airborne ultrasonic electrostatic transducers with conductive grooved backplates: tailoring their...

Ultrasonics 37 (1999) 493–503www.elsevier.nl/locate/ultras

Airborne ultrasonic electrostatic transducers with conductive groovedbackplates: tailoring their centre frequency, sensitivity and bandwidth

Luis Pizarro a,*, Dominique Certon a, Marc Lethiecq a, Bernard Hosten ba LUSSI/GIP ULTRASONS, 2 bis bd. Tonnelle, BP 3223, 37032 Tours Cedex, France

b LMP, Universite de Bordeaux I, 351 Cours de la Liberation, 33405 Talence Cedex, France

Received 2 February 1999; received in revised form 21 July 1999

Abstract

Several electrostatic transducers with metallic backplates have been manufactured. Some backplates have been subjected tocoarse polishing, whereas others were first polished and then rectangular grooves were machined. Experimental investigationshave been performed in order to study the influence of the groove widths, their layout and the value of DC bias on the transducerbehaviour. These results are compared with theoretical predictions using the Helmholtz resonator model. The equation of thevibration of a membrane submitted to a mechanical stress, but neglecting the effects of its elasticity, was also compared withexperimental results. This shows that the Helmholtz resonator model is adapted to the behaviour of the polished backplates,whereas the membrane model explains the vibration over the grooves. By choosing the groove widths and layout (i.e. polished-to grooved-surface ratio), the centre frequency, the bandwidth and the sensitivity of the transducer can be optimised for a givenapplication. © 1999 Elsevier Science B.V. All rights reserved.

Keywords: Air-coupled ultrasound; Capacitance transducer; Electrostatic transducer

1. Introduction as a few micrometres. The behaviour and performanceof the transducers depends on the type of backplate. If

Electrostatic transducers have been developed in the the backplate surface introduces a layer of air betweenlast few years for the generation and detection of the conductor and the membrane (i.e. coarsely polishedultrasonic waves in air. Their main advantage is a wide or some micromachined surfaces), then the transducerbandwidth and a sufficiently high sensitivity for non- behaves as a capacitor microphone [6,8,9,11]. In thecontact applications such as non-destructive evaluation case of larger-scale micromachining, the membrane can[1–5]. Several manufacturing techniques have been stay in contact with the surface between the cavities,reported in the literature, and these can be classified and then a drum-type behaviour of the membrane overinto two groups. The first includes all the devices using the cavities is observed [20,21].a metallic backplate, the second those based on doped The transducers manufactured here are based on asilicon. In the first group, some have used a coarse backplate made of tempered steel that is first polishedpolishing technique to obtain the desired surface rough- and then into which rectangular grooves are micro-ness of the backplate [6–9], whereas others have machined. A thin plastic membrane with the externalmachined V-shaped grooves [10–16 ], or introduced cir- face metallised is then deposited on the backplate. Acular holes by laser ablation [17–19] or by photolitho- DC bias is finally applied between the backplate andgraphy [20]. In the second group, micromachining the front face of the membrane.technologies used in the microelectronics industries are In this way, as an electrical signal is applied to theused, mainly photolithography [21–27]. Such technol- transducer a membrane displacement is produced andogies permit the manufacture of a periodic distribution converted into an airborne ultrasonic wave. Conversely,of circular cavities with diameters and depths as small as an ultrasonic wave reaches the transducer, it produces

a membrane displacement, which is transformed into anelectrical signal.* Corresponding author.

