Optically interrogated MEMS pressure sensor...

14th Int Symp on Applications of Laser Techniques to Fluid Mechanics Lisbon, Portugal, 07-10 July, 2008

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Optically interrogated MEMS pressure sensor array

Lukas Prochazka1, Thomas Roesgen

2

1: Institute of Fluid Dynamics, ETH, Zurich, Switzerland, [email protected]

2: Institute of Fluid Dynamics, ETH, Zurich, Switzerland, [email protected]

Abstract A novel pressure measurement technique is developed with the objective to simplify wind tunnel investigations while providing full field measurement capability with highest measurement performance. The static pressure distributions are recorded by interferometric means using an array of silicon diaphragm microresonators as pressure sensing elements. They are remotely excited into free vibration using a capacitive ultrasound transducer (CUT) transmitting high power narrow band (50 – 80kHz) acoustic noise. Dependent on their quasi-static deflection caused by a pressure load the resonance frequency varies with an average pressure sensitivity of 3Hz/Pa at 70kHz. Further requirements for the system performance are a maximum pressure, temporal and spatial resolution of 0.2% of FS (10kPa), 1Hz and 5mm, respectively.

1. Introduction

During aerodynamic and mechanical load performance surveys in wind tunnel testing it is often

desirable to obtain the surface pressure on a specific component of the investigated model with high

spatial resolution, good measurement accuracy and low uncertainty. Despite the great demand for

static pressure measurements only the pressure tap method tends to be used in commercial and

industrial wind tunnel facilities. This technique, where the transducers are linked by a tube to the

measurement locations, provides high measurement performance but is also very costly due to the

required mechanical and hydraulic interfacing in the test model. The application of pressure

sensitive paints for low speed testing at atmospheric pressure conditions is often unsuitable due its

low sensitivity and large temperature cross-sensitivity (Liu et al. 2005 and Airaghi 2006).

The objective of the present work is to develop an alternative pressure field measurement technique

with measurement performance (i.e. high accuracy, low measurement uncertainty) comparable to

the pressure tap system but more simple handling and installation requirements. The requirements

on the measurement performance are summarized in Table 1.

Table 1 Required measurement performance

Temporal resolution

Spatial resolution

Measurement range

High sensitivity

Low measurement uncertainty (drift, cross sensitivity, …)

≈ 1Hz

> 5mm

< 10kPa

0.2% FS or 20Pa

In order to define a suitable design concept different techniques and devices for static pressure

measurements of different application fields were compared and evaluated regarding the above

mentioned criteria (Prochazka et al. 2005). Based on this study two basic conclusions were

identified.

1. The need for high measurement performance (sensitivity, uncertainty) in low speed wind

tunnel testing requires a pressure sensing device with a deflecting sensing element operating

in differential mode, e.g. membrane or diaphragm.


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2. To allow for a simple installation of the device, the sensor unit has to be mounted directly

on the model surface and its interrogation has to be wireless (no electrical or hydraulic

interfacing)

Wireless interrogation of a sensor array by electromagnetic waves in the microwave or radio

frequency (RF) range has the advantage of permitting an operation even without visual contact

between sensor and interrogation unit. On the other side the transmitter on the sensor side

(electronics and power supply) would again require complex installation within generally small

wind tunnel models. The delivery of the necessary power by RF over a distance of around 2 meters

would require high transmitting power with the potential for electromagnetic interference with other

highly sensitive electronic devices within the wind tunnel environment. An optical interrogation

causes no electromagnetic interference and the necessary visual contact between sensor and detector

is usually given due to several observation points in a test section. The challenging task hereby is to

detect the expected small signal carrying the pressure information within a bright, vibrating

background and over the mentioned interrogation distance.

The concept which is regarded as the best suitable measurement technique for the mentioned

application is based on tiny and passive MEMS pressure sensor dies which are mounted directly on

the model surface and are optically interrogated using a highly sensitive active pixel CMOS lock-in

detector.

