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A Fabry–Pérot fiber-optic ultrasonic hydrophone for thesimultaneous measurement of temperature andacoustic pressure

Paul Morrisa�

Department of Medical Physics and Bioengineering, University College London, Gower Street, LondonWC1E 6BT, United Kingdom

Andrew HurrellPrecision Acoustics Ltd., Hampton Farm Business Park, Higher Bockhampton, Dorchester DT2 8QH,United Kingdom

Adam ShawQuality of Life Division, National Physical Laboratory, Hampton Road, Teddington, Middlesex TW11 0LW,United Kingdom

Edward Zhang and Paul BeardDepartment of Medical Physics and Bioengineering, University College London, Gower Street, LondonWC1E 6BT, United Kingdom

�Received 9 October 2008; revised 11 March 2009; accepted 16 March 2009�

A dual sensing fiber-optic hydrophone that can make simultaneous measurements of acousticpressure and temperature at the same location has been developed for characterizing ultrasoundfields and ultrasound-induced heating. The transduction mechanism is based on the detection ofacoustically- and thermally-induced thickness changes in a polymer film Fabry–Pérot interferometerdeposited at the tip of a single mode optical fiber. The sensor provides a peak noise-equivalentpressure of 15 kPa �at 5 MHz, over a 20 MHz measurement bandwidth�, an acoustic bandwidth of50 MHz, and an optically defined element size of 10 �m. As well as measuring acoustic pressure,temperature changes up to 70 °C can be measured, with a resolution of 0.34 °C. To evaluate thethermal measurement capability of the sensor, measurements were made at the focus of ahigh-intensity focused ultrasound �HIFU� field in a tissue mimicking phantom. These showed thatthe sensor is not susceptible to viscous heating, is able to withstand high intensity fields, and cansimultaneously acquire acoustic waveforms while monitoring induced temperature rises. Theseattributes, along with flexibility, small physical size �OD�150 �m�, immunity to Electro-MagneticInterference �EMI�, and low sensor cost, suggest that this type of hydrophone may provide apractical alternative to piezoelectric based hydrophones.© 2009 Acoustical Society of America. �DOI: 10.1121/1.3117437�

PACS number�s�: 43.35.Yb, 43.35.Wa, 43.30.Xm �TDM� Pages: 3611–3622

I. INTRODUCTION

Ultrasonic hydrophones are widely used to characterizediagnostic and therapeutic medical ultrasound fields. In bothcases there are stringent measurement requirements to bemet.1 The increasing prevalence of higher frequency ultra-sound in diagnostic imaging devices, such as very high fre-quency ocular and intravascular scanners, calls for measure-ment bandwidths that extend to several tens of megahertz. Inaddition, to minimize spatial averaging errors at these fre-quencies, an element size smaller than the acoustic wave-length ��75 �m at 20 MHz in water� is desirable. Althoughpiezoelectric polyvinylidene �PVdF� needle and membranehydrophones can readily meet the bandwidth requirement,

a�Author to whom correspondence should be addressed. Electronic mail:
[email protected]
J. Acoust. Soc. Am. 125 �6�, June 2009 0001-4966/2009/125�6

achieving element sizes on this scale with adequate sensitiv-ity is problematic as sensitivity decreases with decreasingelement area.

Characterizing therapeutic ultrasound fields poses asomewhat different measurement challenge. The hydrophonebandwidth and element size requirements are often less de-manding than for diagnostic ultrasound measurements sincethe fundamental frequencies used tend to be in the lowmegahertz range. The greater challenge lies in withstandingthe hostile environment that can be produced, particularlywhere high amplitude or high intensity focused beams areemployed, such as in extracorporeal lithotripsy or high-intensity focused ultrasound �HIFU� therapy. In the lattercase, for example, the cavitational activity and temperaturerises induced at clinical intensities can readily damage PVDFhydrophones. A further requirement for procedures that ex-ploit thermal effects to provide a therapeutic benefit is mea-surement of the induced temperature rise within the sound
field. This is most commonly achieved using wire thermo-
© 2009 Acoustical Society of America 3611�/3611/12/$25.00

couple probes. These are acoustically robust, sensitive, andsmall enough to be inserted into tissue or a tissue mimickingphantom for in-situ measurements.2,3 However, they can suf-fer from significant measurement errors due to viscous self-heating induced by the ultrasound field.4 Although the thin-film thermocouple5–7does not suffer from this limitation, itslarge size prevents it from being used in-situ. Furthermore, itis fragile and susceptible to damage by high intensity fields.It would also be advantageous if both the temperature riseand the acoustic pressure could be measured simultaneouslyat the same location. This would enable the two measurandsto be correlated in order to study the relationship betweenacoustic and thermal effects, for example, to elucidate therole that cavitation plays in enhancing heating in HIFUtherapy. To our knowledge, this type of dual measurementcapability has not previously been demonstrated using non-optical methods.

Fiber optic hydrophones offer the prospect of overcom-ing some or all of the above limitations. The most practicallyapplicable type of fiber-optic hydrophone is the extrinsictype. This employs a fiber-optic downlead to deliver light toand from an optical sensor, or some transducing mechanism,at the end of the fiber. In principle, the sensitive region isdefined by the spot size of the incident illumination. This, inturn, is given by the fiber core diameter which is typically afew microns for a single mode optical fiber. This offers theprospect of providing acoustically small element sizes at fre-quencies of tens of megahertz. Other generic advantages in-clude the small fiber diameter ��150 �m�, a flexible probetype configuration, robustness, low cost, immunity to EMI,and in some cases the ability to measure temperature as wellas pressure. Several types of extrinsic fiber-optic hydrophonehave been investigated for measurement applications inmedical ultrasound. The simplest is based on the detection ofpressure-induced changes in the Fresnel reflection coefficientat the tip of an optical fiber.8–11 This approach offers verylow sensor cost, technical simplicity, and the potential forexceptionally wide bandwidth, limited by the wavelength oflight and thus potentially extending to several hundred mega-hertz. Furthermore, there is the potential to measureultrasound-induced heating since the Fresnel reflection coef-ficient is dependent on temperature as well as pressure. Thefundamental disadvantage, however, is low acoustic sensitiv-ity with a broadband noise-equivalent-pressure �NEP� typi-cally of the order of 1 MPa. This limits its application to themeasurement of temporally stable signals that can be signalaveraged over many cycles or the output of high amplitudesources, such as shockwave lithotripters.9,12,13