E-mail address: [email protected] (L. Pizarro) The aim of this work is to study the effect of the

0041-624X/99/$ – see front matter © 1999 Elsevier Science B.V. All rights reserved.PII: S0041-624X ( 99 ) 00033-5

494 L. Pizarro et al. / Ultrasonics 37 (1999) 493–503

width of the grooves, of their layout and of the value 2.2. Mechanical displacement measurementof the DC bias on the centre frequency, the bandwidth

The amplitude of mechanical displacement atand the sensitivity of transducers.different locations on the membrane surface of eachMeasurements of the mechanical displacement of thetransducer was measured using a laser interferometermembrane and the pulse electro-acoustic response have(SH-140 Nd:Yag, 100 mW from BM Industries) with abeen performed. The transducer design and the experi-sensitivity of 10 mV A−1, a 20 kHz–5 MHz bandwidthmental set-up will first be described. Then, the theoreticaland a focal beam-width of 20 mm (Fig. 2). A functionbackground will be recalled and predicted results willgenerator (Hewlett Packard HP3314A) was used tobe compared with experiments and discussed.drive the transducer providing a −16 V pulse with~1 MHz bandwidth at −6 dB. The signal issued fromthe laser interferometer was digitised and stored in amicrocomputer. Finally, by computing the fast Fouriertransform (FFT), the frequency response was obtained.2. Materials and methods

2.1. Transducer design and fabrication 2.3. Electro-acoustic response

The electro-acoustic response of the different trans-The transducers consist of a disc-shaped temperedducers was measured by generating and receiving ultra-steel backplate with a diameter of 18 mm and a thicknesssonic waves by two identical transducers facing eachof 4.6 mm. A polishing procedure, using a diamondother. A pulse generator (1035PR from Accu-Tron) waspaste (grain size smaller than 1 mm), was applied toused to drive the transmitter providing a −200 V wide-backplates to reduce the roughness of the surface inband pulse. The pulse repetition rate was always 200 Hz.contact with the membrane. Then, all polished back-The received signal was amplified by a charge amplifierplates were micromachined with different groove(model CA6/C from Cooknell ). This signal was thenarrangements. Two groove widths were considered, withdigitised and displayed on a digital oscilloscopedifferent groove pitches. Backplates with grooves of(HP54504A from Hewlett Packard), and stored in awidth 200 mm have pitches of 400 and 300 mm; back-microcomputer. The frequency response was obtainedplates with grooves of width 100 mm have pitches ofby computing the FFT. Both transducers had the same300, 200 and 150 mm. The groove depth is alwaysDC bias (Fig. 3).100 mm. An example of layout is shown in Fig. 1. The

The distance between transducers was maintained atmembrane is a Mylar film of thickness 2.5 mm that was80 mm in order to avoid large diffraction effects onpretensioned and bonded to a metallic ring.measurements.

3. Theoretical background

This section, devoted to the study of mechanicalbehaviour of the membrane in metallic substrate trans-ducers, is a first approach towards a complete transducermodel in such configurations. Here however, only resultsof simple mechanical theory will be presented in viewof the interpretation of experimental results. Differenttheoretical approaches will be considered, according tothe type of backplate, i.e. rough or grooved.

3.1. Rough backplates

This analysis is based on the work by Carr andWykes [7–9]. The basic hypothesis is that the roughnessof the substrate surface creates a large number ofmicroscopic air bubbles trapped under the membraneand that they can be considered as a single layer of airwith a thickness dag. The transducer is thus assumed toFig. 1. SEM micrograph of a grooved backplate (groove width 100 mm,

pitch 300 mm). be a simple two-layer system laying on a semi-infinite

495L. Pizarro et al. / Ultrasonics 37 (1999) 493–503

Fig. 2. Mechanical displacement measurement set-up.

Fig. 3. Electro-acoustic response measurement set-up.

substrate (Fig. 4). Such a transducer can then be ana-lysed in a similar way as for condenser microphones.

When the membrane is submitted to an excitationvoltage v(t), the electrostatic forces are opposed to twomechanical forces: the tension T with which the mem-brane is maintained at its border and the pressure insidethe air layer trapped between the membrane and thesubstrate. The tension governs the resonances of the

Fig. 4. Schematic view of rough backplate transducer.membrane, similar to those of a drum [28], at frequency


f calculated using

J0A2pfRS rS

T B=0, (1)

where J0(x) is the zero-order Bessel function, rS themass per unit surface of the membrane and R the radiusof the transducer.