2. Method

The pressure sensing system is based on the optical interrogation of an array of passive MEMS

pressure sensors (silicon diaphragms with a sealed cavity) which may be embedded in a highly

elastic filler material and are mounted directly on the model surface (Fig.2.1). Dependent on the

surface geometry, the model size and application, a thin (0.2 mm) sensor foil with a certain number

of sensing elements or individual tags equipped with a single sensor are conceivable. In case of

small-scale models, where even 0.2mm thick and 5mm large obstacles could alter the flow

characteristics, the tags may be recessed in tiny milled cavities in the model surface (Fig.2.2).

Fig. 2.1 Measurement setup including the remote imaging

interferometer and the ultrasound transducer

Fig. 2.2 Sensor design and tags mounted flush to the

surface of the model (one possible arrangement)

As a measure for the applied pressure the resonance frequency of the clamped silicon diaphragm is

used for the present setup which depends on its quasi-static deflection and thus on the applied

pressure. We believe that in combination with a wireless and remote readout capability of the

sensors of about 2 meters the measurement of frequency instead of static deflection is more accurate


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and robust. An intensity modulating signal with the particular frequency may arise due to the

interference between the object and reference beam, due to an alternating speckle pattern or even

due to the deflection of the reflected light on the vibrating diaphragm. The resonant sensor

characteristic of a ‘digital’ frequency output signal (i.e. basically independent of analog levels)

simplifies the installation and adjustment of the optical setup significantly.

The readout is achieved in a two-stage configuration. First, an ultrasound transducer (UT) (emitting

band-limited noise) excites the diaphragms into free oscillation. After that, the actual resonant

vibration responses of the individual sensing elements are measured with an imaging interferometer

(or a speckle setup). This device is based on a highly sensitive, “active-pixel” CMOS camera

(CSEM, Switzerland) that works as a full field lock-in detector with a spatial resolution of 144 x 90

pixels (> 104 measurement points) (Beer et al. 2005). It can resolve signal frequencies up to 100kHz

while providing very strong background (i.e. static image) rejection and a frame rate of 6kHz.

Fig. 2.3 CUT (SensComp, Inc.) used in the present setup

and its power supply (self-made). Transducer diameter ≈

40mm

Fig. 2.4 Transmitting characteristic of the CUT

(SensComp, Inc.) operated with 200 VoltDC and

350 VAC p-p (Sound pressure measured at 1m from

transducer)

Using ultrasound allows the excitation of several sensing devices in parallel without altering the

sensing behavior due to heating as would be a problem with laser excitation or due to specific

actuators as in the electrical case. The difficulty associated with this excitation principle is mainly

related to the high losses caused by strong attenuation of ultrasound in air and thus the need for a

high power UT to allow for sufficiently high excitation over an operating distance of about 2

meters. High power UTs for applications in air are mainly of capacitive type (CUT). Advantages of

CUTs are their high transfer coefficient (high transmitting efficiency in air) in combination with a

wide bandwidth and the simple design that is reflected in their low cost (Manthey et al. 1992 and

Rafiq et al. 1991). Figure 2.3 shows the CUT used in the present setup (SensComp, Inc.) with its

large noise generating diaphragm of about 40mm and its DC and AC power supply. At almost

maximum operating power (200 VDC and 350 VAC p-p) the specific CUT provides a sound pressure

around 120dB (at 1m from transducer) in a wide frequency range from 50 to 100 kHz (Fig. 2.4).

3. Experimental verification of the sensing principle

The pressure sensing behavior of a plate resonator can be described by simple semi-empirical

formulas (Fabula et al. 1997), eq. 3.1 – 3.

( ) ( )2

25.110

+==

t

ddfdf resres (3.1)


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)1(122

16.35)0(

22 νρπ −==

E

a

tdf res (3.2)

+

−

=

32

258.12.4

)1(16

t

d

t

d

a

tEp

ν (3.3)

Eq. 3.1 describes the variation of the resonance frequency as function of the static diaphragm

deflection d and its thickness t whereas eq. 3.2 gives a formula for the resonance frequency of a

square, non deflected, clamped diaphragm. The variable a represents the edge length of the

diaphragm and the material properties are defined by the Young’s modulus E, the Poisson’s ration ν

and the material density ρ. The static deflection due to loading of the diaphragm, i.e. pressure load

p, can be calculated iteratively using the implicit formula given by eq. 3.3.

