement is 32 μm The force sensitivity is measured to be up to 1.7 nN/m, and the corresponding displacement sensitivity is 1.5 kV/m The bandwidth frequency of this presented sensing microgripper is measured as 29 Hz The MDD and MDF are estimated to be nm and 770 nN, respectively The fabrication process is based on conventional bulk micromachining and polymer filling, and it is CMOS compatible The characteristics of this sensing microgripper will make the manipulation of small objects more efficient, more accurate, and less tiring than with currently available grippers due to its large jaw displacement and sensing sensitivity The presented sensing microgripper could be used in automatic systems for microassembly and in microrobotics In addition, the microgripper could be of use in living cell handling or in minimally invasive surgery, provided the working temperature is lowered and the electronics are properly isolated from the liquid environment ACKNOWLEDGMENT The authors would like to thank the DIMES-IC Processing group for the technical support, P J F Swart of the Electronic Components, Technology and Materials group for the help, G de Graaff of the Electronics Instrumentation Laboratory for the help with the electronic and mechanical measurements, and J Wei, M Saadaoui, and H W van Zeijl of the Electronic Components, Technology and Materials group and S L Paalvast and W J Venstra of the Precision and Microsystems Engineering Department for their suggestions and 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Sperling, Introduction to Physical Polymer Science Hoboken, NJ: Wiley, 2006 [34] J H van Zanten, W E Wallace, and W Wu, “Effect of strongly favorable substrate interactions on the thermal properties of ultrathin polymer films,” Phys Rev E, Stat Phys Plasmas Fluids Relat Interdiscip Top., vol 53, no 3, pp R2053–R2056, Mar 1996 Trinh Chu Duc received the B.S degree in physics from Hanoi University of Science, Hanoi, Vietnam, in 1998, the M.Sc degree in electrical engineering from Vietnam National University, Hanoi, in 2002, and the Ph.D degree from Delft University of Technology, Delft, The Netherlands, in 2007 His doctoral research concerned piezoresistive sensors, polymeric actuators, sensing microgrippers for microparticle handling, and microsystems technology He is currently an Assistant Professor with the Faculty of Electronics and Telecommunication, College of Technology, Vietnam National University Gih-Keong Lau received the B.Eng (with first-class honors) and M.Eng (by research) degrees in mechanical engineering from Nanyang Technological University (NTU), Singapore, in 1998 and 2001, respectively, and the Ph.D degree from Delft University of Technology, Delft, The Netherlands, in 2007, where his research topics were polymer microactuators and microfabrication From 2001 to 2003, he was a Research Associate with the Centre for Mechanics of Microsystems, NTU, where he worked on the topology optimization of compliant mechanisms and piezoelectric actuators for hard disk drives and, since 2008, has been an Assistant Professor with the School of Mechanical and Aerospace Engineering His current research interests are electroactive polymer actuators and their microfabrication 1555 J Fredrik Creemer (S’97–A’01–M’03) received the M.Sc degree in electrical engineering from Delft University of Technology, Delft, The Netherlands, in 1995, the Diplôme d’Études Approfondis in electronics from the Université Paris-Sud, Orsay, France, in 1996, and the Ph.D degree (cum laude) from Delft University of Technology, in 2002 His doctoral research explored the effect of mechanical stress on bipolar transistor characteristics He was an Analog Chip Designer, with SystematIC Design from 2002 to 2003 In 2003, he was with the Kavli Institute of Nanoscience, as a Postdoctoral Researcher In 2006, he was an Assistant Professor with the Laboratory for Electronic Components, Technology and Materials, Delft Institute of Microsystems and Nanoelectronics, Delft University of Technology His research interests are microelectromechanical system microreactors, transmission electron microscopy, and microsystems technology Dr Creemer was the recipient of the Else Kooi Award 2002 for the research described in his dissertation and, in 2006, a Veni Grant Pasqualina M Sarro (M’84–SM’7–F’07) received the Laurea degree (cum laude) in solid-states physics from the University of Naples, Naples, Italy, in 1980, and the Ph.D degree in electrical engineering from Delft University of Technology, Delft, The Netherlands, in 1987, where her thesis dealt with infrared sensors based on integrated silicon thermopiles From 1981 to 1983, she was a Postdoctoral Fellow with the Photovoltaic Research Group, Division of Engineering, Brown University, Providence, RI She then joined the Delft Institute of Microsystems and Nanoelectronics, Delft University of Technology, where she is responsible for research on integrated silicon sensors and microelectromechanical systems (MEMS) technology In December 2001, she became the A van Leeuwenhoek Professor, and, since 2004, has been the Head of the Electronic Components, Materials and Technology Laboratory She has authored or coauthored more than 350 journal and conference papers Dr Sarro was the recipient of the EUROSENSORS Fellow Award in 2004 for her contribution to the field of sensor technology In April 2006, she became a member of the Dutch Royal Academy of Science, and in November 2006, she was elected an IEEE Fellow for her contributions to micromachined sensors, actuators, and microsystems She is a member of the technical program committees for several international conferences (IEEE MEMS, IEEE Sensors, EUROSENSORS, and Transducers), the Technical Program Cochair for the First IEEE Sensors 2002 Conference, and the Technical Program Chair for the Second and Third IEEE Sensors Conference (2003 and 2004) She is the General Cochair of IEEE MEMS 2009 She is also a member of the AdCom of the IEEE Sensors Council  ...CHU DUC et al.: MICROGRIPPER WITH LARGE JAW DISPLACEMENT AND INTEGRATED FORCE SENSORS 1547 TABLE I GEOMETRY OF THE SENSING MICROGRIPPER Fig Schematic drawing of the sensing microgripper [10]–[12]... room-temperature first-order piezoresistive coeffi3 T is the lateral cients reported in [13]), and Il = (1/12)Wcan CHU DUC et al.: MICROGRIPPER WITH LARGE JAW DISPLACEMENT AND INTEGRATED FORCE SENSORS. .. can be estimated by MDD = ujaw Vout /Vn MDF = Fjaw Vout /Vn (6) where ujaw is the sensing microgripper jaw displacement and Fjaw is the lateral force applied to the jaws of the sensing microgripper