Considering an experimental tension evaluatedaround 20 N m−1 and a radius of 9 mm, these reso-nances are of the order of a few kilohertz, which ismuch lower than the useful bandwidth of our transduc-ers. The effect of the tension T can thus be neglected.

The pressure in the air layer can be calculated usingHelmholtz resonator theory [29]. For a mechanical

Fig. 5. Amplitude of membrane displacement versus frequency for adisplacement y(t) of the membrane, this pressure PAir rough backplate transducer.can be calculated using

ment as a function of frequency. One can notice that inPAir=Pagy(t) with Pag=

cP0

dag, (2) the useful bandwidth of our transducer the frequency

response is almost flat, since the resonance only appearsaround 1.4 to 1.5 MHz. Since losses have been neglected,where c is the adiabatic constant of air and P0 thethe amplitude of resonance goes to infinity. In practice,atmospheric pressure.losses will lead to a finite amplitude, which can beNewton’s law applied to the displacement of themodelled by a resistor in the transducer equivalentmembrane leads to the behaviour equationcircuit [11,20,21,23,25].

rS∂2y

∂t2+Pagy(t)=−pe(t), (3)

3.2. Grooved backplates

where pe(t) is the pressure due to electrostatic forces.For such substrates we will first study the behaviour

For a harmonic excitation voltage [ pe(t)=Pe ejvt], oneof the membrane over a simple groove and then extend

can easily obtain the solutionthe results to the entire surface of the transducer, sincethe grooves have a periodic spacing [7,9,20,21,27]. From

y(t)=Pe

rSv2−Pagejvt, (4) a theoretical point of view, our grooved transducers

have a similar behaviour to those manufactured by Carrand the resonance frequency of the system is and Wykes [7]. Their devices show a centre frequency

dependence with the membrane properties (membranewidth and mechanical tension) but not with the depthf

0=

c0

2pS rAirrMhdag

, (5)of the grooves. The transducer behaviour is welldescribed by a simple membrane resonator. This is not

where rAir is the density of air, c0 the speed of sound in the case for the devices manufactured by Hietanen et al.air, rM the density of the membrane and h its thickness. [14], where the centre frequency corresponds to theAs the roughness of the substrate is decreased, the resonance of the air trapped in the cavity. In such aequivalent air layer thickness is reduced and the reso- case, the Helmholtz resonator model is well adapted asnance frequency of the device increases. for rough backplate transducers.

If the transducer is considered as a plane capacitor, The system is schematically represented in Fig. 6: athe expression of the electrostatic pressure Pe [21] membrane of thickness h covers a groove of width a,depends on the square of the applied voltage: the depth dag and length L with L&a.polarisation voltage VP and the amplitude of the har- We assume infinite elasticity for the membrane, thenmonic excitation Vac; by making VP&Vac, we obtain the only mechanical force is due to the tension T (a

precise measurement of T is difficult, but it was evaluatedPe=

e0VPVac

(dag+h)(dag+h/eM), (6) experimentally at around 25 N m−1). With the depth of

the groove (i.e. the thickness of the air cavity) beingaround 100 mm, no significant pressure variation will bewhere e0 and eM are the permittivity constant of the

vacuum and the relative permittivity of the membrane observed and the resonance frequency of the cavity willbe well below our useful bandwidth. The displacementrespectively.

Fig. 5 shows the amplitude of the membrane displace- of the membrane on the edges of the groove is assumed


Table 1Calculated resonance frequencies for a membrane thickness of 2.5 mm

Groove width (mm) Membrane tension (N m−1)T=20f0 (kHz) T=30f0 (kHz)

200 189 232100 378 463

Fig. 6. Schematic view of grooved backplate transducer.

to be zero. This hypothesis is probably not strictlyverified in practice, and the effective displacement onthe edges should be limited to the roughness of thesubstrate between two grooves. Finally, with the lengthof the groove being much larger than its width, theproblem can be written in a single dimension, i.e. acrossthe width. Fig. 7. Averaged membrane displacement over a groove: groove width

