Out-of-plane electrothermal actuators insilicon-on-insulator technology

Des declencheurs electrothermiques hors planavec une technologie silicium-sur-isolateur

See-Ho Tsang, Dan Sameoto, and M. Parameswaran∗

A novel electrothermal micro-actuator capable of actuation after being rotated 90 out of plane is presented. The actuator is fabricated and characterizedusing the Micragem silicon-on-insulator (SOI) process developed by Micralyne Inc. The actuator is designed to be rotated out of plane by means of aserpentine spring design that provides electrical contact to the substrate while applying a downward force to hold the actuator in place. The eventualpurpose of this actuator is to adjust the position of a micromirror at a 90 angle to the substrate. The actuator performs as expected in the horizontalposition and can be assembled using standard microprobes. The actuator performance and the theory of spring operation are presented using an analyticalheat transfer model, finite element modelling, and device measurements.

On presente un nouveau micro-declencheur electrothermique capable de declenchement apres avoir ete tourne 90 hors plan. Le declencheur est fabriqueet caracterise en utilisant le processus silicium-sur-isolateur (SOI) Micragem developpe par Micralyne Inc. Le declencheur est concu pour etre tourne horsplan en utilisant une conception a ressort en serpentin qui fournit un contact electrique au substrat pendant qu’on applique une force de haut en bas pourgarder le declencheur en place. Le but de ce declencheur est d’ajuster la position d’un micro-miroir a 90 au substrat. Le declencheur travaille comme prevudans la position horizontale et peut etre assemble en utilisant les micro sondes standards. La performance du declencheur et le principe de fonctionnementdu ressort sont presentes en utilisant un modele de transfert de chaleur analytique, une modelisation par elements finis et des mesures de dispositifs.

I. Introduction

Out-of-plane actuation is desired for microelectromechanical sys-tems (MEMS) devices such as grippers, scanners, mirrors, and shut-ters. When assembled out of plane, micro-actuators can adjust andmanipulate objects or optical systems that are positioned vertically rel-ative to the substrate. However, these components must be fabricatedin plane and assembled afterwards to achieve out-of-plane orientation.Typically, out-of-plane actuators are designed with hinges [1]–[6] toallow the devices to rotate out of plane. However, the creation of hingesis limited to multilayer processes, and designs using hinged structurescannot be employed in single-layer processes. Furthermore, hinges aresusceptible to mechanical jamming, which reduces the yield of assem-bled structures. In order for devices to be rotated out of plane in single-layer processes, structures require torsion beams or flexures [7]–[8].Torsion beams are generally a poor option because of the lengthsrequired for compliance, as well as significant variability in failurestrength due to torsional stress.

An improved flexure design for out-of-plane rotation based on ser-pentine springs has been designed, simulated, and tested. The serpen-tine spring design requires smaller chip area to fabricate, is stressedprimarily in bending, and serves to hold the vertical structure in placeonce assembled. The rotation out of plane in this design is achieved bypushing the actuator with a micromanipulator probe tip. The serpentinespring suspension system translates the forward motion into a rotationout of plane after a deflection of several tens of micrometres in plane.After designing and modelling the device, we fabricated and tested theout-of-plane actuator using the Micragem silicon-on-insulator (SOI)process [9].

∗See-Ho Tsang, Dan Sameoto, and M. Parameswaran are with the Instituteof Micromachine and Microfabrication Research, School of Engineering Sci-ence, Simon Fraser University, 8888 University Drive, Burnaby, B.C. V5A 1S6.E-mail: stsanga, dsameoto, param@sfu.ca

In our design, the actuator that is rotated out of plane is a chevron-style electrothermal actuator [10]. The chevron design was chosen forits simplicity and reliability. Electric current that is passed through theactuator is dissipated in thin beams that heat and expand. The ther-mal expansion of the beams is amplified by geometric constraints andmechanical amplifiers.

