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Sensors and Actuators A 156 (2009) 134–144

Contents lists available at ScienceDirect

Sensors and Actuators A: Physical

journa l homepage: www.e lsev ier .com/ locate /sna

igh speed laser scanning using MEMS driven in-plane vibratory grating:esign, modeling and fabrication

u Du a,b, Guangya Zhou a,∗, Kelvin K.L. Cheo a, Qingxin Zhang b, Hanhua Feng b,in Yang b, Fook Siong Chau a

Micro and Nano Systems Initiative, Department of Mechanical Engineering, National University of Singapore, Singapore 117576, SingaporeInstitute of Microelectronics, Singapore 117685, Singapore

r t i c l e i n f o

rticle history:eceived 1 October 2008eceived in revised form 27 December 2008

a b s t r a c t

In this paper, we report a design, modeling and fabrication of a microelectromechanical system (MEMS)based in-plane vibratory sub-wavelength diffraction grating scanner for high optical efficiency, high speedlaser scanning applications. This paper demonstrates that with careful design, optical efficiency of more

ccepted 16 February 2009vailable online 28 February 2009

eywords:iffraction gratingicro-scannersicro-opto-electromechanical systems

than 75% is experimentally achievable with a simple gold-coated binary sub-wavelength grating. Highfrequency scanning of 55.01 kHz with an optical scan angle of 5.2◦ is achievable even when operatedin air. The device demonstrated has a 1 mm diameter grating. Through consideration of the mass ofthe suspension springs, stress alleviation beams and fabrication imperfections, a theoretical model isproposed which can estimate the dynamic performance close to the actual operating condition.

© 2009 Elsevier B.V. All rights reserved.

MOEMS)

. Introduction

There are numerous applications spanning a wide range ofndustries that utilize and require high speed laser scanningechnology, such as automotive head-up displays, head-worn dis-lays [1–4] and other wearable displays for mobile computingr personal electronic devices. Currently, much work in MEMS-ased laser scanners focuses on micro-scanning mirrors, due toheir outstanding advantages such as having a low mass, highcanning frequency and low power consumption. However, dueo well-known non-rigid-body deformation of the thin mirrorlate, the optical performance of the micro-scanning mirror isegraded during high speed scanning. To solve this problem,EMS-based diffraction grating scanners [5,6,12] were developed,hich utilize rotational in-plane motion rather than rotational

ut-of-plane motion (as in a conventional scanning mirror [7,8])ausing the laser beam to scan. MEMS vibratory grating scan-ers have the potential to scan at high frequencies without theptical performance degradation due to dynamic non-rigid-body

eformation (dynamic aberration). In previously proposed diffrac-ion grating scanners, the optical efficiency is much lower asompared with micro-mirror scanners [5,6] because the inci-ent light is diffracted to multiple orders and the zeroth-order

∗ Corresponding author. Tel.: +65 65161235; fax: +65 67791459.E-mail address: mpezgy@nus.edu.sg (G. Zhou).

924-4247/$ – see front matter © 2009 Elsevier B.V. All rights reserved.oi:10.1016/j.sna.2009.02.011

beam cannot be used for scanning. Increasing the scanning fre-quency also poses significant structural design challenges. Theoptical efficiency problem can be solved by using properly designedsub-wavelength binary diffraction grating while accurate model-ing and design of the devices can be achieved by a systematicapproach.

In this paper, we demonstrate the design, modeling and fabri-cation of a micromachined high speed and high optical efficiencydiffraction grating scanner, which can experimentally achieve anoptical efficiency of more than 75% and scanning at 55.01 kHz.

2. Scanner optical design

2.1. Operation principle of the vibratory grating

The operation principle of the vibratory grating is shown in Fig. 1.The diffraction grating lies in the XOY plane and its grating linesorientated parallel to the X-axis. The incident laser beam lies inYOZ plane and illuminates the diffraction grating with an incidentangle of �i. When the grating platform rotates about Z-axis in XOYplane meanwhile the diffracted beams (except the zeroth order)will scan accordingly, which is shown in Fig. 1.

2.2. Simulation of grating profile for high efficiency laser scanning

In this work, the binary diffraction grating period is designed tobe 500 nm, which is smaller than the wavelength of He–Ne laser

Y. Du et al. / Sensors and Actuators A 156 (2009) 134–144 135

seAntttIcgdtd

apmstwdei

Fig. 1. Operation principle a MEMS vibratory diffraction grating.

ource (632.8 nm) we use. According to the well-known gratingquation, only the zeroth and first order diffraction beams exist.lso, the first order diffraction beam fulfills the bow-free scan-ing condition [6,9] when the incident angle is properly set. Sincehe sub-wavelength diffraction grating [13,14] is adopted, wherehe grating period is smaller than the wavelength, scalar diffrac-ion theory is no longer valid to predict the diffraction efficiency.n this case, numerical simulations using more accurate rigorousoupled-wave analysis (RCWA) method were adopted to investi-ate the relationship between the diffraction efficiency and theetails of the binary grating profile. The simulation result showshat the diffraction efficiency of the sub-wavelength is polarizationependent.