100 mm, Vp=100 V, Vac=1 V and T=22 N m−1.The displacement of the membrane y(x, t) when it issubmitted to an electrostatic pressure pe(t) is then gov-erned by the differential equation tude along its width can also be calculated using

∂2y(x, t)

∂t2=

T

rMh

∂2y(x, t)

∂x2−

pe(t)

rMh. (7) j:=YPC1−

tan(ka/2)

ka/2 D. (10)

This mean displacement amplitude is proportional toConsidering zero displacement boundary conditionsthe amplitude of the dynamic electrostatic pressure[ y(0, t)=y(a, t)=0] and only harmonic solutionsthrough the term YP, whose frequency dependence fol-[ pe(t)=Pe sin(vt)], the amplitude of the membrane dis-lows a 1/v2 law.placement is given by

Fig. 7 shows the mean displacement amplitude as afunction of frequency for a groove width of 100 mm, aj(x)=YPG1−cos(kx)−C1−cos(ka)

sin(ka) Dsin(kx)H, (8)polarisation voltage of 100 V, an amplitude of excitationvoltage of 1 V and a tension of 22 N m−1. A resonance

with frequency in the useful bandwidth of transducers isobserved (around 400 kHz).

YP=Pe

rMhv2, k=

v

c, c=S T

rMh.

4. Results and discussionThe frequency associated with the n-order resonance isthen

4.1. Mechanical displacement measurement

fn=

n

2aS T

rMh. (9) The results presented here concern the impulse

response of the membrane displacement obtained withthe experimental set-up and procedure presented inConsidering the system as a plane capacitor, the

amplitude of the dynamic electrostatic pressure Pe is Section 2. For each transducer, we have measured thisdisplacement as a function of the polarisation voltage.given by Eq. (6) as for a rough backplate.

Table 1 gives the resonance frequencies of the system Figs. 8 and 9 show the time domain impulse responseand frequency response for a transducer with 100 mmfor two groove widths (100 and 200 mm) and two

membrane tensions (20 and 30 N m−1) for a membrane wide grooves and a pitch of 150 mm in the centre of agroove and at mid distance between two adjacentdensity of 1400 kg m−3.

The mean value of the membrane displacement ampli- grooves (on a polished area). Fig. 10 shows the results


Fig. 9. Experimental membrane displacement of a micro-machinedFig. 8. Experimental membrane displacement of a micromachinedtransducer (groove width 100 mm, pitch 150 mm), at the centre of atransducer (groove width 100 mm, pitch 150 mm), at the centre of arough surface, for 50, 100 and 150 V DC biases: (top) time responsegroove, for 50, 100 and 150 V DC biases: (top) time response andand (bottom) frequency response normalised by the excitation fre-(bottom) frequency response normalised by the excitation frequencyquency response.response.

for a rough backplate transducer at different locations reproduced on the membrane surface. The frequencyresponses show that the displacement has a moderateon the surface. In order to be able to give spectrum

amplitudes in angstroms per volt, a deconvolution of bandwidth (around 20%) and a centre frequency around500 kHz, which corresponds to the first calculated reso-the measurements by the excitation voltage was

performed. nance over a groove. The position of this resonanceallows an estimation of the mechanical tension of theFor the grooved backplates, a high displacement

amplitude is observed in comparison with that of rough membrane: a value around 35 N m−1 is found. Betweenthe grooves the spectra have maximum amplitudesbackplates. Over the centre of a groove, a maximum

displacement of 150 A is measured, compared with a around 20 A V−1 at frequencies around 500, 900 and1300 kHz. The first frequency can be associated withmaximum of 55 A on a rough backplate. The maximum

amplitude over the centre of a groove is obtained when the resonance of the grooves. The amplitude at thisfrequency, when low polarisation voltages are applied,the polarisation voltage is 150 V, whereas the optimum

voltage is around 50 V over the polished parts of the cannot be neglected compared with the amplitude mea-sured over a groove. Also, the vibration over the groovesbackplate. The results show that, over the grooves, the