The Micragem technology is appropriate for suspended thermal ac-tuators because the current path can be selectively controlled in a singlelayer of single-crystal silicon (SCS). The resistivity of the SCS in theMicragem process is very high (2.756Ω-cm) [9]. A standard thermalactuator made using SCS with the same resistivity would require volt-ages greater than 20 V for significant current to flow. To reduce thisthreshold, the metal layer available in the Micragem technology canbe patterned on top of the thermal actuators to provide thermal energythrough Joule heating. Operating the chevron actuator in this mannercan lower the required voltage to less than that of a pure silicon actua-tor, and a monolithic silicon structure can be selectively heated.

II. Design concept and theory

Micragem is an SOI process developed by Micralyne Inc. that uses sur-face and bulk micromachining techniques to produce micromechani-cal and microfluidic devices [9]. An SOI wafer consisting of a 10 µm–thick SCS layer bonded to a 525 µm–thick Pyrex substrate is used. Theprocess begins with the Pyrex wafer, which undergoes two isotropicwet etches to create trenches of 2 µm, 10 µm, or a combined 12 µmdepth. After this is done, a metal layer (METAL1), consisting of 50 nmof titanium, 50 nm of platinum, and 200 nm of gold can be depositedand patterned inside the trenches. Once the substrate is processed, theSCS wafer, held by a handle wafer, is anodically bonded to the Pyrexto create the structural layer. Then the handle wafer is removed, anda second metal layer (METAL2), consisting of 10 nm of chrome and75 nm of gold, is deposited and patterned. Finally, a deep reactive ion

Can. J. Elect. Comput. Eng., Vol. 31, No. 2, Spring 2006

98 CAN. J. ELECT. COMPUT. ENG., VOL. 31, NO. 2, SPRING 2006

Figure 1: Description of the Micragem process. The sequence (a) to (d) illustrates themajor processing steps for creating a sample structure.

etching (DRIE) etch patterns the SCS mask and is designed to producesilicon features with the second metal layer self-aligned on top. TheDRIE step is used to form the thin chrome/gold-covered beams for thechevron actuators. Fig. 1 illustrates the basic processing steps of theMicragem process.

A. Design componentsThe basic dimensions and the design of the out-of-plane electrothermalactuator are shown in Fig. 2 and Fig. 3, respectively. The structure ofthe device consists of three functional components: a serpentine springsystem that electrically and mechanically connects the actuator to thesubstrate; a two-bar chevron electrothermal actuator; and mechanicalamplifiers that convert the small displacement of the chevron actuatorinto significant deflection.

B. Serpentine spring designThe serpentine springs on either side of the actuator are designed toprovide rigidity in plane while allowing the structure to rotate out ofplane when the device is pushed forward using a micromanipulatorprobe tip. Previous research has indicated that the in-plane stiffness ofMEMS flexures is highly dependent on spring shape, while the out-of-plane stiffness is not [11]. The serpentine springs are designed tobe deflected in plane until a toggle point is reached, after which thedevice rotates out of plane. The out-of-plane moment is initially verysmall, but if a small downward displacement is included at the bottomlip of the actuator, then it can create an out-of-plane moment that willcontribute to an earlier rotation. Once the actuator has been rotated outof plane, the serpentine springs produce a large downward force be-tween the lower lip of the actuator and the substrate. The downwardforce provides a friction contact with the substrate that can prevent thestructure from collapsing once assembled. A model of the assembleddevice is shown in Fig. 4. In the Micragem fabricated devices, eachserpentine spring is 2.25 loops in length at a 1:1 (height-to-width) as-pect ratio. The number of loops and the aspect ratio were selected tominimize the stress in the springs during and after assembly.