Higher detection sensitivities, comparable to those ofPVDF hydrophones, have been obtained by exploiting inter-ferometric methods. A promising approach that has been ex-plored for medical ultrasound metrology is based on the de-tection of acoustically-induced changes in the opticalthickness of a solid Fabry–Pérot interferometer �FPI� locatedat the tip of an optical fiber. Two variants of this approach,distinguished by the materials and methods used to fabricatethe FPI, have been demonstrated. One involves sputteringonto the fiber tip, a multilayer dielectric sensing structure
comprising a central SiO2 spacer sandwiched between two
3612 J. Acoust. Soc. Am., Vol. 125, No. 6, June 2009

stacks of alternating SiO2 and Nb2O5 layers—the latter formthe mirrors of the FPI.14 The small thickness of themultilayer structure �typically a few microns� enables band-widths extending to several hundred megahertz to beachieved.15 Although this type of sensor has not been evalu-ated for the purpose of measuring ultrasound-induced heat-ing, its potential for measuring temperature has been demon-strated by detecting laser-induced thermal transients.16 Thesecond approach employs a different design, one in whichthe spacer between the two FPI mirrors is a polymer.17–20

The lower Young modulus of polymers compared to rela-tively hard dielectrics such as SiO2 results in largeracoustically-induced phase shifts and thus higher sensitivityfor a given FPI finesse.21 The thickness of the spacer tends tobe larger �10–100 �m� than the SiO2 spacer referred toabove providing bandwidths in the tens of megahertz range.It also results in a smaller free spectral range, thereby reduc-ing the required wavelength tuning range for the interrogat-ing laser. An early laboratory prototype hydrophone systembased on this type of sensor has been demonstratedpreviously.20 This system employed a free space 850 nm dis-tributed Bragg reflector �DBR� laser diode as the interroga-tion source �the wavelength of which was manually adjustedin order to optimally bias the FPI� and provided a bandwidthof 25 MHz and a peak NEP of 10 kPa �over a 25 MHzmeasurement bandwidth�. This type of sensor has also beenused to detect laser-induced thermal transients22 and a pre-liminary study has been undertaken to evaluate its ability tomeasure ultrasound-induced heating.23

In this paper, we provide a detailed account of a systembased on this approach but incorporating several key techni-cal developments. These include modifying the design of thesensor and the system detection optoelectronics to extend theacoustic bandwidth to 50 MHz and the use of a computercontrolled discretely tunable “telecom” laser to interrogatethe sensor. The latter significantly enhances the practical util-ity of the concept as it enables the FPI biasing procedure tobe fully automated. Furthermore, because the laser is fiber-coupled, stability problems associated with coupling free-space beams into single mode fibers are eliminated. A majoradvance has been the development and implementation of asensor interrogation algorithm that enables simultaneousmeasurements of temperature and acoustic pressure to bemade for the first time.

Section II A describes the acoustic and thermal trans-duction mechanisms of the sensor and the interrogationscheme used to recover pressure and temperature simulta-neously. Sections II B–II D describe the procedure used tofabricate the sensor, the design of the instrumentation re-quired to interrogate it, and the practical implementation ofthe interrogation scheme. In Sec. III, the acoustic perfor-mance in terms of sensitivity, linearity, frequency response,and directivity are discussed, and in Sec. IV the thermal sen-sitivity and response time are described. The dual measure-ment capability of the system, demonstrated by making mea-surements of ultrasound-induced heating in a tissue phantom,
is discussed in Sec. V.
Morris et al.: The Fabry–Pérot fiber-optic hydrophone

II. THE FABRY–PÉROT FIBER OPTIC HYDROPHONE

A. Principles of operation

A schematic of the tip of the fiber-optic hydrophone isshown in Fig. 1�a�. The sensing element is a FPI which com-prises a thin �l�10 �m� Parylene-C polymer film spacersandwiched between a pair of gold mirrors. Light emitted bya tunable laser is incident on the FPI and is multiply reflectedfrom both mirrors and interferes as it re-enters the fiber.Acoustically- or thermally-induced changes in the opticalthickness of the polymer spacer produce a correspondingphase shift between the light reflected from the two mirrors.This is demodulated to obtain a measure of pressure or tem-perature. In Secs. II A 1 and II A 2, the transduction mecha-nisms for each measurand are described.

1. Acoustic transduction mechanism

The variation in reflected optical power, Pr, from a FPIas a function of phase � is termed the phase interferometertransfer function �ITF� where � is given by

� =4�nl

�, �1�

where n and l are the refractive index and thickness of thepolymer spacer, respectively, and � is the optical wavelength.For a plane parallel FPI formed with non-absorbing mirrorsand illuminated with a collimated beam, the phase ITF takesthe form of the Airy function, an example of which is shownin Fig. 1�b� along with its derivative. In order to make anacoustic measurement, the laser wavelength is adjusted sothat �=�b, where �b is the phase corresponding to the peakderivative of the phase ITF. At this wavelength, the sensitiv-ity and linearity are at a maximum and the FPI is said to beoptimally biased. Under these conditions, a small

Polymer spacer (Parylene-C)

Fibre cladding

Fibre core

l

10 m�

12

5m

�

Gold mirrors

(a)

Reflected optical power, P

r

P h a s e , φφφφ

I T F

I T F '

O p t i m u m p h a s e

b i a s p o i n t , φ

d Pr

d φ

(b)

FIG. 1. �a� Schematic of FPI sensor. �b� Phase �ITF� and its first derivative�ITF��. Operation at optimum phase bias point �b for the linear detection ofa small acoustically-induced phase modulation d� is illustrated.

acoustically-induced phase shift d� can be regarded as being

J. Acoust. Soc. Am., Vol. 125, No. 6, June 2009

linearly converted to a corresponding change in the reflectedoptical power, dPr.