amplitude of displacement increases with polarisation and ungrooved areas are coupled and both contributeto the energy radiated in air. The other resonance peaks,voltage, which is in agreement with Eq. (8) of the

theoretical study. However, over the polished parts of at high frequencies, can be associated with smallHelmholtz resonators, i.e. air bubbles trapped betweenthe backplates, the effects of polarisation voltage are the

opposite: an increase of voltage Vp tends to press the the membrane and the polished backplate. Indeed, thesehigh frequency resonances disappear when the polarisa-membrane on the backplate, thus hindering its vibration

when submitted to an impulse voltage. This effect can tion voltage is increased and the trapped air is evacuatedas the membrane is pressed on the backplate.be observed directly, since, when the voltage Vp is

increased, the microstructure of the polished surface is For the rough backplate transducers (Fig. 10), a


Fig. 11. Scan of frequency response of the membrane displacement ofa micromachined transducer (groove width 200 mm, pitch 300 mm andDC bias 100 V ).

be clearly identified and the polished areas betweengrooves are characterised by a very low displacementamplitude in the 100–900 kHz range. At the centre of

Fig. 10. Experimental membrane displacement of a rough backplate each groove, a resonance peak around 250 kHz, corre-transducer, at three different locations and for a DC bias of 100 V:

sponding to the calculated fundamental mode of cavity,(top) time response and (bottom) frequency response normalised byis observed. At higher frequencies, more complex ampli-the excitation frequency response.tude distributions appear. In particular, at 750 kHz thethird harmonic, black and white spots corresponding tonodes and antinodes of vibration, is observed. Moreshort time response is observed with, at certain locations,

a ring-down, which is probably associated with the surprising are the peaks observed around 500 kHz thatseem to have been generated. This could be explainedpresence of a trapped air bubble. When the measurement

point is scanned over the surface, large changes in both by a lack in symmetry of the boundary conditions onthe two edges of each groove. Finally, at around 1 MHz,the centre frequency and the ring-down are observed.

Such dispersions lead us to expect a large bandwidth many peaks appear both over the grooves and betweenthem. Such complex spectra are difficult to analyse;for rough backplate devices. Moreover, the measured

frequency responses at different locations are relatively however, one can imagine that the peaks are due totrapped air bubbles over the polished parts of theflat in the 100 kHz–1 MHz frequency range.backplate, which would behave as Helmholtz resonators.

4.2. Scan of frequency response4.3. Electro-acoustic response

In order to obtain a global view of the vibration ofthe transducer surface, the measurement point was This section presents the results of transmission-mode

electro-acoustic responses measured according to thescanned along a direction perpendicular to the grooveaxis. Then, for each position, the spectrum of the procedure described in Section 2.3. Figs. 12–16 concern

grooved backplates and Fig. 17 was obtained using adisplacement was calculated. Fig. 11 shows such a scanfor a transducer with a groove width of 200 mm and a rough backplate. For each case, an adjustment of the

damping on the impulse generator (i.e. value of resistorpitch of 300 mm. The horizontal axis is the frequency(kilohertz) and the vertical axis is the position of the in parallel with the transducer) was performed in order

to optimise the amplitude of the time response. In somemeasurement point (micrometres). The amplitudes ofthe spectra are represented in grey scale (white for cases, however, a saturation of the receiver was observed,

so a slight electrical mismatch was introduced.maximum, black for zero). The positions of grooves can


Fig. 14. Experimental electro-acoustic response of three microma-Fig. 12. Experimental electro-acoustic response of two micromachinedchined transducers with the same groove width (100 mm) and differenttransducers having the same polished- to grooved-surface ratio: (top)pitches (solid line: pitch 300 mm; dashed line: pitch 200 mm; dotted line:200 mm groove width; (bottom) 100 mm groove width.pitch 150 mm) for a DC bias of 50 V: (top) time response; (bottom)frequency response; a.u.: arbitrary units.

be defined by the fundamental resonance of the mem-brane over the groove. The centre frequency of a devicecan be tailored simply by changing the groove width.The bandwidths measured here are longer than thoseobtained using the laser probe. Such an increase inbandwidth could also be explained by the contributionof a large number of Helmholtz resonators due to airbubbles of different sizes trapped over the polished partsof the backplate.

