E-1

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 193-194
16-18 November, 2005

Low-loss and Constant-Group-Delay SAW Filters
Employing Cu-Based R-SPUDTs
H.L Li, J.G. Wen, K. Hashimoto, T. Omori, and M. Yamaguchi (Chiba University)
1. Introduction
Recently, much of an effort has been made to
develop low loss surface acoustic wave (SAW)
filters with constant group delay[1,2,3].
The transversal-type filters designed by theapodization of an interdigital transducer (IDT) offer
superior group delay provided that relatively large
insertion loss is accepted. It is known that asingle-phase unidirectional transducers (SPUDTs)
is an effective countermeasure for reducing theinsertion loss. However, the group delay
deteriorates with a decrease in the insertion loss [3].
The authors recently proposed[4] that low-loss
and constant-group-delay filters can be developedby using a resonant SPUDT (R-SPUDT)[5], andshowed how the technique is effective by anexperiment using Al-electrode/128oYX-LiNbO3-substrate structure. However, because of a smallreflectivity of an Al electrode, the variable range inof the reflectivity is limited unless an extraordinary
large film thickness (>0.09λ) is employed. The
limitation results in poor device performance.
The paper describes the use of Cu as an electrodematerial aiming at designing low-loss andconstant-group-delay SAW filters of excellent
performance.
Theoretical analysis shows that a reflectivity of a
Cu electrode is about two times larger than that ofan Al electrode of the same thickness. Because ofthis, Cu is effective in widening the variable range
in the reflectivity. In addition, the electric
conductivity of Cu is about two times larger thanthat of Al, it is also beneficial for reducing an
ohmic loss.
The constant-group-delay filter is designed andfabricated. Experimental results are in fairly good
agreement with the design; the group delay
deviation of 30 ns was realized over the range of±6.5 MHz at the center frequency of 490 MHz.
The minimum insertion loss and -3 dB bandwidth
were 6.1 dB and 8.5 MHz, respectively.
2. Theoretical analysis
The input admittance of infinitely long
EWC/SPUDTs (see Fig. 1) was calculated for Al

and Cu electrodes on 128oYX-LiNbO3 substrate.
Wc

+

-

-

pI
Fig.1 unit section of EWC/SPUDTThe software MSYNC[6] was used for thenumerical calculation. Then, the normalizedreflectivity (|κ12|pI/π) (pI: periodicity) was estimatedfrom the calculated admittance by using thetechnique described in [6] for a given metallization
ratio Wc/pI (Wc: center electrode width, see Fig. 1)
and electrode thickness h/pI. The width of the side
electrodes is 0.125pI, and the center-to-center
distance between the electrodes are 0.375pI.
Figs.2 (a) and (b) show the estimated reflectivityof the SPUDT for Al and Cu electrodes,
respectively. It is seen that the variable range in thereflectivity of Cu electrodes is expanded to be about
two times larger than that of Al electrodes for agiven film thickness.
Next, it was investigated how filter performances
is improved by the expanded variable range in |κ12|
of Cu.
For the optimization, the electrode thickness was
fixed at 3%λ, and the following specifications wereused as constraints: insertion loss less than 2 dB, -3
dB fractional passband larger than 1.8%, groupdelay error less than 20 ns and IDT length less than
52 periods.
Results of the optimization are shown in Fig. 3(a).
It is seen that the insertion loss for Cu (dashed line)
is about 2 dB smaller than that for Al (dotted line).
The difference is mainly due to the power leakage
to the backward direction of the SPUDTs. Fig. 3(b)shows the group delay for Cu (dashed line). It isalmost constant in transition bands as well as in a
passband. The group delay deviation is about 20 ns
and does not seem to be badly affected by theincreased |κ12|.
The designed device was fabricated on a
128oYX-LiNbO3 substrate. Experimental results arealso shown in Fig. 3(a) and (b) by solid lines, where
the minimum insertion loss at the centre frequencyof 490 MHz and -3 dB bandwidth were 6.1 dB and- 193 -

8.5 MHz, respectively. The group delay deviationof 30 ns was realised over the range of ±6.5MHz.

0
Cu Sim.
Cu Exp.
Al Sim.
-10

0.006
h/PI

-20

0.005

Insertion loss [dB]

0.005
0.01
0.02
0.03
0.04

|κ 1 2 |p I/π

0.004

-30

-40

-50

0.003

-60

0.002
-70
440

450

460

470

480

490

500

510

520

530

540

550

Frequency (MHz)0.001(a)Amplitude response

0
0.1

0.15

0.2

0.25

0.3

0.35

0.4

Sim
Exp.