The sensitivity, the reflected optical power modulationper unit acoustic pressure, at �b is given by

dPr

dp=

dPr

d�

d�

dp, �2�

where p is the acoustic pressure. dPr /d� is the first deriva-tive of the ITF at �b and is termed the optical phase sensi-tivity. It is dependent on the incident optical power and thefinesse of the FPI. The finesse is in turn defined by the mirrorreflectivities and the phase dispersion due to the divergenceof the incident beam and non-uniformities in the spacerthickness.

d� /dp is termed the acoustic phase sensitivity and rep-resents the magnitude of the optical phase shift produced perunit pressure. The change in phase arises from a change inthe optical thickness of the spacer and may be caused by twomechanisms, a change in physical thickness or a change inrefractive index. In general, the acoustic phase sensitivity canbe written as

d�

dp=

4�

��n0

dl

dp+ l

dn

dp�Pl�k� , �3�

where dl is a small change in the spacer thickness and dn isa small change in the refractive index. Pl�k� is a frequencymodifying term that accounts for the spatial variation ofstress within the spacer that occurs when an acoustic wave isincident on the sensor and is dependent on the geometry,structure, and physical properties of the fiber tip.24 Equation�3� indicates that d� /dp depends on the elasto-mechanicaland elasto-optic properties of the spacer, although for thepolymer �Parylene-C� spacers used in the sensors describedin this study, the elasto-optic effect can be neglected.25 Equa-tion �3� can therefore be written as

d�

dp=

4�n

�

l

EPl�k� , �4�

where E is the Young modulus of the spacer. Equation �4�shows that the sensitivity is proportional to the thickness ofthe spacer and inversely proportional to the Young modulusof the material. Thus a sensor with a thicker spacer will havea higher sensitivity but reduced bandwidth since the sensorresponds to the spatial average of the stress over its sensitivevolume.

2. Thermal transduction mechanism

When subjected to a change in temperature, the opticalthickness of the FPI spacer will change due to thermal ex-pansion and the change in refractive index �the thermo-opticeffect�. In principle, the resulting phase shift could be recov-ered by optimally biasing the FPI and measuring the re-flected optical power modulation as described in Sec. II A 1.Indeed, such an approach has been employed previously tomake temperature measurements using FPIs.16,22,26,27 How-ever, this method has several limitations. First, it is expectedthat for a polymer spacer, even a relatively small temperature
change of only few degrees will induce a phase shift that
Morris et al.: The Fabry–Pérot fiber-optic hydrophone 3613

exceeds the phase range over which the ITF is linear aroundthe optimum bias point �b. As well as compromising thelinearity of the temperature measurement, a phase shift ofthis magnitude will shift the bias point to the extent that it nolonger corresponds to the peak derivative of the ITF thusreducing the acoustic sensitivity. This will clearly be prob-lematic if simultaneous acoustic and temperature measure-ments are required. In addition, unlike acoustic measure-ments, the relatively long timescale �milliseconds toseconds� of thermal excursions makes temperature measure-ments susceptible to low frequency noise such as that due tofluctuations in the laser output power. For these reasons, thefollowing alternative interrogation scheme which relies ontracking the optimum bias wavelength has been imple-mented.

Figure 2 shows the reflected optical power as a functionof wavelength. This is termed the wavelength interferometertransfer function. At temperature T0, an optimum bias wave-length, �0, can be found which corresponds to �b, the phasebias at the peak derivative of the phase ITF in Fig. 1�a�. Thuswe can write

�b =4��nl�0

�0, �5�

where �nl�0 is the optical thickness at temperature T0. If thewavelength ITF is measured at a different temperature, T1, adifferent wavelength, �1, is now found to correspond to �b

and we can write

�b =4��nl�1

�1, �6�

where �nl�1 is the optical thickness at T1. From Eqs. �5� and�6�, it can be seen that the change in the optimum bias wave-length, ��=�1−�0, which occurs for a temperature change�T=T1−T0, is given by

�� =4�

�b��nl� , �7�

where ��nl� is the change in optical thickness caused by thetemperature change. Providing ��nl� is directly proportionalto the change in temperature, the shift in the optimum biaswavelength will also be proportional to the temperature

Reflected optical power, P

r

W a v e l e n g t h , λλλλ

T0

T1

λ0 λ1

FIG. 2. Effect of temperature on the wavelength ITF. A change in tempera-ture from T0 to T1 produces a linear shift in the optimum bias wavelengthfrom �0 to �1.

change. This wavelength shift could, in principle, be mea-


sured by continually sweeping the laser wavelength andtracking the wavelength ITF as it shifts along the horizontalaxis with temperature. If temperature alone is being mea-sured, this approach would be feasible provided that the timeto sweep the laser wavelength is small compared to the time-scale of the temperature change. However, since the FPI isonly transiently at the optimum phase bias point, �b, it wouldbe difficult to make simultaneous ultrasound measurementswith this method.

For this reason, an alternative interrogation scheme wasdeveloped. This requires initially tuning the laser output tothe optimum bias wavelength and thereafter continuouslymonitoring the reflected optical power Pr. As soon as Pr

changes due to a temperature-induced phase shift, the laserwavelength is immediately re-tuned so as to return Pr to itsinitial value. By continually re-adjusting the laser wave-length to maintain a constant Pr, the thermally-induced shiftin the optimum bias wavelength, and therefore the tempera-ture change, can be tracked over time. Providing the laserwavelength is rapidly re-tuned �before the thermally-inducedphase shift becomes appreciable�, the system will always beoperating at, or close to, �b thus allowing an acoustic mea-surement to be made. In this way ultrasound waveforms canbe acquired at the same time that the temperature is beingmonitored.

The sensitivity of the temperature change measurementusing this method is dependent on several factors: first, thethermo-mechanical and thermo-optic properties of theParylene spacer. The phase change, d�, induced by a tem-perature change, dT, can be written as

d�

dT=

4�l

��n0� +

dn

dT� , �8�

where � is the coefficient of �linear� thermal expansion andn0 is the refractive index at T0. For Parylene-C,28 �=3.5�10−5 °C−1 but no data exist for the thermo-optic coeffi-cient dn /dT. However, Zhang et al.29 proposed an empiricalformula for the calculation of dn /dT of polymers based onthe value of �. Using this, a value of 6�10−5 °C−1 is ob-tained. Unlike the elasto-optic effect, the thermo-optic can-not therefore be neglected. Second, the optical phase sensi-tivity �dPr /d�� influences the thermal sensitivity since itdefines the minimum detectable reflected power variation,which is used to determine when it is necessary to re-tune thelaser to maintain a constant Pr. In addition, the minimumdetectable temperature change depends on the wavelengthtuning resolution of the laser.