As shown in Fig. 2, the diffraction efficiency can be as highs 80% for TM polarization, when the depth of grating grooves isroperly set. For TE polarization, the diffraction efficiency increasesonotonically with the increasing depth of the grating grooves and

aturates at about 45%. Consequently, high optical efficiency (closeo that of a coated mirror) can be achieved with a simple binary sub-

avelength diffraction grating, provided that the grating grooveepth is precisely controlled in the micro fabrication process (forxample at 125 nm) and a TM polarized incident laser is utilized forllumination during scanner operation.

Fig. 2. Diffraction efficiency as a function of grating groove depth.

Fig. 3. Schematic illustration of the micromachined in-plane vibratory grating scan-ner.

3. Scanner structure design and modeling

Fig. 3 shows the schematic illustration of the micromachinedin-plane vibratory grating scanner.

From Fig. 3, we can see that a platform with sub-wavelengthdiffraction grating is connected to the substrate through six singlebeam flexures. Two sets of electrostatic comb-drive resonators areattached symmetrically opposite to the grating platform throughone flexure each. Each single beam flexure has eight pairs of stressalleviation beams connected perpendicularly, which is used toreduce its axial stress during large deformation.

In this design, we use the symmetrically translational motionwith opposite direction to excite the rotational motion of the gratingplatform. Compared with the method that the grating is directlydriven by electrostatic circular comb-drive actuator, the rotationrange will not be decided by the displacement range of the drivingactuator but by the design of the mode shape. The mode shape canbe designed by properly choosing the mass of the micro-resonatorand the stiffness of the suspension springs. Consequently, we canpotentially achieve high frequency scanning with large scanningamplitude by using this kind of driving method.

The supporting springs, which were symmetrically connected tothe grating platform, are used to reduce the stress in each spring butkeeping scanning frequency the same. In this case, if we increasethe number of the supporting with a same scanning frequency, therotation range can be increased. On the contrary, if we keep thedimension of each spring the same but increase the number ofsupporting springs, the scanning frequency will increase but therotation range will not change. Consequently, we need to put asmany supporting springs as possible in the allowable space.

3.1. Initial model of the system

The dynamics of the MEMS in-plane grating scanner can be con-sidered to be a 2 degree-of-freedom (DOF) vibration system. Herewe make the following assumptions to obtain the simplest dynamicmodel of the system:

(1) The dynamic model of the system is linear under small defor-mations.

(2) Mass of all the flexure suspensions in the system is ignored andonly the moment of inertia of the grating plate and the mass ofthe micro-resonator are considered.

(3) Stress alleviation beams are considered to be ideal, which canrelease the axial stress but not reduce the spring constant of thesingle beam flexure.

136 Y. Du et al. / Sensors and Actuato

Fv

(

u

E

Tdbs

[

Wsm

u

b

platform has a length of 880 �m, width of 31 �m and thickness of

ig. 4. Schematic illustration of the simplified 2 DOF undamped free spring-massibration system.

4) Grating plate and micro-resonators are treated as rigid-bodyso that all the suspension beams can have a perfect boundarycondition.

Fig. 4 shows the schematic view of the simplified 2 DOFndamped spring-mass vibration system.

During vibration, the total kinetic energy of the system is:

k = 212

mx2 + 12

J�2 = mx2 + 12

J�2 (1)

o calculate the potential energy of the suspension beam undereformation, we begin by looking at the stiffness matrix of a singleeam. Fig. 5 shows the local variables defined for a single suspen-ion beam.

The stiffness matrix of the above beam is:

k] = EI

L3

⎡⎢⎢⎢⎢⎢⎢⎢⎢⎢⎢⎢⎣

AL2

I0 0 −AL2

I0 0

0 12 6L 0 −12 6L

0 6L 4L2 0 −6L 2L2

−AL2

I0 0

AL2

I0 0

0 −12 −6L 0 12 −6L

0 6L 2L2 0 −6L 4L2

⎤⎥⎥⎥⎥⎥⎥⎥⎥⎥⎥⎥⎦

(2)

here L is the length, I is the moment of inertia and A is the cross-ectional area of the beam and E is the Young’s modulus of theaterial.

The corresponding vector of local variables is then:[ ]T

� = u1 v1 �1 u2 v2 �2 (3)

We can get the vectors of local variables for both connectioneam and supporting beam under small deformation assumption.