240

Metallization ratio (Wc/pI)(a) Al electrodes
220

0.012

h/pI
0.01

Groupdelay [ns]

0.005
0.01
0.02
0.03
0.04

200

180

|κ12|pI/π

0.008

160

0.006

140

0.004
460

470

480

490

500

510

520

530

Frequency [MHz](b) Group delay response
Fig. 3 The theoretical and experimental
frequency response of RSPUDT filter
Acknowledgement

0.002

0
0.1

0.15

0.2

0.25

0.3

0.35

0.4

metallization ratio (Wc/pI)(b) Cu electrodes
Fig. 2 Reflectivity per period of Cu and Al
electrodes on 128°LN substrateThis work was partially supported by the
Grant-in-Aids for Fundamental Scientific Researchfrom the Ministry of Education, Sports, Science andCulture.

3. Conclusion
This paper described the application of Cu to thedevelopment of low-loss SAW filters with constantgroup delay in transition bands as well as in a
passband. Theoretical analysis suggested that Cu
electrodes could offer the expanded variable range
in the reflectivity, which is two times larger than Al
electrodes of the same film thickness.
A low-loss SAW filter with constant group delay
was designed and fabricated. Experimental results
were fairly agreeable with the design, and small
group delay deviation of 30 ns was realised over therange of ± 6.5MHz at the center frequency of490MHz.

References
1. G.Martin, F.Z.Bi, D.P.Chen, and C.S.Lam:Proc.
IEEE Ultrason. Symp., (1998) p.165
2. Hiromi Yatsuda:IEEE Trans. Ultrason.,Feroelec., Freq. Contr., 45, No.1( 1998) p.41.
3. C.S.Hartmann and B.P.Abbott: Proc. IEEE
Ultrason. Symp., (1989) p.79
4. H.L Li, J.G. Wen, K. Hashimoto, T. Omori, andM. Yamaguchi: Proc. IEEE Ultrason. Symp.,(2005) [to be published]
5. P. Ventura, M. Solal, P. Dufilie, J. M. Hode andF. Roux: Proc. IEEE Ultrason. Symp., (1994)
p.1
6. K.Hashimoto. J.Koskela and M.Salomaa: Proc.
IEEE Ultrason. Symp., (1999) p.93

- 194 -

E-2

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 195-196
16-18 November, 2005

Ultrasonic Linear Actuator Using Elastic Hinge Structure
Masahiro Takano*, Akichika Nakashima*, Mitsuhiko Taka* and Takaaki Ishii**(*Industrial Research Institute of Ishikawa, **University of Yamanashi)

1. Introduction
In recent years, downsizing is the trend in the field ofelectronic and machine industries and high resolution of
nanometer level is required for measurement andmanufacturing. To realize such demand, precise actuatoris strongly required. In addition to high resolution, somespecial conditions such as non-magnetic interference, use
in vacuum, have to be considered. For that purpose,
ultrasonic actuator is one of the candidates1-3).
Especially, ultrasonic linear actuators are in strong
demand.
The objective of this research is to realize theultrasonic linear stage using ultrasonic linear actuatorwith high resolution and high thrust. The effective
method to obtain ultrasonic elliptical motion was
intensively investigated. Then the proposed actuator
was manufactured and was applied to the linear stage.
Its characteristics were measured and discussed.

2.3 Effect of the elastic hinge structure.
ANSYS 8.1 was used for finite element method (FEM)
analysis. Langevin transducers were used for excitation.
The analysis model is shown in figure 2. Resonancefrequency was 63 kHz and vibration modes with elastic
hinge and without elastic hinge are summarized in figures
3 and 4. In figure 3, coupling vibration was excited.
On the other hand in figure 4, almost independentvibration modes were obtained with elastic hinge
structure. The piezoelectric response was simulated by
FEM under the condition that both transducers were
excited with 90 degrees phase shift. It is shown in figure
5. The displacement at the end of PZT as a function ofdriving frequency was estimated.
The larger
displacement can be obtained with hinge structure. The
elliptical motion was formed in both conditions withelastic hinge and without elastic hinge. Figure 6 showsthe example of the elliptic locus.
Piezoelectric
ceramics 1(PZT 1)

2. Design of the ultrasonic actuator.
2.1 Driving principle of the ultrasonic actuator.
Driving principle of the ultrasonic linear actuator isshown in figure 1. Elliptical motion of the tip attachedto the slider intermittently drives it. The rotation
direction determines the moving direction of the slider.
Slider

Hinge
Piezoelectric
ceramics 2(PZT 2)Vibrator(a) Without elastic hinge (b) With elastic hinge
Fig. 2 Basic configuration of the ultrasonic actuators.