B. Sensor fabrication

The sensors were fabricated in batches of 32 as follows.A 2 m length of 1550 nm single mode optical fiber �outerdiameter of 125 �m and core diameter of 9 �m� is cleavedusing an ultrasonic fiber cleaver �PK technology FK11�. Thefibers are then loaded into a mount for the deposition of thereflective coatings and polymer spacer. Thin gold films areused for the reflective coatings, with the reflectivity of thecoating being controlled by the thickness of the film. A nu-
merical model of the ITF was developed in order to inform

the choice of mirror reflectivity.24 It was found that a frontmirror reflectivity of approximately 75% and a back mirrorreflectivity of 98% �the maximum achievable with gold be-tween 1500 and 1600 nm� were optimal. The gold coatingsare deposited onto the tip of the optical fiber via a standarddc sputtering process. The polymer spacer was formed byvapor deposition of Parylene-C, poly�chloro-para-xylene�.20

Sensors were fabricated according to the process outlinedabove with a spacer thickness of 10.4 �m. In order to pro-tect the sensing element, a second Parylene layer 2 �m thickwas deposited over the tip of the fiber, as shown in Fig. 3.

There are several advantages of the above fabricationprocess. The use of a vapor phase deposition process to formthe polymer spacer enables a highly conformal coating withexcellent surface finish, good optical clarity, and uniformityof thickness to be achieved. These attributes allow a highquality FPI with good fringe visibility and finesse to be pro-duced. The thickness can also be precisely controlled��0.1 �m� to design sensors with specific free spectralranges and acoustic bandwidths. Furthermore, the use of allvacuum deposition methods allows batch fabrication of largequantities of sensors with high repeatability at low unit cost.

C. The interrogation unit

Figure 4�a� shows a schematic of the system used tointerrogate the sensor. The components within the dotted boxmake up the interrogation unit which has been developedinto a fully integrated portable prototype �Fig. 4�b��. Thesensor downlead is terminated with an FC/APC connectorwhich is inserted in the front panel of the interrogation unit.Light from a tunable laser is delivered to the downlead bymeans of a 2�2 fiber-optic coupler. Light reflected from theFPI is then routed to an InGaAs photodiode again via thecoupler. A second photodiode is used to monitor the directoutput from the laser. The photodiode measuring the lightreflected from the sensor has both ac and dc coupled outputs.The dc coupled output is connected to an analog to digitaldata acquisition device which is connected to a control per-sonal computer �PC� via a universal serial bus �USB� inter-face. This allows measurement of the reflected optical powerfrom the sensor in order to measure the ITF and monitortemperature changes. The ac coupled output �3 dB cut-offfrequency: 50 kHz� is connected to an oscilloscope for mea-surement of acoustic waveforms.

The laser used in the system is an AltoWave 1100 tun-able laser from Intune Technologies Ltd. The laser has a 40

Optical fibredownlead

125 m�

10 m�10.4 m�

2 m�

Fibre core

Parylene-Cspacer

Parylene-Cprotective coating

Reflectivegold coatings

FIG. 3. Schematic of the FPI sensing structure deposited at the tip of theoptical fiber.

nm tuning range in the telecoms C-band �1528–1568 nm�


and is based on the sampled-grating �SG� DBR design. SG-DBR lasers are monolithically integrated semiconductor la-sers and have no moving parts. This allows rapid, electronicwavelength tuning which is controlled via an RS232 commu-nications interface. The AltoWave 1100 has an optical outputpower of approximately 8 mW, which is constant over itstuning range �0.2 dB�. It is tunable over 600 discrete wave-length channels, separated by a constant optical frequency of8.33 GHz ��0.06 nm increment in wavelength� and can per-form a linear sweep through all 600 channels in a time of120 ms. It is also capable of random channel to channeltuning in approximately 200 �s although in practice this islimited to approximately 2 ms due to the time taken to com-municate with the PC via the RS232 interface.

D. Implementation of sensor biasing scheme

The following describes the practical implementation ofthe interrogation schemes described in Secs. II A 1 andII A 2.

1. Acoustic measurement

To acquire an acoustic waveform, it is necessary to biasthe FPI, as described in Sec. II A 1. In principle, this can beachieved by sweeping the laser through its 40 nm wave-length range and measuring the reflected optical power Pr.

Fibre coupledtuneable laser

( nm)��

ControlPC

2x2 coupler

Digitisingoscilloscope

Transducerdriver

Ultrasoundtransducer

Sensordownlead

USB dataacquisition

card

Interrogationunit

USB 2.0hub

RS232-USB

DC

AC

Fibreconnector

(a)

USBconnection

Sensor downlead

SignaloutputSignaloutput

(b)

FIG. 4. �a� Schematic of the fiber-optic hydrophone system. The compo-nents contained in the dotted box form the interrogation unit shown in thephotograph �b�.

Since the laser channels are separated by a constant optical


frequency, and frequency is proportional to phase, plotting Pr

as a function of channel provides a direct measure of thephase ITF. The latter can then be differentiated and the lasercan be tuned to the wavelength that corresponds to the peakphase derivative. However, it was found that this approachwas compromised by the heating of the FPI due to absorp-tion of the laser light. This self-heating effect corrupts themeasurement of the ITF since the optical thickness of theFPI varies during the wavelength sweep. To mitigate this, theITF was obtained as rapidly as possible by performing asingle sweep through the 40 nm wavelength range of thelaser at the maximum tuning speed. Since the FPI is illumi-nated for only 120 ms, the self-heating is negligible enablingan accurate measurement of the ITF to be obtained. How-ever, a difficulty was then found to arise when the laserwavelength was subsequently tuned to the optimum biaspoint. The FPI is now illuminated for an extended period andthis results in a temperature rise large enough to cause theITF to shift. The FPI is then no longer optimally biased.