Although experimental data associated with the plate resonator can be found in literature (Defaÿ et

al. 2002, Smits et al. 1985 and Uttamchandani et al. 1987) they are insufficient to establish a

reliable validation of the above mentioned approach. Therefore experimental tests with common

pressure sensor dies were carried out using a pressure and temperature controlled test cell. The

pressure outside the resonator and in the cavity behind the diaphragm could be adjusted

independently of each other between 0.8bar below and above ambient pressure with an accuracy

better than 1mbar (< 0.1% FS). The temperature was set and controlled with a Peltier heater and

two heat exchangers to values between 20 and 35ºC with an accuracy better than 0.02K (Fig.3.1).

Fig. 3.1 Experimental setup for the investigation of sensor behavior

Due to the hermetically sealed cell and limited internal space it was not possible to combine

acoustic excitation and pressure control inside the cell. Instead an external, pulsed laser diode

(Flexpoint 10mW, Laser Components GmbH) was used to heat the diaphragm locally and thus to

generate mechanical stresses which excite the diaphragm into vibration with a certain excitation

frequency. The dynamic amplitude was then measured with a Laser Doppler Vibrometer OFV-551


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(Polytec GmbH) from outside of the test cell as well (Fig.3.1). During a test run the excitation

frequency was swept across the range of interest with frequency steps up to 20Hz in the

neighborhood of the resonance peak. The resonance frequency was then found at the maximum of

the amplitude spectrum (Fig. 4.2).

Figure 3.2 shows a comparison between the theoretical approach and experiment (red crosses)

carried out with a 1.4mm (left) and 1.1mm (right) large square diaphragm both about 10µm thick.

The static diaphragm deflection was varied by an adjustment of the back pressure in the cavity

between 0 and -900mbar. The pressure and the temperature in the test cell were kept constant

during the measurement. The red dashed line represents the resonance frequency as function of the

pressure load acting on the diaphragm determined with the above mentioned approach and the

dotted black line gives its derivative in respect to the pressure load (i.e. pressure sensitivity). Due to

the uncertainty in the value of diaphragm thickness of the individual sensor dies the particular

thickness for calculation was adjusted to match the resonance frequency at zero static deflection

based on theory and experiment. As is obvious, eq. 3.1 strongly overestimates the pressure

sensitivity determined by the experiment and under the aforementioned assumption of thickness

setting. Therefore eq. 3.1 was adjusted by a linear term inside the square root and the coefficients

are determined using the best correlation with all conducted measurements, eq. 3.4. Furthermore the

new approach rather underestimates the pressure sensitivity compared with the old one and

therefore seems to be more suitable to find the right resonator design for the present application.

( ) ( )2

95.021.010

+−==

t

d

t

ddfdf resres (3.4)

Fig. 3.2 Resonance frequency as function of static diaphragm deflection (diff. pressure across diaphragm) for a 10µm

thick silicon diaphragm with an edge length of 1.4mm (left) and 1.1mm (right). Red crosses: experiment, red solid:

eq. 3.1, red dashed: eq. 3.4. black dotted: pressure sensitivity derived from eq. 3.1 and black dash-dot derived from

eq. 3.4

4 Design considerations

The criteria regarding the sensor design partially followed from observations made during the above

mentioned experimental tests. The selection of the appropriate sensor design is based on the

validated approach given by eq. 3.2, 3.3 and 3.4.

Operating frequency range

The sensing devices are designed to operate in a frequency range between 50 and 80 kHz. At higher

frequencies the attenuation of ultrasound which increases with the frequency squared becomes too

high and thus excitation over a longer distance difficult. Furthermore the pressure sensitivity of a


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plate resonator with chip dimensions suitable for the present application is maximized in this

frequency range. Working at frequencies considerably below 50 kHz is not suitable because

environmental vibrations could influence the measurement or interaction between sound and the

boundary layer (BL) or the resonator and BL could alter the flow characteristics. Scenarios like an

interaction with the Tollmien-Schlichting waves in a laminar BL and thus an influence on the

transition between laminar and turbulent BL, or an alteration of the kinetic energy (larger eddies) of

a turbulent BL are conceivable. The latter interaction scenario could have an influence on the

friction drag or separation behavior of the BL. The main carriers of kinetic energy in a turbulent BL

are the larger eddies with characteristic frequencies far below 10 kHz (at wind speeds around