Elliptic motion
SliderSliderVibrator

Vibrator

Elliptic motion

Elliptic motion
SliderVibrator

Elliptic motion(a) Longitudinal mode α(b) Longitudinal mode β
Fig. 3 Natural modes of the actuators without elastic hinge.

Fig. 1 Driving principle of the ultrasonic linear actuator.

2.2 Elastic hinge structure.
Two standing waves with 90 degrees phase shift haveto be excited to obtain a traveling wave. In general,
piezoelectric actuators are used to excite the vibration andthey were rigidly fixed to the vibrator by glue or bolt.
However, the structure of the actuator sometimes
becomes complicated because some pieces of
piezoelectric element are used and vibration mode
analysis also becomes complicated because of its
sophisticated structure and coupling vibrations.
Therefore, elastic hinge structure was proposed for
actuator design to isolate each vibration mode. Elastic
hinge has low stiffness, hence coupling vibration will be
isolated. Almost pure longitudinal vibration mode will
be excited in the Langevin transducers.
Effective
excitation and energy transmission will be expected.
takano@irii.go.jp, ishii@ms.yamanashi.ac.jp(a) Longitudinal mode α(b) Longitudinal mode β
Fig. 4 Natural modes of the actuators with elastic hinge.

2.4 Manufacturing of the ultrasonic actuator.
Figure 7 shows the proposed ultrasonic linear actuatorusing four Langevin transducers. As discussed above,the elastic hinge is made between vibration coupler andLangevin transducer. Four Langevin transducers werebolted symmetrically on the coupler. The shape of thecoupler was designed by the FEM to match the resonancefrequencies of two longitudinal vibration modes. The- 195 -

phase difference of the applied voltage among fourtransducers were described in figure 7.
Figure 8 is the picture of the actuator installed in thestage system. The cross roller guide type precision stage
is used as slider. The contact surface on the slider side
was hard anodized 30µm in thickness for anti-ware
property. The actuator was held at the nodal point of thecoupler with spring, and pressed onto the contact surfaces
with 30 N normal force. Driving frequency was 62.7
kHz.
Stage displacement was measured by thecapacitance type displacement meter.
The applied
voltage was 60 V and driven by burst mode containing 50
cycles. As shown in figure 9, about 4 nm-resolution was
obtained. Figure 10 shows the relationship between
3.5
Displacement (μm)

3.0stage displacement and applied voltage. Although we
can see insensitive range below 60 V, almost linear
relation was obtained.

3. Conclusions
High resolution linear stage was realized using
precision ultrasonic linear actuator. The elastic hinge
structure between piezoelectric element and vibration
coupler was proposed. This structure was simulated by
the FEM. Its effective roles to isolate each vibration
mode and to obtain larger displacement have been
confirmed. The actuator was manufactured and installedin the stage system. As a result, the maximum velocity
of 100 mm/s and resolution of less than 10 nm wereobtained.

PZT 1
PZT 2

2.5
2.0
1.5
1.0
0.5
0.0
50000

55000

60000

65000

70000

Frequency (Hz)(a) Without elastic hinge.
PZT 1
PZT 2

3.0
2.5

Fig. 82.0

Ultrasonic linear stage.

700

1.5
Displacement (nm)Displacement (μm)

3.5

1.0
0.5
0.0
50000

55000

60000

65000

70000

Frequency (Hz)(b) With elastic hinge.

Fig. 5 Displacement at the end of PZT
as functions of driving frequencies.

600
500
400
300
200
100
0
0.0

1.0

2.0

4

6000

-4

-2

0

2

4

-2
-4
x (μｍ）

Displacement (nm）y （μｍ）

4.0

Fig. 9 Stage displacement.
(Input voltage : 60 [Vp-p], cycles : 50)

2
5000
4000
3000
2000
1000
0

Fig. 6 Lissajous figure at the edge ofthe vibrator (With elastic hinge).

40

60

80

100

Input voltage

120

140

160

（Vp-p）

Fig. 10 Stage displacement as a function of input voltage.
(Cycles : 50)

PZT 1(phase:+90°)
PZT 2(phase:-90°)

PZT4 (phase:0°)

PZT3 (phase:0°)
CouplerFig. 73.0Time (s)Structure of the ultrasonic actuator.