To overcome this, an alternative interrogation schemethat relies on iteratively tuning the laser wavelength in orderto search for Prb, the reflected optical power at the optimumbias point, was implemented. This procedure is illustrated inFig. 5. First, the ITF is measured by sweeping the laserthrough its 40 nm tuning range sufficiently quickly that sig-nificant self-heating does not occur as described above. Fromthe ITF measured in this way �denoted “Measured ITF” inFig. 5�, the frequency f1 and reflected power Prb correspond-ing to the optimum bias point are identified. The laser fre-quency is then tuned to f1. As a consequence, the reflectedpower measured by the photodiode is momentarily equal toPrb. However, self-heating at f1 immediately causes the ITFto shift, �becoming “shifted ITF1”� and so the reflectedpower now decreases from Prb to Pr1. In an attempt to returnthe reflected power to Prb, the laser is tuned to f2. This re-sults in a further �but smaller� shift in the ITF which nowbecomes “shifted ITF2.” At f2, the reflected power Pr2 is stillless than Prb so the laser is tuned yet again in order to ap-proach Prb. This procedure is repeated a number of times.With each iteration, the additional temperature rise due toself-heating becomes progressively smaller and therefore sotoo does the shift in the ITF. Eventually, after n iterations, theFPI attains a constant temperature, the position of the ITF no

Reflectedopticalpower,

Pr

Optical frequency, f

Measured ITF

f1f2fn

Prb

Pr2

Pr1

Shifted ITF1

Shifted ITF2

Shifted ITFn

FIG. 5. Iterative scheme to optimally bias the FPI in the presence of self-heating.

longer changes, and the system converges on Prb which cor-


responds to the optimum bias point fn. In practice, the itera-tion is terminated when the measured reflected optical powerlies within 2% of the value obtained in the previous iteration.The system is then ready to make an acoustic measurement.The entire procedure including acquiring the ITF takes ap-proximately 1 s. In practice, this procedure is implementedas an initialization routine when the system is switched on.Thereafter, the reflected power is monitored and if it changesby more than 2% the laser is retuned so as to maintain aconstant value of Pr.

2. Temperature measurement

Assume that the system has been optimally biased asdescribed above and that the sensor is now subjected to anexternal source of heat which produces a temperature riseover some time interval. As the temperature begins to in-crease, the reflected power Pr will change. If Pr changes bymore than a small amount �2%�, the laser wavelength is au-tomatically adjusted by the control software so as to returnPr to Prb, thus returning the system to the optimum biaspoint. As the temperature continues to rise, this procedure isrepeated. In this way, the time course of the temperature riseis discretely sampled by tracking the changes in the biaswavelength as described in Sec. II A 1. Since the FPI is al-ways maintained at the optimum bias point, ultrasoundwaveforms can be acquired at the same time the temperatureis being monitored.

III. ACOUSTIC CHARACTERISTICS

The acoustic performance of the sensor was measuredusing a substitution calibration method30 at Precision Acous-tics Ltd., Dorchester, UK. This was carried out using a cali-brated 0.4 mm �diameter� PVDF membrane hydrophone�Precision Acoustics Ltd., Dorchester, UK� that acted as areference against which the fiber-optic hydrophone measure-ments were compared. The acoustic field was generated inwater by a 1 MHz planar transducer producing a 25 cycletone burst with an approximate peak to peak pressure of 1MPa at the transducer face. Due to the non-linear propaga-tion of the acoustic field in the water tank, integer harmonicsof the fundamental up to at least 60 MHz are generated. Thisarrangement was used to measure the sensitivity, frequencyresponse, and directivity of the hydrophone.

A. Sensitivity

Figure 6 shows a typical measurement of the shockedwave tone-burst as measured by both the reference mem-brane hydrophone �Fig. 6�a�� and the fiber-optic hydrophone�Fig. 6�b��. In both cases the signals were acquired withoutsignal averaging. The calibration sensitivities of the mem-brane hydrophone and fiber-optic hydrophone were 50 and580 mV/MPa, respectively, at 5 MHz. Comparison of thesignals shows that the signal-to-noise ratio of the fiber-optichydrophone is comparable to that of the membrane hydro-phone. The comparison also reveals significant structure inthe signal from the fiber-optic hydrophone; this is due to theprobe-type geometry of the sensor that causes radial reso-
nances and edge waves that propagate across the tip of the

fiber. These are then detected by the sensor as they cross theactive area after the initial acoustic wave has passed. Thiswill be described in more detail in Sec. III C.

The NEP, is defined as the acoustic pressure which pro-vides a signal to noise ratio of 1. The NEP was obtained byrecording the output of the photodiode over a 20 MHz band-width in the absence of an acoustic signal. The rms value ofthe noise voltage was then computed and multiplied by afactor of 3 to obtain the peak value. The peak noise voltagewas then converted to an equivalent pressure by dividing bythe calibration sensitivity �580 mV/MPa� of the hydrophone.This gave a peak NEP of 15 kPa �at 5 MHz� over a 20 MHzmeasurement bandwidth. By comparison, the peak NEP ofthe membrane hydrophone was 10 kPa and that of a 75 �mPVDF needle hydrophone �Precision Acoustics Ltd.� was 28kPa under the same measurement conditions. Note that thepeak rather than rms noise figures are quoted since theformer provides a more realistic indication of the smallestsignal that can be detected when measuring broadband sig-nals in the time domain.

B. Linearity

Assuming the elastic limits of the polymer spacer are notexceeded, the upper limit of linear acoustic detection is de-termined by the phase range over which the gradient of theITF is nearly constant at the bias point and the acoustic phasesensitivity. The linearity is determined by calculating theequation of the straight line which passes through the opti-mum bias point with a gradient equal to the peak derivative

- 2 0

- 1 0

0

1 0

2 0

3 0

4 0

5 0

6 0

7 0

0 5 1 0 1 5 2 0 2 5 3 0 3 5 4 0 4 5

Amplitude (mV)

T i m e ( µµµµs )

- 2 0

0

2 0

4 0

6 0

2 5 2 6 2 7 2 8

(a)

- 3 0 0

- 2 0 0

- 1 0 0

0

1 0 0

2 0 0

3 0 0

4 0 0

5 0 0

6 0 0

7 0 0

8 0 0

0 5 1 0 1 5 2 0 2 5 3 0 3 5 4 0 4 5

Amplitude (mV)

T i m e ( µµµµs )

- 2 0 0

0

2 0 0

4 0 0

6 0 0

8 0 0

2 5 2 6 2 7 2 8

(b)

FIG. 6. Comparisons of the outputs of �a� a 0.4 mm PVDF membranehydrophone and �b� the fiber-optic hydrophone in response to a “shocked” 1MHz toneburst. Insets show expanded timescale �in �s�.

of the phase ITF. The difference between the ITF and the


straight line is then calculated and the range of phase overwhich the difference is less than 5% and 10% determined.20

The linear pressure range is then given by multiplying thisphase range by the acoustic phase sensitivity. A value for thelatter has been determined experimentally in a previousstudy.20 This was found to be 0.075 rad/MPa for a sensorwith a film thickness of 25 �m at a laser wavelength of 850nm. Since d� /dp is proportional to l /�, this must be con-verted to the appropriate value for a sensor 10 �m in thick-ness interrogated by a laser wavelength of 1550 nm. Theacoustic sensitivity can be found from