60m/s) (Blackwelder et al. 1976 and Schlichting et al. 2000). Thus an interaction with ultrasound is

not plausible. All studies where active flow control using acoustic sound or oscillating wall

structures was investigated, were carried out in the lower audible frequency range (Yarusevych et

al. 2006, Nishri et al. 1998, Ahuja et al. 1984, Greenblatt et al. 2000 and Sinha 2001).

Pressure sensitivity

Despite the fact that pressure sensors with a large sensing area (i.e. diaphragm) are generally more

sensitive to pressure variations due to the larger static diaphragm deflection, the measurement

sensitivity of a plate resonator is also influenced by its damping characteristics. The main damping

source of a plate oscillator results from acoustic radiation in air (Stemme 2001) which is dependent

on the diaphragm area squared, the vibrating frequency and the air density (Kinsler et al. 1982). The

influence of the two former parameters on the acoustic damping is shown in Figure 4.1. Basically

the damping is reduced (i.e. the pressure sensitivity increased) with decreasing diaphragm area. This

gain in measurement performance is reduced if the device operates in a frequency range below 100

kHz where the influence of the frequency on the acoustic damping is high and where the damping is

reduced with decreasing frequency (i.e. increasing diaphragm area). Figure 4.2 shows the

experimentally determined damping characteristics of a 1.1mm (blue circles) and 1.4mm (red star)

large diaphragm, expressed as a quality factor which is inversely related to the damping of an

oscillator (Q = fres/∆f3dB) (Stemme 2001).

Fig. 4.1 Damping fraction of the acoustic radiation as

function of the diaphragm size (R radius of a circular

diaphragm) and operating frequency

Fig. 4.2 Quality factor as function of the static diaphragm

deflection and area. 1.1mm (blue circles, 1.4mm (red

stars).

It is obvious that despite the higher resonance frequency (see Fig. 3.2 (right)) the smaller diaphragm

leads to higher Q-factors, i.e. lower damping compared with the larger device. The weak decay of

the Q-factor with decreasing pressure in the cavity (increasing static deflection or resonance

frequency) may be interpreted as an influence of the frequency on the acoustic damping.

Further restrictions toward an increasing diaphragm size are the required spatial resolution of 5mm

and a device size which should not modify the model geometry (contour). Sufficiently high pressure


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sensitivity and the capability of being acoustically excited and optically interrogated give the

limiting minimum size.

In general the maximum pressure sensitivity becomes higher if the diaphragm thickness decreases.

This is not necessarily true if the overall pressure sensitivity averaged over the required

measurement range is considered. For a very thin diaphragm the pressure sensitivity decays rapidly

after passing its maximum and may fall below the sensitivity of a thicker diaphragm. Furthermore

with a thinner diaphragm the fragility increases and a patterning of the upper surface (for optical

reasons as we will see later) becomes more difficult or unfeasible.

To be able to perform signed static pressure measurements the diaphragm must have a certain initial

static deflection which depends on the particular, signed measurement range (e.g. -6 kPa to 4 kPa).

The initial deflection is adjusted by setting the right pressure in the hermetically sealed cavity below

the diaphragm during the fabrication process.

A pre-stressed operation of the resonator is also favorable if its measurement performance is

considered. Figure 4.3 shows amplitude spectra of a 1.4 and 1.1mm large plate resonator for an

increasing static deflection, i.e. decreasing pressure in the cavity from 0 to -600mbar (-900mbar for

the right Figure) (from left to right). The shape of the individual amplitude spectrum gives certain

information about the oscillation behavior of the resonator. One may see that the resonance peaks at

zero or low pre-stressing are high and narrow and lead therefore to quality factors higher 100. If one

considers the larger diaphragm a sudden degradation of the oscillation behavior occurs at relatively

small static deflection (Fig. 4.3 left). The high drop of the resonance peak disappears again with

further increasing the pre-stressing of the diaphragm. The oscillation behavior improves again and

remains almost constant. Only a slight decay and broadening of the resonance peaks with increasing

static deflection remains visible. It probably is caused by the stiffening of the diaphragm together

with a constant excitation power and the broadening from the increasing resonance frequency (i.e.