Fig. 5. Local variables defined wrt suspension beam.

rs A 156 (2009) 134–144

Where:⎧⎨⎩

�uc =[

0 −x 0 0 −R� �]T

�us =[

0 0 0 0 −R� �]T

(4)

with everything in place, the potential energy of the system is then:

Ep = 212

k0x2 + 212

�uTc [kc]�uc + 6

12

�uTs [ks]�us (5)

where:

[kc] = EIc

L3c

[12 −12 6Lc

−12 12 −6Lc

6Lc −6Lc 4L2c

]and [ks] = EIs

L3s

[12 −6Ls

−6Ls 4L2s

]

Applying the Lagrange formulation, we can obtain the linear modelof the system to be:[

m 00 J

][x�

]+

[nf k0 + kc1 −c1kc1−ncc2kc2 nsk� + nckc2

][x�

]= 0 (6)

where:

k0 = 12EIf

L3f

, kc1 = 12EIc

L3c

, c1 = R + Lc

2,

kc2 = 12EIc

L3c

(R2 + LcR + L2

c

3

), k� = 12EIs

L3s

(R2 + LsR + L2

s

3

),

c2 = R + (Lc/2)

R2 + LcR + (L2c /3)

, Ic = 112

W3c T, Is = 1

12W3

s T,

If = 112

W3f T and J = 1

2�Si�R4T.

nf: Number of folded beam suspensions in each micro-resonator;nc: number of connection beam suspensions; ns: number of sup-porting beam suspensions; k0: spring constant of each foldedbeam suspension; kc1: coupling spring constant for resonator; c1:rotational to translational motion coefficient; kc2: coupling springconstant for grating platform; c2: translational to rotational motioncoefficient; k�: spring constant for grating platform; Ic, Is, If: areamoment of inertia of connection spring, supporting spring andfolded beam suspension respectively; Wf: width of a single beamin folded beam suspension; �Si: density of silicon.

Through Eq. (6), we can obtain the natural frequencies and modeshapes for both the 1st and 2nd vibration modes. These are used toget a preliminary estimate of the vibration characteristics of thesystem.

We attempted a design where the diameter of the grating plateis 1 mm. Each micro-resonator has a total number of 293 movablefingers on one side with finger length of 60 �m, width 5 �m, thick-ness of 10 �m, finger gap of 3 �m and initial finger overlap length of30 �m. The micro-resonator is suspended by 53 sets of folded beamsuspensions with beams of length 95 �m, width 5 �m and thick-ness 10 �m. Each single beam flexures connected to the grating

10 �m. Besides, each stress alleviation beam has a length of 338 �m,width of 5 �m and thickness of 10 �m. The material properties ofthe single crystal silicon used in the model are shown below inTable 1.

Table 1Material properties used in the calculation.

Parameter Value

Young’s modulus 169 GPaPoisson’s ratio 0.28Density 2330 kg m−3

Y. Du et al. / Sensors and Actuators A 156 (2009) 134–144 137

Table 2Summary of theoretical and FE simulation results.

Mode Model without consideringthe mass of the beam

Model considering themass of the beam

FE simulation

1stmode

Frequency (Hz) 76527 67371 66940Frequency deviation from FE simulation 14.32% 0.64% N/AMode shape (◦/�m) 0.5126 2.3200 2.3026Mode shape deviation from FE simulation 77.74% 0.76% N/A

2m

Frequency (Hz) 79466 77895 77550

um

3

taes

sTot{

f

c

E

E

⎪⎪⎪⎨⎪⎪⎪⎩

finner(u) =L3

−L2

+ 22

fouter(u) =(

3u2

L2− 2u3

L3

)x

2

(13)

ndode

Frequency deviation from FE simulation 2.47%Mode shape (◦/�m) −0.2911Mode shape deviation from FE simulation 468.55 %

For the proposed prototype grating scanner, the calculated nat-ral frequency and mode shapes using the simple linear dynamicodel are compiled in Table 2.

.2. Weight influence of the suspension

However, a more rigorous model is needed to take into accounthe effects of the connection, supporting springs and the stresslleviation beams to obtain a more accurate estimation. We willstablish a model for the connection springs first. The model andmall angle approximations are shown below in Fig. 6.