Figs. 14–16 show the electro-acoustic responses as afunction of the grooved surface ratio tR (33%, 50% and66%) and for three different polarisation voltages (50 V,100 V and 150 V ). The groove width is always 100 mm.Fig. 13. Experimental electro-acoustic frequency response of two

micromachined transducers having the same polished- to grooved-sur- At 100 and 150 V, the transmitted energy increasesface ratio: (solid line: groove width 100 mm and pitch 200 mm; dashed with the grooved surface ratio. This shows that, asline: groove width 200 mm and pitch 400 mm; a.u.: arbitrary units). observed with the mechanical displacements in Figs. 8

and 9, the highest displacements are those over thegrooves. At frequencies much higher than the fundamen-Figs. 12 and 13 show the time and frequency

responses for two different groove widths (100 mm and tal resonance of the membrane over the groove thespectra are practically zero. Indeed, the transmitted200 mm), but with identical grooved surface ratios

(tR=50%). A relatively low bandwidth response is energy is low and the attenuation in air is high.With a polarisation voltage of 150 V most of theobserved, comparable to the displacement measure-

ments. The centre frequency of the system appears to energy is around 400 kHz, and for grooved surface


Fig. 16. Experimental electro-acoustic response of three microma-chined transducers with the same groove width (100 mm) and differentpitches (solid line: pitch 300 mm; dashed line: pitch 200 mm; dotted line:pitch 150 mm) for a DC bias of 150 V: (top) time response; (bottom)

Fig. 15. Experimental electro-acoustic response of three microma-frequency response; a.u.: arbitrary units.

chined transducers with the same groove width (100 mm) and differentpitches (solid line: pitch 300 mm; dashed line: pitch 200 mm; dotted line:pitch 150 mm) for a DC bias of 100 V: (top) time response; (bottom)frequency response; a.u.: arbitrary units.

last graph, Fig. 17, shows the transmission responses ofrough backplate transducers for polarisation voltages of50, 100 and 150 V. The amplitude of the transmittedratios of 33% and 50% some contribution is found

around 500 kHz and 1 MHz. This behaviour is in signals is lower than with the grooved backplates, andthus the sensitivity is lower. On the other hand, theagreement with the displacement measurements of

Fig. 9, where the polished parts of the substrate have bandwidth is much higher and values over 100% of –6 dB relative bandwidths are obtained. Such a result isbeen shown to give a significant contribution to the

vibration of the membrane. This can explain why, in in agreement with the displacement measurement ofFig. 11, where a relatively flat frequency response wasFig. 14, the effects of the grooved surface ratio on the

sensitivity of the system are contrary to those in Figs. 15 observed in the 100 kHz–1 MHz frequency range. Theresonance peak due to the air layer trapped under theand 16.

In Fig. 15–17, one can finally notice, with more or membrane, found to be around 1.6 MHz, does notcontribute to the transmitted signals because of theless amplitude, a peak around 100 kHz. This is probably

due to the resonance of the air cavity in the grooves attenuation in air between the two transducers.The optimum polarisation voltage is found to bethemselves. Indeed, Eq. (5) for a Helmholtz resonator

applied to a rectangular cavity of 100 mm depth gives a around 100 V. Here, two effects are in competition: first,as shown by Eq. (4), an increase of the polarisationresonance mode around 100 kHz. This peak was attenu-

ated by the demodulation electronics of the laser voltage should lead to an increase of the transmittedenergy; second, as the polarisation voltage increases, theprobe (and thus not observed with displacement

measurements). air bubbles are evacuated and the number of Helmholtzresonators on the transducer surface decreases, thusHere, it is less attenuated than the higher frequency

peaks due to the air gap between the transducers. The decreasing the active area.


is reached beyond which performances decrease becausethe membrane is then pressed against the backplate andthe air layer vanishes.