The bending of the spring system is very nonlinear, making an an-alytical solution for the mechanical deflection impractical. However,finite element modelling (FEM) can be effectively applied to solve forthe spring deflection and stress distribution. The description of materialproperties for SCS supplied by Micralyne lists an empirical measure-ment for the elastic modulus of 129 GPa and a typical Poisson ratio of0.26 [9]. These values were used in our FEM simulations along with afixed displacement input to the bottom edge of the actuator to simulateassembly. The resulting model was generated as shown in Fig. 4.

As determined by the finite element solution, a maximum stress of2.58 GPa occurs on the first loop of the spring assembly adjacent tothe anchor. Because the yield strength for SCS is in the order of 7 GPa,a safety factor of approximately 2.7 is achieved. From the nodal so-lution of the model, a downward force of 2.82 mN and a lateral forceof 0.933 mN are generated by the serpentine springs after assembly.These forces must be matched by equal forces on the bottom edge of

Figure 2: Dimensions of the out-of-plane thermal actuator, shown in micrometres. Thethickness of the layer is 10 µm.

Figure 3: Diagram illustrating the basic design and components of the out-of-plane ther-mal actuator. The gray areas represent the metal layer used for electrical conduction.

the actuator for the structure to be in equilibrium. Because the gravita-tional force on the microstructure is negligible, the friction force (Ff )between the actuator and substrate is governed by Ff = µFN , whereµ is the coefficient of static friction and FN is the normal force. Solv-ing for the required coefficient to keep the actuator in place, we findthat µ must be equal to or greater than 0.33. However, because thecoefficient of friction must be measured empirically for the device, re-lying only on the friction to keep the actuator in place is impracticable.Therefore in order to increase the lateral stability, ridges were addedto the substrate along the area of the expected contact to act as a me-chanical stop. The height of each ridge was designed to be 2 µm, andit was anticipated that the mechanical stop would hold the actuator inplace after assembly.

Choosing an optimal design for the serpentine spring requires aniterative process through several variations of spring length, numberof loops, and aspect ratios. The fabrication schedule limited the de-signs that could be tested using the Micragem process. The majorityof designs fabricated varied the number of loops and the length of thesprings. The 2.25-loop design with 325 µm–long springs performedthe best in experimental trials. After the Micragem fabrication dead-line, another set of structures was fabricated using a single layer ofpolysilicon in the PolyMUMPs process [11] to further test the feasi-bility of the serpentine spring design. The spring designs fabricatedusing PolyMUMPs were 130 µm in length with 2.25 loops and a 1:2(height-to-width) aspect ratio. The structures fabricated using this pro-cess performed as anticipated and were easily assembled with a singlepush from a micro-actuator probe tip. Fig. 5 shows an assembled struc-ture using the serpentine spring system.

TSANG / SAMEOTO / PARAMESWARAN: OUT-OF-PLANE ELECTROTHERMAL ACTUATORS 99

Figure 4: Finite element model of the device after assembly, showing the stress concentra-tions of the compliant spring system. The arrow highlights the location of maximum stresswith a value of 2.58 GPa.

Figure 5: An SEM image of a PolyMUMPs fabricated structure illustrating the concept ofthe serpentine spring assembly system.

C. Chevron electrothermal actuator designChevron-style electrothermal actuators with mechanical amplifierswere selected instead of other electrothermal actuator designs becauseof their simplicity and reliability of actuation. The motion that theseactuators produce is linear, and the force is very high, allowing themechanical amplifiers to produce large deflections. The chevron elec-trothermal actuators produced in the Micragem process are actuatedby applying a current across the beams. The conduction path for heat-ing the beams is defined by selectively patterning a chrome/gold metallayer on the surface of the device. The chevron electrothermal actua-tors are powered by heating the metal layer on top of the thin bars andrelying on thermal conduction to heat the silicon below. This methodis very similar to that used for other electrothermal actuators fabricatedwith an electrically insulating material [12]–[13]. The electrical resis-tivities of the silicon and gold as provided by Micralyne are 2.756Ω-cm and 2.2×10−6 Ω-cm respectively [9]. In the Micragem process thegold layer is 75 nm thick, and the silicon layer is 10 µm thick. There-fore, the thickness of the silicon is approximately 133 times that of thegold. Solving for the sheet resistance, we find that the theoretical sili-con resistance is approximately 10 000 times that of the gold. Because