References
1) T. Funakubo, T. Tsubata, Y. Taniguchi, K. Kumei, T. Fujiwara
and C. Abe : Ultrasonic Linear Motor Using Multilayer
Piezoelectric Actuators, Jpn. J. Appl. Phys., 34 Part 1, No.5B(1995) 2756
2) T. Furukawa et. al., J. Acoustical Society of Japan, 55-8
(1999) 554 (in Japanese)
3) C, Yun, T. Ishii, K. Nakamura, S. Ueha and K. Akashi : A
High Power Ultrasonic Linear Motor Using a Longitudinaland Bending Hybrid Bolt-Clamped Langevin TypeTransducer, Jpn. J. Appl. Phys., 40 (2001) 3773.

- 196 -

E-3

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 197-198
16-18 November, 2005

A Miniature Multi-Degree-of-Freedom Ultrasonic Motor◎Yasuyuki Gouda, Daisuke Koyama, Kentaro Nakamura, and Sadayuki Ueha
(P & I Laboratory, Tokyo Inst. of Tech.)

1 Introductionz

There are various demands for small-sized
Multi-Degree-of-Freedom(MDOF)
actuator.
Conventional
MDOF
actuator
consistsofelectromagnetic motors and gears, and it has someproblems of noise, complex structure, and large size.
Ultrasonic actuator has been studied for its advantages
in silent operation, simple structure, small size, andhaving a holding force. MDOF ultrasonic actuator has
been prototyped using a combination of a longitudinal
vibration and two orthogonal bending vibrations of arod[1][2].
In this study, we aim to miniaturize the MDOF
ultrasonic actuator and discuss the structure suitablefor fabricating on a substrate with the minimum
number of PZT elements. The actuator consists of acylinder fixed on a substrate and a PZT ring. Electrodeof the PZT is divided into four parts and they are
driven by the phased voltages to excite the 1stlongitudinal
and
2ndbendingvibrations
simultaneously. Height and diameter of the cylinder
are decided so as to tune the resonance frequency of
the 1st longitudinal vibration to the 2nd bending one.
Actuator performances are evaluated experimentallyfor a prototype of a 7-mm-diameter stainless steel
cylinder with a 0.5-mm thickness PZT ring. We
confirmed the rotation of the ball rotor for each of thethree axes and obtained 15.8 rps using a 6-mm ballrotor.

2 Structure and the operation principle
The actuator consists of a stainless steel cylinder
fixed on a substrate and a PZT ring as shown in Fig. 1.
The PZT ring is uniformly polarized in the thickness
direction, and bonded close to the bottom of thecylinder. Electrode of the PZT ring is divided into fourparts, and we can excite a longitudinal vibration andtwo orthogonal bending vibrations simultaneously by
selecting the electrodes. By combining thesevibrations, elliptical motion takes place at the top
surfa
gouda@sonic.pi.titech.ac.jpyxBall rotor

CylinderSubstrate

PZT with
4-divided electrodeFig. 1 Proposed structure of themulti-degree-of-freedom ultrasonic actuator.
z

y

x

z

z

y

+sinωt+sinωt－sinωt －sinωtyxx+sinωt+sinωt+sinωt+sinωt・Bending 1 ・Bending 2
・Longitudinal
Fig. 2 Method to excite each vibration.
surface of the cylinder, and a ball rotor located on thecylinder rotates because of the friction force. Figure 2
shows the method to excite each vibration. To excite
the X-bending vibration, we apply voltages with thephase difference of 180 degrees to two electrodes thatare diagonally located each other. The Y-bending
vibration is excited in the same way as the X-axis
vibration but a pair of electrodes of 90°-shifted is
selected. To excite the Z-longitudinal vibration, weapply the same voltages to all the divided electrodes.
Figure 3 shows the method to rotate the ball rotor. To
achieve the X-axis rotation, we utilize the Y-bending
vibration and the Z-longitudinal vibration. To achieve
the Y-rotation, we use the X-bending vibration and the
Z-longitudinal vibration. To achieve the Z-axis
rotation, we use the X-bending vibration and the
Y-bending vibration. In each case, the phasedifference between the two selected vibrations shouldbe 90° to obtain elliptical vibration locus.

- 197 -zyzX-axis Rotation

x

cosωt
cosωt+sinωt

cosωt

cosωt

3 Frequency design

Z-axis Rotationcosωt

cosωt+sinωt

cosωt+cosωt

cosωt+sinωt

cosωt－cosωt

cosωt－sinωt

X-bending and Z-longitudinal
Fig. 3 Method to rotate the ball rotor.
Z-axisX-bending and Y-bending

20

92.0

Ball diameter ・10 mm

Rotation speed(rps)Resonance frequency(kHz)cosωt－sinωt

cosωt－sinωtyxx

Y-bending and Z-longitudinal

91.0
2nd bending mode

90.0
89.0

z

Y-axis Rotationy

1st longitudinal mode

0
・5 mm

-10

88.0
87.0
0.0

10

-20
0

1.0
2.0
3.0
Inner diameter of the cylinder(mm)

・6 mm
60

120 180 240 300
Phase difference(deg.)360

Fig. 4 Resonance frequencies of the cylinder.