increasing acoustic damping). The sudden drop in oscillation behavior could be located in the

vicinity of the maximum pressure sensitivity. As can be seen in figure 3.2 (right) this point is not

entirely reached and probably therefore no sudden drop of the resonance peak occurs in the case of

the smaller diaphragm (Fig. 4.3 right). That could confirm the assumption of different stress

conditions in the diaphragm before and after the point of maximum pressure sensitivity.

Temperature cross-sensitivity

An important requirement for the design is to keep the temperature cross-sensitivity far below the

measurement resolution of the system to avoid the need for correction of the pressure data. A single

material sensor design (single crystal silicon) with no electronics or intermediate oxide and nitride

Fig. 4.3 Characteristics of amplitude spectra with increasing static diaphragm deflection (from left to right) for a

large sensor with 1.4mm, Pcav: 0- -600mbar (left fig.) and small sensor with 1.1mm large diaphragm,

Pcav: 0- -900mbar (right fig.)


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layers together with a proper choice of the embedding material should minimize stresses due to

thermal expansion coefficient mismatch.

Fig. 4.4 Temperature cross-sensitivity of the large

sensor (1.4mm diaphragm edge length) embedded in

soft epoxy for a cavity pressure between 0 and

-800mbar (increasing static diaphragm deflection). A

constant sensitivity in the two given temperature ranges

(21 and 25ºC, 25 and 31ºC) is assumed

Fig. 4.5 Change in resonance frequency of the larger

sensor due to varying pressure in the test cell (i.e. air

density) between -500 and 300mbar. The static

diaphragm deflection of the sensor was kept constant

during the measurement. Solid line represents the

analytical approach (Di Gioanni 1982)

The experimental investigation of the influence of different embedding materials (silicon, rigid and

soft epoxy) on the temperature sensitivity of the sensing device did not provide conclusive results,

probably because of an unsuitable sensor fitting within the sample holder for such kind of

investigation. Further work on this subject will follow using the actual devices with a closed cavity.

Despite that an interesting observation could be made, namely a decreasing temperature cross-

sensitivity with an increasing static deflection of the diaphragm (i.e. an increasing pre-stressing)

(Fig. 4.4). The values in figure 4.4 are based on the assumption of constant temperature sensitivity

in the range between 21 and 25ºC and 25 and 31ºC and also take into account the uncertainty in

resonance frequency determination (error bars). The data were recorded for a large sensor

(diaphragm edge length 1.4mm) embedded in soft epoxy and cured at 35ºC for several hours.

The influence of variations of the surrounding gas density due to a change in temperature on the

vibration behavior is analytically described in literature (Di Giovanni 1982) and was also verified

by experimental tests (Fig. 4.5). Figure 4.5 shows the change in resonance frequency with

increasing pressure (from -500 to 300 mbar) and thus density within the test cell and compares it

with the analytical approach (solid line). During all measurements a slight pre-stressing of the

diaphragm (diff. pressure 25mbar) was preserved by adjusting the pressure in the cavity. If one

assumes that the air density changes due to temperature variations then the resulting cross-

sensitivity is small and should not exceed 4Hz/K at fres of about 70 kHz (compare temperature

sensitivity due to packaging, Fig. 4.4).

Dimensioning of the diaphragm

To evaluate the optimal size and thickness of the diaphragm the pressure sensitivity characteristics

for different diaphragm dimensions (edge length 1 to 2mm and thickness 3 to 10µm) was computed

using eq. 3.2 to 3.4. Subsequently the data were analyzed so as to fulfill following requirements:

(1) Maximum operating frequency range 50 to 80 kHz

(2) Minimum measurement range 10 kPa

(3) Minimum diaphragm pre-stressing corresponds to the point where maximum pressure

sensitivity occurs


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Fig. 4.6 Mean, max. and min. pressure sensitivity and operating

frequency range for different diaphragm dimensions identified

with an evaluation procedure

Fig. 4.7 Retro-reflecting characteristic of square

grooves with perpendicular side walls and

different aspect ratios (AR), β minimum incident

angle for retro-reflection, x usable part of the

bottom surface for retro-reflection and at a certain

incident angle α

Figure 4.6 shows the pressure sensitivity of all designs which fulfill the above mentioned criteria.