��

�p�

��2,l2�= ��

�p�

��1,l1��1

l1�� l2

�2� . �9�

Noting that, in this case �1=850 nm, �2=1550 nm, l1

=25 �m, and l2=10 �m, give

��

�p�

�1550 nm,10 �m�= 0.016 rad/MPa. �10�

The linear pressure range �to within 5%� of the fiber-optichydrophone used to obtain the waveform shown in Fig. 6�b�is 4 MPa to 7.5 MPa. If a reduced linearity of 10% can betolerated, this range becomes 6 MPa to 10 MPa. The pres-sure range is not symmetric as the ITF is not symmetricalabout the bias point. The ITF reflectance minimum is alsoasymmetric, due to the use of metallic reflective coatings inthe FPI. These metallic coatings introduce additional phasechanges on reflection and transmission leading to the asym-metry in the ITF. As a result, the peak positive phase deriva-tive is smaller than the peak negative derivative.

C. Frequency response

Figure 7 shows the frequency responses of three nomi-nally identical fiber-optic hydrophones—the responses arenormalized for ease of comparison. The calibration sensitivi-ties of each hydrophone �hydrophones 1–3� at 5 MHz are209, 255, and 580 mV/MPa, respectively—the variation insensitivity is due to the fact that the output power of the laserwas set to a different value for each hydrophone. The uncer-tainties associated with these measurements are as follows:1–15 MHz, 14%; 16–20 MHz, 18%; 22–30 MHz, 23%; and30–50 MHz, 40%, based on combined systematic and ran-

0

0 . 2

0 . 4

0 . 6

0 . 8

1

1 . 2

0 1 0 2 0 3 0 4 0 5 0

Normalised response

F r e q u e n c y ( M H z )

H y d r o p h o n e 1



FIG. 7. Measured frequency responses of three typical fiber-optic hydro-phones.

dom uncertainties. It should be noted that the frequency re-


r freq

sponses shown in Fig. 7 are those of the hydrophone systemand thus include the finite bandwidth of the photodiodes �3dB bandwidth �50 MHz�. Thus, while the response of thehydrophone decreases toward 50 MHz, the intrinsic responseof the sensors is expected to extend beyond this. The re-sponses of all three sensors are in close agreement up to 25MHz indicating good sensor-to-sensor uniformity. Abovethis frequency, there is greater variation which is consistentwith the higher uncertainty in the measurement. The varia-tion may also be due to small differences in the geometriesof the fiber tips, which will have greater influence at higherfrequencies.

Figure 7 also shows that the frequency response of thehydrophone is significantly non-uniform. A detailed investi-gation of the frequency response, including a comparisonusing a finite difference simulation, has shown that the non-uniformities in the response arise from diffraction of theacoustic wave at the tip of the sensor.24 The diffractioncauses multiple edge waves to propagate and reverberateacross the tip of the sensor. The presence of two diffractingboundaries �that of the fused silica fiber and Parylene spacer�leads to waves propagating with more than five differentwave speeds: longitudinal in water, longitudinal and shear inthe Parylene, and longitudinal and shear in the fused silicafiber as well as several interface wavespeeds. It is the fre-quency dependent interaction of all of these waves that isresponsible for the complex structure in the frequency re-

(c)

(a)

FIG. 8. Directional response of a fiber-optic hydrophone. Response shown fo

sponse.


D. Directivity

The directional response of a fiber-optic hydrophonewas measured up to a frequency of 20 MHz at 1 MHz inter-vals, as shown in Fig. 8. The multitude of wave types thatinteract within the sensor and contribute to its output meansthat the directivity inevitably differs from that of an idealrigid disk receiver. It can be seen that for frequencies up to10 MHz, the directional response is well behaved with thesensitivity decreasing with increasing angle and the variationacross the angular range increasing with increasing fre-quency. At 10 MHz, the sensitivity drops approximately 6 dBfor a 90° angle of incidence. Above 10 MHz the behaviorchanges significantly. The measurement at 13 MHz showsthat the maximum sensitivity is no longer obtained at normalincidence, but at approximately 45°. At 15 MHz, the sen-sitivity appears to oscillate as a function of angle, but with amaximum drop in sensitivity of just 1.8 dB across the full180° range. As the frequency is increased to 18 MHz, twolarge drops in sensitivity of greater than 25 dB appear in theresponse at approximately 30°. At 20 MHz, the nulls havereduced in magnitude but occur closer to normal incidence.

IV. THERMAL CHARACTERISTICS

The thermal performance of the sensor was character-

-10

-5

0

-90 -60 -30 0 30 60 90

sitivity(dB)

Angle (degrees)

-25

-20

-15

-10

-5

0

-90 -60 -30 0 30 60 90

Normalisedsensitivity(dB)

Angle (degrees)

16 MHz17 MHz18 MHz19 MHz20 MHz

)

)

uencies: �a� 1–5 MHz, �b� 6–10 MHz, �c� 11–15 MHz, and �d� 16–20 MHz.

(b

(d

ized in terms of its sensitivity, linearity, and response time.


A. Sensitivity

The sensitivity was measured by placing the tip of thefiber in a water bath at room temperature, T0. The sensor wasthen placed into a second water bath held at second tempera-ture, T1, while the change in the optimum bias wavelength,��, was recorded. The temperature change, �T=T1−T0, wasmeasured using a pair of thermocouple probes in a differen-tial measurement configuration. This was repeated for arange of values of T1, from 25 °C to 80 °C, and the resultscan be seen in Fig. 9. In the linear region of the graph �up to�T=45 °C�, a change of 1 nm in the bias wavelength corre-sponds to a temperature change of 5.18 °C. The photodiodenoise-equivalent optical power is much less than the changein reflected power that occurs over a single wavelength tun-ing step. Hence, the temperature measurement resolution isdetermined largely by the minimum wavelength step bywhich the laser can be tuned, approximately 0.06 nm, whichleads to a temperature resolution of approximately 0.34 °C.

B. Linearity and dynamic range

The calibration data in Fig. 9 show that the response ofthe hydrophone to temperature is linear up to a temperaturechange of approximately 45 °C, which in this case corre-sponds to a temperature of 70 °C. Above this, there is anincrease in the gradient of the curve. This is consistent witha glass transition in the Parylene spacer. Below the glasstransition temperature, Tg, the Parylene is in a hard, glassystate and acts elastically. In this state, the thermal expansionof the polymer is linear and reversible. Above Tg, theParylene is in transition between glassy and rubbery moduliand viscoelastic losses are significant.31 Hence the sensor iscurrently limited to measuring temperatures up to 70 °C.