Fig. 6 Rotation speed vs. the phase difference.

To drive the actuator effectively, we have to matchthe resonance frequencies of the 1st longitudinal
vibration and the 2nd bending vibration[2]. We
analyzed these resonance frequencies of a fixed-freecylinder through finite element method. The stainlesssteel cylinder of 7 mm in diameter and 14 mm inheight is assumed to be fixed on a rigid substrate. As
shown in Fig. 4, these resonance frequencies match at
89.469 kHz when the inner diameter is 1.711 mm.

Z-axis rotation at 84.833 kHz. Using the setup shownin Fig. 5, we measured the rotation speed of the ballrotor as a function of the phase difference between theelectrodes. Rotation speed and direction are changed
according to the phase difference as shown in Fig. 5,
and we obtained 15.8 rps for clockwise direction and13.8 rps for counterclockwise direction with the ballrotor of 6 mm in diameter.

4 Operating characteristics

We
proposedanewstructureofmulti-degree-of-freedom ultrasonic motor suitable forminiaturization, and confirmed the X-axis rotation andY-axis rotation at 84.920 kHz, and Z-axis rotation at
84.833 kHz. For the Z-axis rotation, we obtained 15.8
rps using a 6-mm-diameter ball rotor.

Pattern sheet

Amp.

Amp.

Phase Shifter

Amp.

Phase Shifter

Amp.

Phase Shifter

5 Conclusions

6 Acknowledgements
Photo
Interrupter

A part of this study was supported by the
Grant-in-Aid for Scientific Research of Japan Society
for the Promotion of Science, #17360038.

Substrate

References
Fig. 5 Experimental setup.
According to the analysis results, we made a
prototype actuator using a cylinder with the innerdiameter of 1.7 mm and a PZT ring with
0.5-mm-thickness. We confirmed the X-axis rotation
and Y-axis rotation at 84.920 kHz, as well as the[1]Takefumi Amano, et al. “An Ultasonic Actuatorwith Multi-Degree-of-Freedom Using Bending andLongitudinal Vibrations of A Single Stator”, Proc.
IEEE Int. Ultrasonic Symp., pp667-670, 1998.
[2]M.
Aoyagi,
et
al.
“A
Novel
multi-Degree-of-Freedom Thick-Film UltrasonicMotor” IEEE Transactions on Ultrasonics,
Ferroelectrics, and Frequency Control, vol. 49, No.
2, 2002.

- 198 -

E-4

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 199-200
16-18 November, 2005

A piezoelectric rod micro ultrasonic motor with shear-bending modes
Cunyue Lu, Tieying Zhou, Yu Chen, Kai Lu
Department of Physics, Tsinghua University, Beijing 100084, P R China
This paper proposed a new mechanism for exciting shear-bending vibration in a piezoelectric ceramic rod. A
mini-ultrasonic motor prototype with this concept was designed. Both finite element method (FEM) analyses
and experimental results demonstrate the concept.
KEYWORDS: ultrasonic motor; shear-bending mode; piezoelectric ceramic rod