The markers represent the mean (average over measurement range) and the bar the maximum and

minimum sensitivity. The maker type indicates the corresponding diaphragm thickness and the

different colors the operating frequency range.

If one considers the advantage of smaller but thicker diaphragms (sensor size, acoustic damping and

mechanical strength) a plate resonator with an edge length of 1.4mm and a thickness between 5µm

and 6µm working in a frequency range between 50kHz and 80kHz was identified as the most

suitable configuration. Based on the assumption that the minimum sensitivity limits the

measurement performance (i.e. ≈ 2Hz/Pa), the resonance frequency has to be measured with an

uncertainty of 40Hz to meet the required resolution of 0.2% FS.

Adaptation of the sensor design for an improved optical interrogation

The main advantage of an imaging measurement technique is the possibility to interrogate many

sensors at the ‘same’ time. That requires an interrogation capability without the need for a precise

alignment of sensor and detector which is mainly fulfilled by frequency instead of displacement

measurement in the present application. Additionally the top surface of the diaphragm which

modulates the incoming light must have the capability to direct as much as light as possible to the

detector.

One possibility to obtain such an optical characteristic is to artificially roughen the top surface to

create a certain degree a diffusive reflectance. Because of the thin sensor diaphragm (t = 5-6µm) the

roughness height is limited to about the wave length of the incoming light (λ ≈ 700nm). Thus, only

weak diffusive reflectance may be expected. Besides the drawback of a diffusive surface losing

light due to hemispheric scatter, a further difficulty could arise during the manufacture of such a

rough surface on a thin diaphragm with common micro fabrication processes.

The most suitable sensor top surface would feature retro-reflecting properties, which are usually

obtained with a surface structure being similar to an array of corner prisms. As described in

literature there are ways to fabricate corner prism arrays with an element size down to 20µm on

silicon substrates (Neudeck et al. 1996 and Reimer et al. 1999). The drawbacks are the difficulty to


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further scale down the structures to suitable dimensions for the present application (1 to 2µm) and

special and unusual fabrication processes like the selective epitaxial growth technique or grey tone

lithograph.

With common wet and dry etching processes triple prisms are not producible on silicon substrates

but structures with partially retro-reflecting properties can be fabricated. Such structures are tiny 45º

V-grooves (Fig. 2.2) or square grooves with perpendicular side walls with respect to the top surface

(Fig. 4.7) both with an edge length between 1 and 2µm and a depth of 1µm. The former structures

can be manufactured by wet anisotropic etching using a surfactant added TMAH solution and an

alignment of the groove edge with the ⟨100⟩ direction (45º to the wafer flat) on a {100}-silicon

wafer (Sekimura 1999). The main drawback is that only incoming light parallel to the groove edges

is reflected back to the light source. Square grooves are simply fabricated using dry etching

processes which allows for highly anisotropic etching and thus perpendicular walls to the top

surface. It is further conceivable to smooth the side walls and the bottom surface performing a short

oxidation and subsequent stripping of the oxide in a buffered HF etchant. Though square structures

have a triple prism in each of their corners the retro-reflecting property works only for a certain

incident angular range which is dependent on the aspect ratio of the structure (Fig. 4.7). Shallow

structures works better for flat and deep ones for steep irradiation and thus a mixed array with

structures of different aspect ratios should be preferable. Nevertheless a blind spot range remains.

Due to structure dimensions near the wavelength of the interrogation light diffraction on the

patterned surface should appear which would hinder the interrogation process. Therefore the

individual grooves are pseudo-randomly distributed over a regular grid with a pitch of 2µm. As can

be shown the power spectral density spectrum of such a grid is independent of spatial frequency, i.e.

white-noise-like (Yashchuk et al. 2007).