Such an optimum voltage is also observed for groovedbackplate transducers. For low polarisation voltages,the entire membrane surface (i.e. over and between thegrooves) shows significant vibrations. For high voltages,only the membrane over the grooves themselves contrib-utes to the vibrations. The devices with highest sensitivityare those with the highest grooved surface ratio. Inparticular, under high polarisation voltages, only thegrooved surface radiates acoustic energy.

Finally, the use of grooved backplate is shown toallow an adjustment of the centre frequency of a trans-ducer through the device of the groove width. Thebandwidth of such devices is, however, significantlylower than that of rough backplate devices. Future workwill concern the development of specific models in orderto predict the electro-acoustic behaviour of such trans-ducers, as well as design and experimental proceduresin order to control their performances.

References

[1] D.W. Schindel, D.S. Forsyth, D.A. Hutchins, A. Fahr, Air-cou-pled ultrasonic NDE of bonded aluminium lap joints, Ultrasonics35 (1) (1997) 1–6.

[2] D.W. Schindel, D.A. Hutchins, Applications of micromachinedcapacitance transducers in air-coupled ultrasonics and non-destructive evaluation, IEEE Trans. UFFC 42 (1) (1995) 51–58.

[3] D.W. Schindel, D.A. Hutchins, Through thickness characterisa-Fig. 17. Experimental electro-acoustic response of a rough transducer tion of solids by wideband air-coupled ultrasound, Ultrasonics 33for three DC biases (solid line: 50 V; dashed line: 100 V; dotted line: (1) (1995) 11–17.150 V ): (top) time response; (bottom) frequency response; a.u.: arbi- [4] W.A. Grandia, C.M. Fortunko, NDE applications of air-coupledtrary units. ultrasonic transducers, in: IEEE Ultrasonics Symposium Proceed-

ings Seattle, USA vol. 1 (1995) 697–709.[5] B. Hosten, D.A. Hutchins, D.W. Schindel, Measurements of elas-5. Conclusion

tic constants in composite materials using air-coupled ultrasonicbulk waves, J. Acoust. Soc. Am. 99 (4) (1996) 2116–2123.

Electrostatic airborne ultrasonic transducers with [6 ] D.W. Schindel, D.A. Hutchins, L. Zou, M. Sayer, Capacitancedevices for the generation of air-borne ultrasonic fields, in: IEEEmetallic backplates have been designed, manufacturedUltrasonic Symposium Proceedings Tucson, USA vol. 2 (1992)and characterised. Experimental results were also com-843–846.

pared with simulations. The influence of the backplate [7] H. Carr, C. Wykes, Diagnostic measurements in capacitive trans-surface roughness on transducer performance was ducers, Ultrasonics 31 (1) (1993) 13–20.

[8] C. Wykes, H. Carr, W.S.H. Munro, F. Nagi, S. Pomeroy, M.studied in two particular configurations: rough back-Rafiq, Advances in air-borne ultrasonic sensors, in: Ultrasonicsplates and grooved backplates. For the latter, differentInternational Conference Proceedings Vienna, Austria (1993)groove widths and grooved surface ratios were 667–670.

investigated. [9] H. Carr, Electrostatic transducers for airborne ultrasonics, Thesis,For rough backplates, the experimental results show University of Nottingham, UK, 1989.

[10] J. Hietanen, P. Mattila, J. Stor-Pellinen, F. Tsuzuki, H. Vaataja,that the air trapped between the membrane and theFactors affecting the sensitivity of electrostatic ultrasonic trans-substrate behaves as a Helmholtz resonator whose reso-ducers, Meas. Sci. Technol. 4 (1993) 1138–1142.