The model essentially considers the deformation profile of thepring due to a rotation at the end connected to the grating platform.he deflection of the connection spring is approximated by a thirdrder polynomial expression that fulfills the boundary condition ofhe beam which are given below:

f (0) = R sin � ≈ R�, f (L) = xf ′(0) = tan � ≈ �, f ′(L) = 0

(7)

The resulting deflection along the beam is a function of x and �

(u) =(

(2R + L)u3

L3− (3R + 2L)u2

L2+ u + R

)� +

(3u2

L2− 2u3

L3

)x

(8)

The additional kinetic energy for the connection spring can beomputed by

k c =∫ Lc

0

12

f 2(u)dm = 12

�SiWcT

∫ Lc

0

f (u)du (9)

Resulting in

k c = 12

�SiWcTLc

(1335

R2+ 1105

L2c + 11

105LcR

)�2+ 1

2�SiWcTLc

1335

x2

+ 12

�SiWcTLc

(13

210Lc + 9

35R)

x� (10)

Fig. 6. Model of the connection suspension beam.

0.44% N/A−0.0518 −0.05121.17% N/A

We can simply get the additional kinetic energy of the supportingbeam by letting x = 0, resulting in

Ek c = 12

�SiWcTLc

(1335

R2 + 1105

L2s + 11

105LsR

)�2 (11)

Next we attempt to take into account the effects of the folded beamsthat are distributed throughout the device. The model used is sim-ilar to the connection springs and shown in Fig. 7.

The deflection of the beam is also approximated by a third orderpolynomial expression that fulfills the boundary condition of thebeam. The boundary conditions are shown below:{

f (0) = x1, f (L) = x2

f ′(0) = 0, f ′(L) = 0(12)

For the inner two beams,{x1 = x

x2 = (x/2)

and for outer two beams{x1 = 0

x2 = (x/2)

The resulting deflection and velocity profile along the beam isagain expressed as a function of x,⎧ (

2u3 3u2)

x

Fig. 7. Model of the folded beam suspension.

1 ctuato

s

E

R

E

E

p

c 2 c Si c c 210 c 35

38 Y. Du et al. / Sensors and A

The additional kinetic energy for one set of folded beam suspen-ion can be computed as

k f = 2

∫ Lf

0

12

[f 2inner(u) + f 2

outer(u)]dm

= �SiWf T

∫ Lf

0

[f 2inner(u) + f 2

outer(u)]du (14)

esulting in

k f = 12

�SiWf TLf4835

x2 (15)

The modified kinetic energy of the whole system is then:

˜k = Ek + Ek c + Ek s + Ek f (16)

Applying the Lagrange formulation, with only the kinetic com-onent changed, we can obtain the linear model of the system to

Fig. 8. FE models showing the vibration profiles for (a) out-of-plane 1st mode, (

rs A 156 (2009) 134–144

be:[m + m1 Jcmc J + J1

][x�

]+

[nf k0 + kc1 −c1kc1−ncc2kc2 nsk� + nckc2

][x�

]= 0

(17)

where:

m1 = �SiWcTLc1335

+ nf �SiWf TLf4835

,

Jc = 12

�SiWcTLc

(13

210Lc + 9

35R)

,

m = 1n � W TL

(13

L + 9R)

,

J1 = nc�SiWcTLc

(1335

R2 + 1105

L2c + 11

105LcR

)+ ns�SiWsTLs

(1335

R2 + 1105

L2s + 11

105LsR

)

b) out-of-plane 2nd mode, (c) in-plane 1st mode, (d) in-plane 2nd mode.

Y. Du et al. / Sensors and Actuators A 156 (2009) 134–144 139

asvB

ttut

qmripsopt

3

oostetbaa

TS

M

1

2

Fig. 9. FE torque angle curves for connection beam suspensions.

Finite element (FE) simulations were done to verify if thepproximations were reasonable. We compared the model repre-ented in Eq. (6) with the modified version which considered thearious beams in Eq. (17) and the finally with the FE simulations.oth the 1st mode and 2nd mode frequencies were compared.

Fig. 8 shows some results of the vibration modes obtainedhrough FE simulations. The out-of-plane frequencies were foundo be 8552 Hz for 1st mode and 22,253 Hz for 2nd mode. These fig-res are significantly lower than the in-plane modes which are inhe range of 60 kHz, thus not too significant during operation.

The modified model shows closer approximated vibration fre-uencies than the simple linear model in Table 2. The effect of theultiple suspensions springs and folded beams has an effect of

educing the effective frequencies in the device as expected. Theres a significant frequency drop for the 1st resonating mode com-ared to the 2nd mode, besides a significant increase in the modehape for the 1st mode and a decrease in the 2nd mode. As the massf the resonating structure is significantly larger than the gratinglatform, the inclusion of the beam mass has a greater influence onhe frequency of the grating platform.