We intend to develop fabrication processes for a rough, diffusive surface and for the two above

mentioned (partially) retro-reflective groove patterns and compare them regarding their capability

to be detected by coherent illumination from different observation points.

5 Sensor fabrication

As mentioned before the targeted low temperature cross-sensitivity of a highly sensitive pressure

sensor requires a sensing device made out of one material to minimize stresses due to thermal

expansion coefficient mismatch. Thus the common fabrication technique for pressure sensors where

the sensing diaphragm is manufactured by wet anisotropic etching from the back side of the wafer

with a precise etch stop at the highly doped area represented by the diaphragm is not applicable for

the present application.

The plate resonator with a hermetically sealed cavity is fabricated using low temperature direct

bonding of a silicon-on-insulator (SOI) and a silicon 4 inch wafer. The thin silicon wafer (200 µm)

forms the support for the diaphragm and features on its front side the 30µm deep pressure sensor

cavities. They are manufactured by common photolithography processes and wet anisotropic

etching in KOH (Fig. 5.1a and b). The device layer of the SOI wafer with a total thickness variation

(TTV) of less than 0.5µm forms the 6µm thick sensing diaphragm. Its front side is structured with a

retro-reflecting pattern (45º V- or 90º square grooves) performing e-beam lithography and

dependent on the groove geometry wet (surfactant added TMAH solution) or dry anisotropic

etching processes (Fig. 5.1c and d). E-beam lithography is necessary to obtain square structures

with preferably small rounding of their corners. The patterned surface is then bonded (low

temperature fusion bonding) with a further silicon support wafer (200µm) (Fig. 5.1e). For its later

removal by KOH etching the pattern on the SOI wafer is masked with an oxide layer (0.5µm) from

being etched as well. The diaphragm is released from the remaining SOI wafer using KOH (silicon

handle layer) and BHF (oxide sacrificial layer) wet etching (Fig. 5.1f). Low temperature fusion


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bonding is also used for joining the diaphragm (supported by wafer III) and its bottom part

including the cavity. This time the boding has to be performed in a precisely controlled nitrogen

environment to adjust the low pressure of -200mbar in the cavity and therefore the necessary initial

static deflection of the diaphragm (Fig. 5.1g). During the final step the support wafer (wafer III) and

the oxide layer are removed in a KOH and BHF bath respectively (Fig. 5.1h). a)

d)

b)

e)

c)

f)

g)

h)

Fig. 5.1 Fabrication of the sensing device

6 Conclusion

A sensor concept for wall pressure measurements in low speed wind tunnel applications is

presented. The sensor is designed for full field acquisition of a pressure distribution with a discrete

spatial resolution in the millimeter range and a temporal resolution around 1Hz. An integration of

the sensor unit directly on the surface of the model and a wireless interrogation should reduce the

handling and installation costs significantly.

The sensor concept is based on an array of discrete vibrating plate micro-resonators using their shift

of the resonance frequency as a measure for the applied pressure. These resonant structures have to

be excited wirelessly by ultrasound and their resonance frequency is optically interrogated from a

distance of about 2 meters.

The sensor design was optimized with regard to highest pressure sensitivity and measurement

performance and was carried out using a semi-empirical approach which was validated by

experimental tests. The most suitable resonator design has a 5 to 6µm thick diaphragm with an edge

length of 1.4mm. It is expected that the sensor is working in a frequency range between 50 and

80kHz and covers thereby a pressure range of around 10kPa. A certain initial static diaphragm

deflection allows for signed pressure measurement and improves also the measurement

performance. A minimum pressure sensitivity of about 2Hz/Pa and hence an uncertainty in


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resonance frequency measurement of 40Hz is expected to meet the required resolution of 0.2% FS.

A type of retro-reflecting pattern on the top surface of the diaphragm should simplify the optical

interrogation of the sensor array.

In order to reduce the temperature cross-sensitivity the whole sensor is made out of one material,

namely single crystal silicon. The fabrication relies on common photolithography and anisotropic

etching processes and also involves two low temperature direct fusion bonding processes.

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