nance frequency depends essentially on the mean thick- [11] P. Mattila, F. Tsuzuki, H. Vaataja, K. Sasaki, Electroacousticness of this layer. It has been observed that, below this model for electrostatic ultrasonic transducers with V-grooved

backplates, IEEE Trans. UFFC 42 (1) (1995) 1–7.resonance (which occurs between 1 and 2 MHz), the[12] J. Hietanen, J. Stor-Pellinen, M. Luukkala, A model for an elec-transducer exhibits a relatively flat frequency response,

trostatic ultrasonic transducer with a grooved backplate, Meas.which leads to the large bandwidth and high resolution Sci. Technol. 3 (1992) 1095–1097.measured on such devices. Their sensitivity is shown to [13] P. Mattila, J. Hietanen, Bandwidth control of an electrostatic

ultrasonic transducer, Sensors Actuators A: 45 (1994) 203–208.first increase with polarisation voltage, then an optimum


[14] J. Hietanen, J. Stor-Pellinen, M. Luukkala, P. Mattile, F. Tsuzuki, [21] M.I. Haller, B.T. Khuri-Yakub, A surface micromachined electro-static ultrasonic air transducer, IEEE Trans. UFFC 43 (1)K. Sasaki, A Helmholtz resonator model for an electrostatic ultra-

sonic air transducer with a V-grooved backplate, Sensors Actua- (1996) 1–6.[22] D.W. Schindel, D.A. Hutchins, L. Zou, M. Sayer, The design andtors A: 39 (1993) 129–132.

[15] W.S.H. Munro, C. Wykes, Arrays for airborne 100kHz ultra- characterization of micromachined air-coupled capacitance trans-ducers, IEEE Trans. UFFC 42 (1) (1995) 42–50.sound, Ultrasonics 32 (1) (1994) 57–64.

[16 ] G.J. Brown, D. Reilly, Ultrasonic tomographic imaging of solid [23] C. Thielemann, G.M. Sessler, Capacitive silicon sensors for ultra-sound, Acustica 83 (1997) 715–720.objects in air using an array of fan-shaped-beam electrostatic

transducers, Ultrasonics 34 (2–5) (1996) 111–115. [24] A.G. Bashford, D.A. Hutchins, D.W. Schindel, Radiated fieldsof an air-coupled ultrasonic capacitance transducer, Ultrasonics[17] A.G. Bashford, D.W. Schindel, D.A. Hutchins, Micromachined

ultrasonic capacitance transducers for immersion applications, 34 (2–5) (1996) 169–172.[25] K. Suzuki, K. Higuchi, H. Tanigawa, A silicon electrostatic ultra-IEEE Trans. UFFC 45 (2) (1998) 367–375.

[18] D.W. Schindel, D.A. Hutchins, Capacitance devices for the con- sonic transducer, IEEE Trans. UFFC 36 (6) (1989) 620–627.[26 ] P.C. Eccardt, K. Niederer, B. Fischer, Micromachined transduc-trolled generation of ultrasonic fields in liquids, in: IEEE Ultra-

sonics Symposium Proceedings Florida, USA vol. 1 (1991) ers for ultrasound applications, in: IEEE Ultrasonics SymposiumProceedings Toronto, Canada vol. 2 (1997) 1609–1618.301–304.

[19] A.G. Bashford, D.W. Schindel, D.A. Hutchins, Characteristics of [27] I. Ladabaum, X. Jin, H.T. Soh, A. Atalar, B.T. Khuri-Yakub,Surface micromachined capacitive ultrasonic transducers, IEEEultrasonic micromachined capacitance transducers in water, in:

IEEE Ultrasonics Symposium Proceedings San Antonio, USA Trans. UFFC 45 (3) (1998) 678–690.[28] P. Morse, K. Uno, Theoretical Acoustics, Princeton Universityvol. 2 (1996) 955–958.

[20] M.J. Anderson, J.A. Hill, C.M. Fortunko, N.S. Dogan, R.D. Press, 1986.[29] L. Kinsler, A. Frey, A. Coppens, J. Sanders, Fundamentals ofMoore, Broadband electrostatic transducers: modelling and

experiments, J. Acoust. Soc. Am. 97 (1) (1995) 262–272. Acoustics, 3rd edition, Wiley, 1982.

airborne ultrasonic electrostatic transducers with conductive grooved backplates: tailoring their...

Documents