.3. Influence of the stress alleviation beams

Several stress alleviation beams were introduced at the anchorf the suspension springs to help in improving the performancef the device. They permit some translational motion to releasetress at the anchor locations yet minimizing rotational motion athe boundary. In an ideal case, the stress alleviation beams have noffect on the linear harmonic solution. This is necessary as the rota-

ion of the grating platform introduces tension in the suspensioneams. Fig. 9 shows FE simulation results of the improved torque-ngle relationships with and without the addition of these stresslleviation beams compared to the ideal linear spring model.

able 3ummary of theoretical and FE simulation results with stress release springs.

ode Model conof the beam

st mode Frequency (Hz) 67371Frequency deviation from FE simulation 6.88%Mode shape (◦/�m) 2.3200Mode shape deviation from FE simulation 29%

nd mode Frequency (Hz) 77895Frequency deviation from FE simulation 1.70%Mode shape (◦/�m) −0.0518Mode shape deviation from FE simulation 41.92%

Fig. 10. Effect of the stress alleviation beams.

Fig. 10 illustrates how the stress alleviation beams help in reduc-ing stress at the anchor locations. More importantly, it shows aninfluence in affecting the resonance frequencies of the structurewhich we have to account for.

Due to the slight tile �� at the anchor location, the vectors oflocal variables for both connection beams and supporting beamsfor Eq. (4) have to be changed to⎧⎨⎩

�uc =[

0 −x −�� 0 −R� �]T

�us =[

0 0 −�� 0 −R� �]T

(18)

Due to this change, the modified model in Eq. (17) becomes[m + m1 Jcmc J + J1

][x

�

]+

[nf k0 + kc1 −c1kc1

−ncc2kc2 nsk� + nckc2

][x

�

]= 0

(19)

where

c1 = R + Lc(1 − ��/�)2

,

c2 = 12R + 6Lc

12R2 + 12LcR + 4L2c − (2L2

c + 6LcR) ���

,

kc2 = EIc

L3c

[12R2 + 12LcR + 4L2

c − (2L2c + 6LcR)

��

�

],

k� = EIs

L3s

[12R2 + 12LsR + 4L2

s − (2L2s + 6LsR)

��

�

]

FE simulations were again done to compare the overall effectson the system. Results were compared again with the initial linearmodel and the Eq. (20) model in Table 3. Fig. 11 shows some FEsimulation results illustrating the different vibration modes and

sidering the mass Model considering the mass ofthe beam and release spring

FE simulation withrelease spring

63846 630331.29% N/A3.4853 3.31425.16% N/A

77807 765891.59% N/A−0.0352 −0.03653.56% N/A

140 Y. Du et al. / Sensors and Actuators A 156 (2009) 134–144

F a) out2

p1t

taairtcv

3

iTst

ce

˜

ig. 11. FE models with stress alleviation beams showing the vibration profiles for (nd mode.

rofiles. The 1st and 2nd order out-of-plane modes are 6039 and7418 Hz respectively. These figures are still significantly lower thanhe in-plane modes.

Consideration of the stress alleviation beams has a similar effecto the beam mass modeling earlier: both change the mode shapend reduce the natural frequency. However in this case, the modelffects the effective spring constant of the vibrating structures. Sim-larly, as the grating platform has smaller mass, the change in theesonant frequency is more significant for 1st mode. The effects forhe mode shapes are equally significant for 1st and 2nd mode. Wean see from Table 3 that the theoretical model with the stress alle-iation beams show closer agreement to the FE simulation results.

.4. Influence of fabrication imperfections

Another common factor on device characteristics is fabricationmperfections. And the effect of etching profiles is most significant.he plasma etching is widely known to have such errors [11]. Fig. 12

hows a cross-sectional profile of a beam from a previous device andhe model we used based on it.

The cross-sectional model attempts to account for a slight under-ut in the profile as evident in �W. Also a slight slope, which isxpressed as the angle ˛ of the sidewall, in the etching profile is

-of-plane 1st mode, (b) out-of-plane 2nd mode, (c) in-plane 1st mode, (d) in-plane

anticipated. The fabrication imperfections such as undercut duringoxide and silicon etching process, etching slope, etc. will mainlyinfluence the area moment of inertia of each beam flexure suspen-sion.

The relationship between the actual beam width and beamthickness is shown below.

w(u) = 2ctg˛u + W − 2(�W + Tctg˛) (20)

A new area moment of inertia based on these assumptions isthen

I = 112

T(W3 − 3Tctg˛W2 − 4T2ctg2˛W + 2T3ctg3˛) (21)

where W = W − 2�WThe final calculated frequencies for all the models we had

explored so far as well as the measured resonant frequencies aregiven in Table 4. We see a declining resonance frequency as moredetailed compensation is incorporated into the models.