Table 1Thermal model parameters

Thermal conductivity kHeat capacitance CDensity of silicon ρEffective perimeter of actuator section Pi

Cross-sectional area of actuator section Ai

Length of actuator section Li

Voltage across actuator section Vi

Temperature of actuator section Ti

Ambient temperature T∞Distance from end of actuator section xDistance between actuator and substrate ∆z

of this very large difference in electrical conductivity, the silicon canbe neglected when we consider the power generation in the thermal ac-tuator model. Furthermore, the electrical resistivity of the 10 nm–thickchrome is 1.29 × 10−5 Ω-cm, which is equivalent to adding another2 nm of gold in a lumped model. The analytical thermal model usedfor this actuator is based on the one-dimensional thermal model pre-sented in [10]. This model assumes that heat transfer occurs primarilythrough conduction to the substrate through the air and along the sili-con beams. Reference [10] compared the heat transfer rates from con-duction, convection, and radiation at the micro scale and found that fortypical temperatures and device sizes, conduction through the air andsilicon account for more than 99% of total heat transfer. We assumethat the power dissipated in the gold is lost by conduction to both thesilicon and the air, whereas the vast majority of the heat is lost to thesilicon because of its much higher coefficient of thermal conductivity.In the original model of [10], the heat loss is transferred into the sub-strate, which acts as an infinite heat sink. The heat conduction to thesubstrate, in our model, is either through the air between the actuatorand substrate or along the silicon arms. The rate of temperature changein a cross-section of a thermal actuator can be modelled by (1) with theterms as defined in Table 1:

∂T1(x)

∂t=

kSi

CSiρ

∂2Ti

∂x2− kairPi

CSiρAi∆z(Ti − T∞)

+

V 2i

Ri

CSiρAiLi. (1)

Numerical integration is used to determine the steady-state temper-ature profile of the actuator with a given voltage applied. The volt-age and resistance across each section are determined by the effectiveresistivity, dimensions, and total voltage applied to the actuator. Thederived temperature in the model in [10] is relative to the heat sinksassumed to be at either end of the actuator. In a suspended actuatormodel, where infinite heat sinks do not necessarily exist at either end,the one-dimensional model still gives a reasonable temperature profilerelative to that of the whole system because the displacement of theactuator is strongly dependent on the temperature difference betweenthe chevron and the surrounding silicon. Once the actuator is rotatedout of plane, a simple one-dimensional model is less appropriate, andan FEM analysis is necessary to determine the actuator behaviour. Thesteady-state temperature profile of the actuator was calculated usingthe same method as in [10]. Conduction through the silicon beams tothe substrate is dependent on the cross-sectional area, but conductionthrough the air to the substrate is dependent on the effective perimeterof the beam. In the original model from [10], this effective perimeterwas approximately 50% of the actual perimeter. For the dimensionsand material properties of our actuator, the effective perimeter wasfound to be a slightly smaller percentage of the real perimeter becauseof the different aspect ratio of the structure [14].

100 CAN. J. ELECT. COMPUT. ENG., VOL. 31, NO. 2, SPRING 2006

Figure 6: Graph of the increase in temperature as distributed across a single-beam elec-trothermal actuator.

The effective resistivity of the combined gold/silicon beam wasfound by using the resistance of the gold layer and the dimensionsof the silicon to extrapolate an appropriate value equivalent to higher-conductivity silicon. The modelling of the effective resistivity is justi-fied by the insignificant temperature gradient across the cross-sectionof the actuator. The temperature profile with 1 V applied across the ac-tuator is shown in Fig. 6. The relative temperature is measured as anincrease over the ambient temperature of the surrounding air or siliconframe holding the chevron beam. The absolute temperature reached bythe beam is expected to be significantly higher, since the power inputinto the beam will also increase the ambient temperature.