Recently, mini/micro ultrasonic motors have beenof significant research interests. Compared withordinary traveling wave type and standing wave typeultrasonic motors, the shaking head type ones have such
advantages as simple configuration, lower cost and easy
to be miniaturized. So many reported mini ultrasonicmotors belong to shaking head type ones, which use theaxial bending-vibration modes of the stator as theworking modes.1-4) For the purpose of simplify theconstruction, mini ultrasonic motors often use directly apiezoelectric ceramic rod or tube as a stator. Commonly,
previous designs have used the piezoelectric d31 or d33
effect to generate bending vibration. The constant value
for d31 or d33 has a lower value compared with shear
piezoelectric constant d15. Also, the values ofelectromechanical coupling coefficient k31 and k33 aresmaller than that of the shear electromechanicalcoupling coefficient k15.
A design for an ultrasonic motor based on the shear
coefficients has the potential for superior performance,
relative to conventional d31 and d33 based ones. Thus,
this paper presents a special design for an electrode
distribution on a piezoelectric ceramic rod by using theshear piezoelectric d15 effect to excite the bending
vibration modes of the stators.
If an electric field applied to apiezoelectric
ceramic is vertical to the direction of polarization, ashearing strain should be inspired on the piezoelectric
ceramic. Figure 1(a) shows a solid piezoelectric ceramic
rod which is polarized along its axial direction. The
cylindrical surface of the rod is plated with nickel aselectrode. The cylindrical surface electrode was divided
into four pairs (AB, CD, EF and GH) by acid eroding
and each pair has two parts symmetrically along theaxes. For example, the pair AB has part A and part B.
By use of an alternating current (AC) voltage signals(positive and negative), separately on one pairelectrodes, a bending vibration was generated at thestator’s resonance frequency. Figure 1(a) shows thedriving method. Excitation signal A sin (ω t ) are applied
to electrode A, while excitation signal - A sin (ω t ) areapplied to electrode B. At the same time, on thesymmetrical pair of electrodes, excitation signal- A sin (ω t ) are applied to electrode C, and excitation
signal A sin (ω t ) are applied to electrode D. Thus A
bending vibration of a piezoelectric rod is excited asshown in figure 1(b). Similarly, if a A cos(ω t ) voltage

signal are applied to electrode E and H, while a- A cos(ω t ) voltage signal are applied to electrode F andG, the other bending vibration on the piezoelectric rod is
excited. The directions of the two vibrations are verticalto each other. By use of a pair of alternating voltage
signals (sine and cosine) on four pairs of electrodes, ashaking head vibration (wobbling motion) was
generated circumferentially. Because the polarization isalong the axial direction, the piezoelectric ceramic rod
can be well polarized.

(a)

(b)

Fig.1 The electrode distribution and working principle of thepiezoelectric ceramic rod. (a) The electrode distribution. (b)Vibrating mode.

As a prototype example, a compact ultrasonic motor
based on bending vibrating modes of the piezoelectric
rod was fabricated, as illustrated in Figure 2(a).

(a)

(b)

Fig.2 The configuration of a micro motor and its vibrating mode.

It consists of a piezoelectric ceramic stator, a rotor,
and a pressing mechanism including a nut, a spring anda bearing. A copper plate with a slender shaft is bonded
to one end of the piezoelectric ceramic rod. The working- 199 -

bending vibration modes of the stator are as shown infigure 2(b). The local bending vibrating resonancemodes of the piezoelectric ceramic part should be its 1st
local bending vibrating resonance modes.
The role of a piezoelectric stator is to transform the
input electric energy into a mechanical energy. Thistransformation of energy can be described by aneffective electromechanical coupling coefficient, keff,
given asf a2 − f r2
f a2

Fig.4 The measured resonant frequency of the motor

The driving electric voltage of the motor is about
50 V in rms. Fig.5 also shows the relation between therotation speed and the drive frequency. At a resonant
frequency of 72.5 kHz, the motor has a maximum
rotation speed of 109.5 rpm. Under optimum conditions,the maximum speed was slightly increased to 200 rpm.
In the frequency range from 73 to 75 kHz, the motor
speed decreases linearly with the increasing of drive
frequency.
150
Rotational speed(rpm)where fr is the resonant frequency, and fa is theanti-resonant frequency. An ultrasonic motor needs alarge electromechanical coupling factor. The
electromechanical coupling coefficient of thepiezoelectric stator can be modeled using Ansys
software. Fig.3 shows the predicted keff of the bending
modes as functions of the diameter and the length of thepiezoelectric rod. It can be found from Fig.3 thatcoefficient keff increases with the diameter of stator, anddecreases with the length of stator. This modelingrevealed that a piezoelectric rod with a larger diameterand a smaller length would have a larger value of keff.
Thus such a special design presented in this paper forthe electrode distribution on the piezoelectric ceramic
rod is well suitable to a structure of the short fat statorwith a larger value of keff.
L=10mmKeff(%)

0.15

0.05
1

2

3

4

5
D(mm)

6

7

8

Keff(%)

D=6mm

0.3
0.2
0.1
0

4

6

8

10

12
L(mm)

14

16

18

50

60
70
Applied voltage(Vrms)

80

90

100

Applied voltage=50Vrms

60
40

72.5

73

73.5

74
74.5
75
Frequency(kHz)

75.5

76

76.5

77

Fig.5 Characteristics of shear-bending mode motor.

9

0.5
0.4

40

80

20
72

0.1

0

50

100

0.25
0.2

Frequency=72.5kHz
100

0
30

Rotational speed(rpm)K eff2 =agreement with a no-load working condition of the rotor,
as shown in figure 5.