Consideration of the mass of the beams, stress alleviation beamsand finally the fabrication imperfections all reduce the theoreti-cal frequencies as compared to the initial linear model. Significantchanges also occur at the mode shape parameter. As the mass ofthe resonating structure is much larger than the grating platform,

Y. Du et al. / Sensors and Actuators A 156 (2009) 134–144 141

Table 4Summary of theoretical results using different models.

Mode Initial model Model only consider themass of the beam

Model consider the mass ofthe beam and release spring

Model consider the mass of thebeam, release spring andfabrication imperfections

Experimentalresults

1st Mode Frequency (Hz) 76527 67371 63846 57389 550103

2 7−

ttmsstit

4

(t(c

aTawd

dcupwcaw

F

The optical efficiency was measured by measuring the intensityof both the incident and diffracted laser beam. Fig. 16 shows theschematic view of the experimental setup of optical efficiencymeasurement. The beam splitter was used to split the incidentbeam into two beams equally and the intensity of one of them

Mode shape (◦/�m) 0.5126 2.3200

nd mode Frequency (Hz) 79466 77895Mode shape (◦/�m) −0.2911 −0.0518

he consideration of the reduced spring constants due to fabrica-ion imperfections have a larger effect on overall frequency of the

icro-resonating structure. The large quantity of the folded beamuspensions and the narrow beam widths results in being more sen-itive to fabrication errors. This results in a significant reduction forhe 1st mode shape and increase for the 2nd mode shape. Compar-ng all the models, the overall effect of fabrication imperfection ishe most significant.

. Fabrication process

The prototype device was fabricated using silicon-on-insulatorSOI) micromachining technology. The SOI wafer used has a 10 �mhick heavy doped silicon device layer, 2 �m thick buried oxideBOX) layer and a 750 ± 25 �m thick silicon substrate. The fabri-ation process flow is illustrated in Fig. 13.

As shown in Fig. 13, the diffraction grating with 500 nm periodnd 50% duty cycle was transferred by using timed plasma etching.he etching time is strictly controlled so that the grating groove isround 1250 Å. Fig. 14 shows the resulting SEM image of the sub-avelength diffraction grating with 500 nm grating period and 50%uty cycle.

Subsequently 5000 Å PECVD undoped silicon glass (USG) waseposited and patterned by using reactive ion etching (RIE) pro-ess. Next the 10 �m device layer was etched until the BOX layer bysing deep reactive ion etching (DRIE) process to form the grating

late, micro-resonator and connection suspension. The SOI waferas then patterned on the backside followed by another DRIE pro-

ess to remove the silicon and expose the region of grating platend connection suspension. The grating plate and micro-resonatorere released from backside and front side respectively by using

ig. 12. Cross-sectional profile model of etched beam and corresponding SEM image.

.4853 0.1262 N\A

7807 61833 583700.0352 −0.9017 N/A

BOE solution (buffered oxide etchant solution with six parts of 40%NH4F and one part of 49 HF). After that, a 200/5000 Å Cr/Au layerwas evaporated to form the wire bonding pads through a shadowmask. Finally, a 50/800 Å Cr/Au layer was evaporated on the wafersurface to enhance the reflectivity of the diffraction grating.

The SEM image of the final fabricated device is shown in Fig. 15.

5. Experimental results

The optical performance of the prototype MEMS diffraction grat-ing scanner was tested using a linearly polarized He–Ne laser beamwith a wavelength of 632.8 nm. The incident angle is determined tobe 28.4◦ so that the bow-free scanning condition [6,9] is fulfilled.

Fig. 13. Fabrication process flow of the MEMS vibratory grating scanner.

142 Y. Du et al. / Sensors and Actuators A 156 (2009) 134–144

iib(dv

eat8qfitio

W

lower stress level in connection flexures so that the risk of brittlefracture of the connection beam is reduced. Fig. 18 shows the mea-sured frequency response of the MEMS grating scanner at frequencyregion near the resonant frequency of first vibration mode.

Fig. 14. SEM image of micromachined sub-wavelength diffraction grating.

s measured by Photo detector 01. The other split laser beam isllumined on the diffraction grating and the 1st order diffractioneam was measured by photo detector 03. High optical efficiencymore than 75% of the energy of incident light) was experimentallyemonstrated with the TM-polarized laser beam, which matchesery well with the simulation results.

The MEMS grating scanner was operated in atmosphere andach micro-resonator was driven by lateral electrostatic comb-drivectuator using a push–pull mechanism [10]. The schematic view ofhe experimental setup is shown in Fig. 17. The DC bias voltage was5 V and the AC peak-to-peak voltage was 170 V with a driving fre-uency of 55.01 kHz, which is close to resonant frequency of therst vibration mode. A projection screen is put perpendicularly to

he 1st order diffraction with a certain distance. Through measur-

ng the length of the laser scanning trajectory, we can calculate theptical scanning angle, which is estimated to be around 5.2◦.