In addition to the one-dimensional analysis, a more detailed modelwas employed for FEM simulations. A given power dissipation wasconverted to a heat flux on the surface of the beams of the actuator. Asdetermined by our calculations, the heat flux of 3.35 × 107 W/m2 isequivalent to approximately 60 mW of Joule heating on the beams. Us-ing the calculated heat flux, a thermal expansion of 2.5×10−6 K−1 andthermal conductivity of 141 W/m·K for silicon, and a thermal conduc-tivity of 0.03 W/m·K for air, the temperature and displacement weresolved by FEM analysis and are shown in Fig. 7. A maximum temper-ature of 661.1C was reached when the aforementioned heat flux wasapplied, giving a total tip separation, as measured from the centres,of 27.28 µm after the motion was transferred through the mechani-cal amplifiers. A simulated 60.3 mW of power is dissipated across thebeams. The maximum temperature reached, with a value of 661.1C,is located on the centre shuttle, and the surrounding volume above theactuator is modelled as air at 20C, while the volume underneath theactuator, separated by a 10 µm air gap, is modelled as an infinite heatsink at 20C.

III. Implementation and assembly

A. Quality of fabricated devicesThe design as shown in Fig. 2 was implemented using the Micragemprocess, and the devices were returned to us by the Canadian Micro-electronics Corporation. Before experiments were performed, a scan-ning electron microscope (SEM) was used to examine the overall qual-ity of the structures shown in Fig. 8. Based on the SEM images, weconcluded that the SCS structures were produced within the alignmenttolerances expected for this process. Under higher magnification, theridges etched into the substrate to increase frictional contact can beseen. However, we suspect that a minor over-etch occurred, reducingthe definition of the ridges and consequently their effectiveness as amechanical stop. The metal layer on top of the SCS was also poorlydefined and, in some cases, completely missing. The damaged metallayer significantly affects the yield and performance of the chevron ac-tuators because the missing metal causes breaks in the electrical path.

Figure 7: Finite element model of the temperature distribution on the chevron electrother-mal actuators prior to assembly.

The metal layers should self-align to the silicon edge. However, theprocess variation for the etch step caused the metal coverage to recedeand disappear in many cases, as shown in Fig. 8(c).

In addition to the defects discussed above, we discovered anotherimportant flaw under high magnification; namely, the poor sidewalldefinition of the SCS structures. As shown in Fig. 8(d), a ridge onthe sidewalls was created during the DRIE step. Because of the un-evenness of the sidewalls, a large number of stress concentrations candevelop, possibly weakening the silicon components. The rough side-walls may also produce the effect of many small dimples, reducing thearea of contact between the device and the substrate.

B. Assembly of structuresAll out-of-plane electrothermal actuators were designed to be assem-bled manually. In the future this process may be automated, but theforces and displacements required for assembly are larger than anymicro-actuator can presently produce. The advantage of the serpentinespring design is that assembly is theoretically very simple. FEM sim-ulations indicate that pushing the structure forward in plane with aninitial downward deflection is the only action required to assemble it.Fig. 9 shows a four-part sequence for the assembly of the fabricatedout-of-plane actuator. Fig. 10 shows one actuator that is completelyassembled by the method described.

IV. Experimental results

A. Thermal actuator testsBefore assembly, the thermal actuators were tested to confirm the the-ory of operation. The actuators are suspended above the substrate;therefore the silicon frame surrounding the chevron beams had to con-strain them sufficiently for deflection to occur. If the actuator operatedin plane, it would operate once assembled, since the current could bedelivered along the serpentine springs. All deflection was measured asthe distance between the centres of the two mechanical amplifier tips.This measure was chosen in order to provide the largest distance mea-surement and minimize measurement error. Photomicrographs weretaken of the actuator for different voltage inputs. The current at eachvoltage was also measured in order to calculate the input power cor-responding to each image. The microprobes were placed as close tothe actuator as possible so that measurements could be taken of onlythe chevron actuator apart from any resistance losses in the serpentinesprings. National Instruments’ Vision Builder software was used tocalibrate each image and compare the deflection of the amplifier tips.Data for one sample trial are presented in Fig. 11.