20

Fig.3 Calculated keff as a function of diameter (D) and lenght (L)of the stator.

Because of the fact that the calculated value of 0.12
for keff in the bending mode is much smaller than thevalue for k15 of the piezoelectric material itself, it isquite reasonable to believe that there would be a better
configuration for a micro ultrasonic motor to yield asignificantly higher keff.
In our design, the piezoelectric rod is 6mm indiameter, 10mm long and made with PZT-5h soft-typematerial. The frequency response characteristic of therod measured with impedance analyzer is shown infigure 4. The resonant frequency of the working bending
vibration modes is 72.3 kHz.
The rotational speed of the motor varies
proportionally with an applied voltage. This result is inThis paper proposed a new mechanism for exciting
shear-bending vibration in a piezoelectric ceramic rod. A
mini-motor prototype with this concept and a diameterof 6mm was designed, and its characteristics were
measured experimentally. The design concept presented
in this paper is much promising for the applications insmaller motors and especially in these motors with
piezoelectric ceramic rod stators with a large
diameter-length ratio. There is still much space toimprove for higher efficiency and smaller size of motor,
although both FEM analyses and experimental results
have demonstrated the concept.
The authors thanks for the financial supports fromthe National Natural Science Foundation of China(No.50235010) and State Key Laboratory of Acoustics,
Institute of Acoustics, CAS (No.200506).
1. S. Dong, Siak P. Lim, Kwork H. Lee, J. Zhang, Leong C. Lim,K. Uchino: UFFC. 50 (2003) 361.
2. T. Morita, M. Kurosawa, T. Higuchi: Ultrasonics. 38 (2000) 33.
3. S. Cagatay, B. Koc, K. Uchino: UFFC. 50 (2003) 782.
4. T. Zhou, K. Zhang, Y. Chen: Chinese Science Bulletin. 50(2005) 826.

- 200 -

E-5

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 201-202
16-18 November, 2005

Piezocomposite Ultrasonic Transducer for High-Frequency
Wire-Bonding of Microelectronics Devices
Siu Wing Or* and Helen Lai Wa Chan (The Hong Kong Polytechnic University)
Peter Chou Kee Liu (ASM Assembly Automation Ltd.)
1. Introduction
High-frequency (≥100 kHz) ultrasonic wire bonding is anurgently needed technology to produce smaller wire bondsof higher quality in a shorter time as IC packages in the
microelectronics industry are rapidly reducing in size andincreasing in volume. To enable a complete fine-pitch(<50 µm) and high-speed (>20 wires/s) process, axial-mode
transducers operating at a few kiloherz with a moderately
high mechanical qaulity factor (Qm~300) are of prime
importance [1]. While the hard PZT piezoceramic rings usedin state-of-the-art transducers have large axial coupling(kt~0.38 for PKI804), they also have large lateral coupling(k31~-0.38) and high Qm (~1500). The resulting transducers
thus exhibit strong lateral coupling with many non-axial andspurious modes distributed over the hundred kiloherz range.
This causes the supplied energy to be drained from thedesired axial modes into these undesribable modes, thereby
reducing the purity of axial motion of transdcuers. Moreover,high Qm transducers have sharp resonances with limited
bandwidths; they are sensitive to load changes and impose
difficulties in frequency tuning and mode locking. Therefore,
we have developed novel ultasonic transducers possessing
low lateral coupling and broadband charactersitics toalleivate the drawbacks intrinsic in PZT transducers basedon the PZT/epoxy 1-3 piezocomposite technology [2]. In
this paper, a 136 kHz composite transducer is described andits electrical and vibrational characteristics are reported,
together with a PZT transducer of similar structure.
2. Structure of Ultrasonic Transducers
Figure 1(a) shows the solid model of the composite
transducer. The transducer has a length of 47.38 mm and iscomposed of a Langevin-type piezodriver (12.46 mm long),
an ultrasonic horn (34.92 mm long) and a bonding tool. The
piezodriver has a pair of composite rings, each of 12.65 mm
outer diameter, 5.03 mm inner diameter and 1.70 mm
thickness, connected electrically in parallel, mechanically inseries and preloaded by a prestress mechanism formed by
two 4.47 mm thick SST304 slabs and a SST304 prestressscrew. Three pieces of 0.04 mm thick, ring-shaped Cu
electrodes are inserted between the driver components forelectrical connections. The composite rings are self-made
PKI804/Araldite LY5210/HY2954 1-3 piezocomposite with
0.89 PZT volume fraction. They have sufficiently small
epoxy width of 77 µm and can be treated as an effective
homogeneous material for frequencies up to 10 MHz [3].
The ultrasonic horn is formed by two SST304 half-wave
sectional horns. The section next to the piezodriver has acylindrical rod (13.06 mm long, 8.2 mm diameter) attachedto the small end of a conical horn (3 mm long). The second
section is a conical horn (18.86 mm long) with its smaller-------------------------------------------------------------------------------------* E-mail: apswor@polyu.edu.hk