The MEMS grating scanner has two vibration modes as expected.e selected frequency of the first mode as the working because

Fig. 15. SEM image of the MEMS vibratory grating scanner.

Fig. 16. Schematic view of the experimental setup for optical efficiency measure-ment.

it has higher scanning amplitude compared to second mode and

Fig. 17. Schematic view of the experimental setup for prototype scanner.

Y. Du et al. / Sensors and Actuato

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ig. 18. Measured frequency response of the grating scanner at frequency regionear the resonant frequency of first vibration mode.

The measured resonant frequency of the device is about 4.6%ower than the theoretical calculation results we obtained in therevious section. The reduction is mainly due to the imperfec-ions during the micromachining process. The estimated amountf undercut and etching slope might be different than modeled.aterial properties and imperfect boundary conditions of all the

uspension beams are other possible factors. Nevertheless, we hadbtained an estimate that is much closer to the experimental results

f the various effects had not been considered.

. Conclusions and discussions

In this paper, we have introduced the design, fabrication ofnovel micromachined high speed, high optical efficiency laser

canner using a sub-wavelength diffraction grating. High opticalfficiency was experimentally achieved that more than 75% of thencident beam energy can be used for scanning. The current proto-ype scanner is capable of scanning at a high frequency of 55.01 kHzith an optical scan angle of 5.2◦. Compared with other micro-

canners, the scanner has higher scanning speed but lower opticalcanning angle.

For the prototype device, the symmetrically translationalotion with opposite direction is used to excite the rotationalotion of the grating platform. The advantage of this kind of driv-

ng method is that the rotation range will not be decided by theisplacement range of the driving actuator but by the design of theode shape. However large space is required by using this kind of

onfiguration. This is a limitation of this scanner because the size ofhe device cannot be too large since the thickness of the structures only 10 �m.

Consequently, the mass of the micro-resonator, the numberf suspension springs and the stiffness of the suspension springshould be properly designed with considering the requirement ofhe scanning frequency, scanning amplitude as well as the size ofhe device.

We have also demonstrated the validity of a system modelhich takes into account the effects of the springs masses, stress

lleviation beams and fabrication imperfections. All affects the sys-

em significantly. However the influence of fabrication errors washown to have the largest effect on system characteristic. Neverthe-ess, all was shown to affect to certain degrees and future designs

ould have to take them into account. This work provides a moreccurate platform which to base future designs on.

rs A 156 (2009) 134–144 143

References

[1] H. Urey, D. Wine, T. Osborn, Optical performance requirements for MEMS-scanner based micro displays, Proc. SPIE, MOEMS Miniaturized Syst. 4178(2000) 176–185.

[2] H. Urey, D. Wine, J.R. Lewis, Scanner design and resolution tradeoffs for minia-ture scanning displays, in: Proceedings of SPIE Flat Panel Display Technologyand Display Metrology, vol. 3636, San Jose, CA, January, 1998, pp. 60–68.

[3] D. Wine, M.P. Helsel, L. Jenkins, H. Urey, T.D. Osborn, Performance of a bi-axialMEMS-based scanner for micro display applications, Proc. SPIE 4178 (2000)186–196.

[4] R.B. Sprague, T. Montague, D. Brown, Bi-axial magnetic drive for scannedbeam display mirrors, Proc. SPIE, MOEMS Display Imaging Syst. III 5721 (2005)1–13.

[5] G. Zhou, L. Vj, F.S. Chau, F.E.H. Tay, Micromachined in-plane vibrating diffractiongrating laser scanner, IEEE Photonic Technol. Lett. 16 (2004) 2293–2295.

[6] G. Zhou, F.S. Chau, Micromachined vibratory diffraction grating scanner for mul-tiwavelength collinear laser scanning, IEEE/ASME J. Microelectromech. Syst. 15(2006) 1777–1788.

[7] P.M. Hagelin, O. Solgaard, Optical raster-scanning displays based on surface-micromachinedpolysilicon mirrors, IEEE J. Sel. Top. Quantum Electron. 5 (1999)67–74.

[8] R.S. Muller, K.Y. Lau, R.S. Muller, K.Y. Lau, Surface-micromachined micro opticalelements and systems, Proc. IEEE 86 (1998) 1705–1720.

[9] L. Beiser, Holographic Scanning, John Wiley & Sons Inc., New York, 1988, pp.206–215.

10] W.C. Tang, T.C.H. Nguyen, M.W. Judy, R.T. Howe, Electrostatic comb-drive oflateral poly-silicon resonators, Sens. Actuators A 21/23 (1990) 328–331.