TSANG / SAMEOTO / PARAMESWARAN: OUT-OF-PLANE ELECTROTHERMAL ACTUATORS 101

Figure 8: Scanning electron micrographs of the Micragem fabricated device. (a) Overviewof the actuator design illustrating chrome/gold contacts on the substrate for powering theactuator. (b) A closer view of the chevron design showing the chrome/gold metal linesdeposited on the silicon beams. (c) Close-up view of the centre hub of the chevron actu-ator. Note that the chrome/gold layer does not completely reach the edges. The breaks inthe chrome/gold occur on various devices. (d) A close-up of the actuator tip. Note thatthe DRIE is inconsistent, such that a ridge along the side is created. The ridges for themechanical stop on the substrate can also be seen.

Figure 9: Assembly of the out-of-plane actuator. The sequence (a) to (d) shows two mi-cromanipulator probe tips resting on the structure.

Once the device is assembled and powered, current flows throughboth the actuator and the serpentine springs, increasing the total re-sistance and the total required power. Therefore, direct comparison ofdeflection in the assembled position versus the in-plane position canbe done only by examining the deflection versus input current.

During the experiment, excess power causes the metal layer to de-laminate from the SCS surface, melt, or vapourize. From our obser-vations, the maximum tip separation among the actuators tested wasapproximately 42 µm before failure. The maximum input power waslimited to 100 mW because the inconsistent gold coverage caused sig-nificant failure variation over the course of our trials. The stroke forelectrothermal actuators is expected to be linear with respect to the in-

Figure 10: Photomicrographs of the actuator assembled out of plane. (a) Top edge of thesilicon structure surrounding the chevron actuator as viewed from above. (b) Top of themechanical amplifier as viewed from above.

Figure 11: Photomicrographs of the deflection of the microgripper showing the tip sepa-ration versus electrical power input.

put power. Although there is a slight deviation from the linear trendin the experimental data, the general tip separation for this device fallswithin acceptable performance criteria.

B. Experimental comparisonThe behaviour of the chevron electrothermal actuator varies slightlyfrom that of the theoretical model. When power is applied, the tipsof the mechanical amplifiers separate with a gap that is slightly largerthan expected. Fig. 12 shows the deflection response of the tip whenelectrical power is input. The experimental and FEM results are com-pared in this figure.

In our analysis we initially suspected that the deviation from theFEM model resulted from the method by which the power was mea-sured in each case. For the experimental case, the total voltage andcurrent passing through the device are measured directly. For the FEMmodel, the total power is directly applied to the heating arms, and anyheating in other areas of the device is neglected. If the heating in otherareas of the model is included, the expected tip separation can onlydecrease, showing that the difference in power measurement is not thesource of the discrepancy.

102 CAN. J. ELECT. COMPUT. ENG., VOL. 31, NO. 2, SPRING 2006

Figure 12: Graph illustrating the mechanical-amplifier tip separation from the experi-mental data and as predicted by the FEM analysis.

Figure 13: Photomicrograph of a device that failed during assembly. The arrow indicatesthe fracture at a point that is not expected to be a stress concentration. The breakageoccurred during the final push of the assembly, just as the actuator reached an angle ofapproximately 45 relative to the substrate.

Therefore, we can conclude that the deviation between the experi-ment and the FEM model is most likely explained by the large varia-tions in the metal layer across the chevron beams. The variations acrossthe metal layer change the profile of heat conduction so that the modelused in the FEM no longer holds. For future work, in order to fullyunderstand the thermal transfer from the metal layer to the SCS layer,the irregularities of the metal layer will have to be taken into account.