C. Response time

The intrinsic thermal response time of the sensor, basedon the time it takes for heat to diffuse across a 12 �m thickParylene spacer, is approximately 850 �s.22 However, therate at which the system can measure temperature changes islimited by a combination of two factors: first, the acquisitionrate of the USB A-D module which is used to sample the

0

2

4

6

8

10

12

0 10 20 30 40 50 60

Wav

elen

gth

chan

ge,

∆λ∆λ ∆λ∆λ(n

m)

Temperature rise, ∆∆∆∆T (° C)

Measured data

Linear fit

FIG. 9. Change in optimum bias wavelength as a function of temperaturechange.

photodiode output, and second the overhead involved in


communicating with the laser via the RS232 interface. Thesefactors result in a system sampling rate of 200 samples/s forthe temperature measurements. Since the system is only ca-pable of measuring temperature in steps corresponding to thetuning resolution of the laser, there is a maximum temporaltemperature gradient which the sensor can accurately mea-sure. Since tuning one channel corresponds to a temperaturechange of approximately 0.34 °C, the maximum rate of tem-perature change which can be measured is approximately67 °C s−1.

V. MEASUREMENTS OF ULTRASOUND INDUCEDHEATING

In order to demonstrate the applicability of the sensorfor the measurement of ultrasound-induced heating, it wasused to record the temperature rise produced at the focus of aHIFU transducer in a tissue mimicking phantom. These mea-surements were made at the National Physical Laboratory�NPL�, Teddington, UK.

A. Experimental setup

The sensor was embedded in a tissue mimicking phan-tom based on an oil-gelatin emulsion,5 as shown in Fig. 10.To provide a reference for the temperature measurement, athin-film thermocouple �TFT� was also embedded within thephantom approximately 1 cm below the fiber-optic hydro-phone. The TFT was developed at the UK’s NPL in order tomake accurate measurements of ultrasound-induced tempera-ture rises.5,6 The structure of the TFT renders it immune toheating artifacts such as viscous heating, to which wire ther-mocouples are susceptible.4 The fiber-optic hydrophone waspositioned such that the axis of the fiber was perpendicular tothe acoustic axis. This removed the fiber mount from theacoustic path, thereby eliminating acoustic reflections fromthe mount. In order to compare the output of each device, ameasurement was first made with the TFT at the focus. Thesystem was then realigned �including changing the distanceto the transducer� so that the fiber-optic hydrophone was atthe focus and a second measurement made. The heating wasinduced by the output of a 1.07 MHz HIFU transducer witha focal length of 117 mm and a focal spot 3–4 mm in diam-

Water

HIFU transducer(1.07 Mhz)

Thin filmthermocouple

Fibre-optichydrophone

Tissue mimickingphantom (oil/gelatin)

Focal region(max temp rise)

FIG. 10. Experimental setup for making simultaneous pressure and tempera-ture measurements in a HIFU field.

eter and 4–5 cm in length �6 dB�. The heating was con-


trolled by setting the intensity and duration of the �cw� in-sonation. It should be noted that intensity values quoted inthe results below are based on values measured in wateronly. Thus they provide only an approximate indication ofthe values within the phantom.

B. Comparison of temperature-time curves measuredby the fiber-optic hydrophone and thin-filmthermocouple

In the first instance, a comparison of temperature-timecurves obtained by both the fiber-optic hydrophone and theTFT, for a variety of acoustic output settings, was made.Figure 11 shows three temperature-time curves for acoustic

0

2

4

6

8

10

12

14

16

18

20

22

24

0 10 20 30 40 50 60

Tem

per

atu

reri

se(o

C)

Time (s)

Fibre-optic hydrophoneThin film thermocouple

C.

A.

B.

A: I = 51 W/cm2

B: I = 111 W/cm2

C: I = 220 W/cm2

FIG. 11. Comparison of temperature-time curves obtained by the fiber-optichydrophone and the thin-film thermocouple.

0

5

10

15

20

25

30

35

0 10 20

Tem

per

atu

reri

se(°

C)

-0.15

-0.1

-0.05

0

0.05

0.1

0.15

0.2

0 1 2 3 4

Am

plit

ud

e(V

)

Time (µ s)

-0.15

-0.1

-0.05

0

0.05

0.1

0.15

0.2

0 1 2 3 4

Am

plit

ud

e(V

)

Time (µ s)

t = 6 s t = 10 s

FIG. 12. Simultaneous acquisition of a temperature-time curve �top� andinsonation �cw�. For the acoustic waveforms obtained at t=10, 19, and 27 s
the phase shift due to the thermally-induced change in sound speed.

intensities of 51, 111, and 220 W cm−2 ��2.5, 3.7, and 5.4MPa peak-to-peak pressures, respectively�. The insonation at220 W cm−2 was limited to 5 s in order to minimize the riskof damage to the TFT. It can be seen that the shape of thetemperature-time curves measured by the fiber sensor closelymatch those measured by the TFT. This implies that there areno significant viscous heating artifacts in the measurementby the fiber-optic hydrophone. It can also be seen that in thecase of the 111 W cm−2, 30 s insonation, the maximum tem-perature measured by the sensor is slightly higher than thatmeasured by the TFT. This is attributed to a slight error inthe alignment of each sensor in the focal region of the beam.The overall shape remains a good match even in this case.

Additional measurements were made at a range of inten-sities as high as 700 W cm−2 ��12 MPa peak-to-peak�which approaches clinical intensity levels. No damage orchange in performance was observed at this intensity.