11] J. Li, Q.X. Zhang, A.Q. Liu, W.L. Goh, J. Ahn, Technique for preventing stiction andnotching effect on silicon-on-insulator microstructure, J. Vac. Sci. Technol. B 21(2003) 2530–2539.

12] G. Zhou, Y. Du, Q. Zhang, H. Feng, F.S. Chau, High-speed, high-optical-efficiencylaser scanning using a MEMS-based in-plane vibratory sub-wavelength diffrac-tion grating, J. Micromech. Microeng. 18 (2008), 085013.

13] D.W. Carr, J.P. Sullivan, T.A. Friedmann, Laterally deformable nanomechanicalzeroth-order gratings: anomalous diffraction studied by rigorous coupled-waveanalysis, Opt. Lett. 28 (September (18)) (2003) 1636–1638.

14] U. Krishnamoorthy, R.H. Olsson III, G.R. Bogart, M.S. Baker, D.W. Carr, T.P. Swiler,P.J. Clews, In-plane MEMS-based nano-g accelerometer with sub-wavelengthoptical resonant sensor, Sens. Actuators A 145–146 (2008) 283–290.

Biographies

Yu Du received the B.S. degree in Mechanical Engineering from the Xi’an JiaotongUniversity, Shaanxi, China, in 2005. He is currently a Ph.D. student in Micro andNano System Initiative (MNSI) at National University of Singapore (NUS), Singapore.His interest lies in optical MEMS and his current work focuses on the design andfabrication of optical micro devices.

Guangya Zhou received the B.Eng. and Ph.D. degrees in optical engineering fromZhejiang Univeristy, Hangzhou, China, in 1992 and 1997, respectively. He joined theNational University of Singapore (NUS) in 2001 as a Research Fellow. Currently, heis an assistant professor with the Department of Mechanical Engineering, NUS. Hismain research interests include microoptics, diffractive optics, MEMS devices foroptical applications, and nanophotonics.

Kelvin Koon Lin Cheo received the B.S. and M.S. degrees in Mechanical Engineeringfrom National University of Singapore, Singapore, in 2003 and 2005. He is currentlya Ph.D. student in Micro and Nano System Initiative (MNSI) at National University ofSingapore (NUS), Singapore. His interest lies in optical MEMS and his current workfocuses on the system level applications of optical micro devices.

Qingxin Zhang received the B.S. and M.S. degrees in semiconductors devices andphysics from Harbin Institute of Technology, Harbin, China, in 1986 and 1989,respectively, and the Ph.D. degree in microelectronics from the Institute of Micro-electronics, Tsinghua University, Beijing, China, in 1997. After 2 years of being aResearch Fellow with Nanyang Technology University, Singapore, he was with theInstitute of Microelectronics, Agency for Science, Technology and Research, Singa-pore, in 1999, where he is currently a Member of Technical Staff working with theSemiconductor Process Technologies Laboratory. His major interests include MEMSdesign and process, system-on-package and platform technologies for integratedMEMS, CMOS IC, and photonics.

Hanhua Feng received the B.Eng. degree in semiconductor physics and devicesand the M.Eng. and Ph.D. degrees from Huazhong University of Science and Tech-nology (HUST), Wuhan, China, in 1985, 1988, and 1991, respectively. She was alecturer/associate professor in electronic materials and devices for two years withHUST. She was a postdoctoral and research fellow with the Northern Ireland Bio-Engineering Center, University of Ulster, Newtownabbey, U.K., and the School of

Electrical and Electronics Engineering, Nanyang Technological University, Singapore.In January 1998, she joined the Institute of Microelectronics, Singapore, where she iscurrently a MEMS Program Manager. She is familiar with the aspects of wafer fabri-cation, with expertise particularly in CMOS/MEMS integration, PECVD, and plasmaetching, from her many years of working experience in the microelectronics andMEMS areas.

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44 Y. Du et al. / Sensors and A

in Yang received the B.S. degree in materials science from Anhui University ofechnology and Science, China. He received his M.S. degree in materials science frominlin University in 2003 and Ph.D. degree in electronic science and technology fromhanghai Jiaotong University in 2006, China. From 2006 to 2007, he was employed asresearch fellow at the department of mechanical engineering, National Universityf Singapore. In November 2007, he has joined Institute of Microelectronics (IME)

rs A 156 (2009) 134–144

as senior research engineer. His research interests include optoelectronics, MEMSsensors and transducers.

Fook Siong Chau is an Associate Professor in the Department of Mechanical Engi-neering, National University of Singapore, where he heads the Applied MechanicsAcademic Group. His main research interests are in the development and appli-cations of optical techniques for nondestructive evaluation of components and themodeling, simulation, and characterization of micro systems, particularly bio-MEMSand MOEMS.