C. Spring testingAlthough the design was initially intended to require a lateral pushfrom only a single micromanipulator probe tip for assembly, two probetips were necessary in practice to achieve the required path of motionon the Micragem fabricated chips. The primary reason for this wasthe 1:1 aspect ratio of the spring suspension on the fabricated device.This aspect ratio is not sufficiently rigid in plane to produce out-of-plane motion at low displacement. Consequently, there was interfer-ence between the loops of the springs before the necessary rotationwas achieved during assembly. A second probe tip lifted the structureenough to allow clearance in the vertical plane and to avoid interfer-ence between the loops.

Of the 20 devices tested, only three structures could be fully as-sembled because most devices broke before they could be fully ro-tated into the locked position. Serpentine spring fractures occurred atdisplacements far lower than the predicted failure point and in unex-

pected locations. It is our hypothesis that the large number of stressconcentrations or crystal defects due to sidewall roughness negativelyaffected the silicon fracture strength. The stress concentration theory issupported by the significant percentage of breakage at locations of lowstress. An example of such a breakage can be seen in Fig. 13, wherethe initial breaking point is in between the loops near the centre of thebeam. Breakage at this location is unexpected and is most likely dueto defects in the sidewalls or defects in the silicon crystal. We do nothave empirical data on silicon yield strength specific to the Micragemprocess, so this hypothesis cannot be confirmed, but it seems a likelyexplanation for the results observed.

For those devices that survived assembly, even small vibrationswere capable of causing the bottom lip of the device to dislodge fromthe substrate, resulting in the collapse of the assembled device. Theforce of this impact was typically large enough to shatter the mechan-ical amplifiers or serpentine springs. The ridges etched into the sub-strate, which were designed to compensate for the low frictional force,did not supply a large enough divot for the bottom lip of the actuatorto catch. As a result, the stability of assembled devices was very poorand they stayed vertical for only a few minutes.

V. Failure modes and future designs

The critical failure in the fabricated designs can be attributed to thepoor control of the DRIE step of the Micragem process. The DRIEcaused problems with gold coverage and silicon strength. Therefore,the spring strength and thermal actuator performance were highly vari-able because of the processing conditions, causing most of the fabri-cated designs to fail prematurely. Presently the DRIE step is highlyunreliable, but its reliability is expected to improve as the Micragemtechnology matures.

Two problems with the design that can be corrected in the future arerelated to the aspect ratio of the silicon bars and the locking mecha-nism for the actuator. First, a serpentine spring with a different aspectratio (with width larger than thickness) should increase the downwardforce relative to the lateral force and increase device stability. Second,although ridges were etched into the substrate to increase the forceholding the actuator upright, they were insufficient to keep the devicelocked in place. Therefore, a greater pitch will have to be used to en-sure that the isotropic etch defines a higher-quality ridge if this type oflocking mechanism is to be employed again.

If the fabrication issues are resolved and the recommendations forfuture designs are employed, the yield of successfully assembled de-vices is expected to increase significantly.

VI. Conclusion

Using an SOI process, we designed and fabricated the first electrother-mal actuator that can be rotated out of plane. The concept of an elec-trothermal actuator working on a relatively nonconductive structuralmaterial was proved. To the best of our knowledge, this is the firstelectrothermal actuator demonstrated in the Micragem process. Ourdesign also shows that a single-layer process can be used to fabricateout-of-plane actuators and that the basic spring design can potentiallywork in many different processes. The original design produced in theMicragem process had design and fabrication issues that preventedreliable assembly. Although the reliability of the devices was low-ered because of the process variations, the concept of an out-of-planeelectrothermal actuator was demonstrated and tested. Refinements tothis design will allow further improvements in the device yield andperformance.

TSANG / SAMEOTO / PARAMESWARAN: OUT-OF-PLANE ELECTROTHERMAL ACTUATORS 103

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