C. Simultaneous acquisition of ultrasound waveformsand a temperature-time curve

The arrangement shown in Fig. 10 was also used todemonstrate the ability of the system to conduct simulta-neous acoustic and thermal measurements. The results can beseen in Fig. 12, which shows a temperature-time curve for a30 s insonation with an acoustic intensity �Ispta—spatial peak,temporal average intensity� of 220 W cm−2 and severalacoustic waveforms taken at different times during the in-sonation period. The temperature-time curve is of the sameoverall shape as those obtained for lower insonation intensi-ties or shorter heating periods �Fig. 11� with the exception

40 50 60 70e (s)

I = 220 Wcm-2

-0.15

-0.1

-0.05

0

0.05

0.1

0.15

0.2

0 1 2 3 4

Am

plit

ud

e(V

)

Time (µ s)

-0.15

-0.1

-0.05

0

0.05

0.1

0.15

0.2

0 1 2 3 4

Am

plit

ud

e(V

)

Time (µ s)

t = 19 s t = 27 s

tic waveforms �lower row� captured at four different times during a 30 swaveform captured at t=6 s �gray line� is also shown in order to illustrate

30Tim

acous, the


that several rapid temperature fluctuations can be seen overthe heating period. One possible explanation is that thesefluctuations are related to the presence of cavitational activ-ity close to the tip of the sensor. If bubbles form between thefiber-optic hydrophone and the transducer, some of theacoustic energy from the transducer will be reflected awayfrom the sensor; this would lead to a temporary reduction intemperature. Similarly, if bubbles form behind the sensor,they may reflect some of the acoustic beam back toward thesensor thus increasing the local temperature. Cavitational ac-tivity close to the tip may also convert more acoustic energyto heat, thereby increasing the temperature temporarily whilethe activity persists.

The set of acoustic waveforms accompanying thetemperature-time curve shows that as the temperature rises,the speed of sound increases slightly. The waveforms shift tothe left of the trace, and the amplitude of the signal decreasesslightly. These effects are expected since, in general, soundspeed increases with temperature; thus the time-of-flight ofthe wave will be reduced. In addition, as the sound speed andthe acoustic properties of the phantom change with tempera-ture, the focal point of the transducer may move; thus thesensor will no longer be at the focus and the amplitude maydecrease.

A limitation of the current system is that the acousticwaveforms were captured on a digitizing oscilloscope andtransmitted to the control PC via a general purpose interfacebus interface. With this method, it took approximately 2 s toinstruct the scope to capture a waveform and download it tothe PC. Thus, during the 30 s insonation used to obtain thedata shown in Fig. 12, a maximum of only 15 acoustic wave-forms could be captured. Unfortunately, none of these coin-cided with the rapid fluctuations in the temperature-timecurve so it was not possible to test the above hypothesis thatcavitational activity �which may have been evident from theacoustic waveforms� is responsible for these fluctuations.This limitation could be overcome by using an oscilloscopeor PC digitizing card with an on-board segmented memoryarchitecture. This would allow a large number of successivewaveforms to be captured in real time, concatenated in asingle segmented record, and the whole record downloadedto the PC in a single step as described in Ref. 32. A muchhigher waveform acquisition rate could then be achieved,enabling an acoustic waveform to be acquired for each mea-surement point on the temperature-time curve.

VI. CONCLUSIONS

A practical wideband �50 MHz� fiber-optic hydrophonesystem for characterizing diagnostic and therapeutic medicalultrasound fields has been demonstrated. In terms of acousticperformance, its principal advantage is that it can provide asmall element size with significantly higher sensitivity thancan be achieved with PVDF hydrophones of comparable el-ement dimensions. As well as measuring acoustic pressure,the ability to measure temperature changes produced byultrasound-induced heating has been demonstrated. There aretwo distinguishing features of this capability. First the mea-
surement is free of errors due to viscous self-heating that

afflict wire thermocouples conventionally used to make suchmeasurements. Second, it is possible to acquire ultrasoundwaveforms while simultaneously monitoring the tempera-ture. This unique dual measurement capability provides ameans of directly correlating temperature rises with acousticfield parameters. For example, it could be employed to helpunderstand the role that cavitation plays in enhancing tissueheating in HIFU therapy. A further advantage is that, unlikePVDF hydrophones, the fiber-optic hydrophone appears tobe able to withstand the hostile environment produced by aHIFU field at clinical intensity levels. Even if it should fail,the ability to batch fabricate the sensors at low unit costusing all vacuum deposition techniques means that the sen-sor can be inexpensively replaced. Other advantages includesmall physical size and flexibility, biocompatibility, and elec-trical passivity, attributes that suggest the sensor could beused as an implantable probe for making in-vivo measure-ments.

Future work will be directed toward improving both theacoustic and thermal performance of the hydrophone. Pre-liminary experimental and theoretical studies suggest thatthere is significant scope to obtain a more uniform frequencyresponse by appropriately shaping the tip of the optical fiberto reduce the influence of diffraction.24 This also offers theprospect of producing an improved directional response, par-ticularly at higher frequencies where the fluctuations in an-gular sensitivity are most apparent. Although the currentNEP of 15 kPa is sufficient for many medical ultrasoundmeasurement applications, there is significant scope to in-crease sensitivity if required. The noise performance of thecurrent system is limited by the relatively high phase noiseof the SG-DBR interrogation laser. It has previously beenshown that the use of an external cavity laser as the interro-gation source, which has significantly narrower linewidthand therefore lower phase noise, can provide a NEP of 3kPa.33 There is further potential to reduce the NEP by replac-ing the gold coatings currently used to form the mirrors ofthe FPI with dielectric coatings. These are significantly lessabsorbing enabling a higher finesse FPI to be obtained. ANEP of 0.21 kPa has been achieved with free-space illumi-nated sensors fabricated in this way.34

The sensor is currently limited to making temperaturemeasurements up to 70 °C and the thermal resolution is lim-ited by the wavelength tuning resolution of the laser. Theformer may be improved by using an alternative material toform the spacer, such as the Parylene variant Parylene-HT.This material offers similar optical and acoustic properties toParylene-C but with a higher glass transition temperature.The thermal resolution may be increased by monitoring thevariation in the reflected power and using this data, alongwith knowledge of the ITF, to interpolate between successivewavelength tuning steps. It should then be possible toachieve a resolution limited by the photodiode noise voltagerather than tuning resolution of the laser.

In summary, it is considered that the wide bandwidth,small element size, and high sensitivity of this type of sensor,along with its ability to measure temperature, offers a poten-tially useful new measurement tool for characterizing medi-
cal ultrasound fields.

ACKNOWLEDGMENTS

This work was funded by the UK Engineering andPhysical Sciences Research Council �EPSRC�. The authorsacknowledge the support of the National Measurement Sys-tem Programme Unit of the UK Department for Innovation,Universities and Skills �DIUS� for the work carried out atNPL.

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A Fabry–Pérot ﬁber-optic ultrasonic hydrophone for the ... · case, for example, the cavitational activity and temperature ... heating induced by the ultrasound ﬁeld.4 Although