Electrodes

Back Slab
1–3 Piezocomposite
Ring (× 2)
Front SlabBonding Tool
Screw

Piezodriver

Ultrasonic
Horn

(a)

(b)

Fig. 1. (a) Solid model of the 136-kHz composite transducer and (b)photograph of the fabricated composite (left) and PZT (right)
transducers. The two insets of (b) are the composite and PZT rings.

(a)

(b)

(c)

Fig. 2. Deformed shapes of three major natural modes governed by (a)axial, (b) radial and (c) wall-thickness excitations of the piezoelectricring stack.

end (3.6 mm diameter) affixed with a bonding tool by ascrew. The entire horn is capable of amplifying 2.6 folds theaxial displacement output from the piezodriver. Figure 1(b)shows the fabricated composite transducer along with a
PKI804 PZT transducer.

3. Electrical Resonance Characteristics
The electrical impedance spectra of the composite andPZT transducers were measured using an Agilent 4294A
impedance analyzer. By associating the measured series andparallel resonance frequencies with the resonance andantiresonance frequencies computed using an ANSYS finite
element model, the nature of most experimental resonancemodes was identified and grouped into three categories
according to the excitation forces governed by thepiezoelectric ring stack. These categories include: 1) axial
excitation [Fig. 2(a)]; 2) radial excitation [Fig. 2(b)]; and 3)
wall-thickness excitation [Fig. 2(c)].
As shown in Fig. 3(a), the composite transducer has thestrongest resonance at 135.1 kHz, which is the designatedworking mode of the transducer and also is the 3rd axial
mode (M3) with 1.5 longitudinal wavelengths [Fig. 2(a)].
The two lower modes at 54.5 (M1) and 103.5 kHz (M2) are
the 1st and 2nd axial modes with 0.5 and 1 longitudinal
wavelength, respectively. The two higher modes at 181.8(M4) and 193.7 kHz (M5) are a pair of the 4th quasi-axial
modes. The one observed at 225.9 kHz (M6) is the 5th- 201 -

5. Conclusion
A 136-kHz composite ultrasonic transducer for
high-frequency microelectronics wire bonding has beensuccessfully developed using ring-shaped PKI804/Araldite
LY5210/HY2954 1-3 piezocomposite rings having 77 µm
epoxy width and 0.89 PZT volume fraction. Comparing to
its PZT transducer counterpart, the composite transducer has

10k

50

1k

100

0
-50

100

-100

M3

10
125

130

135

140

o

100

Phase Angle ( )

150

10k

50

-150
150

145

1k

0
-50

100

(a)

M1 M2 M3 M4 M6
M5

100

200

300

400

500

600

-150
800
150

700

100k

150
100

10k

10k

100

50

1k

0

o

10
0
100k

-100

-50

100

-100
M3

10
125

130

135

140

-150
150

145

50
0

1k

-50
100
10

(b)

M4 M6
M1 M2 M3 M5

0

100

200

300

400

500

600

700

Phase Angle ( )Electrical Impedance (Ω)

Electrical Impedance (Ω)

150
100k

-100

-150
800

Frequency (kHz)Fig. 3. Electrical impedance spectra for the (a) composite and (b) PZT
transducers.

Log Amplitude (dBm)

8

(a)

6

Voltage (V)

4

30
0
-30
-60
-90

0 100 200 300 400 500 600 700 800 900 1000

Frequency (kHz)20-2Rise Time

-4
0
8

2

4

Effective Bonding Time

6

8

(b)

6
4

Fall Time10 12 14 16 18 20 22 24
Log Amplitude (dBm)

4. Vibrational Characteristics
The vibrational characteristics of both transducers were
evaluated with a Polytec OFV-303/OFV-3001 laser Doppler
vibrometer. An ASM PC-BQM ultrasonic board operating ina constant voltage, digital phase-locked-loop mode was used
to drive the transducers. To initiate an optimal bonding
condition, a 0.15-W burst in 15 ms with a duty cycle of 13%
was supplied to each transducer at resonance (~136 kHz).
The temporal profile of axial displacement at the tip of theultrasonic horn was examined by measuring the temporalprofile of axial velocity. By applying FFT to the steady-state
portion of the displacement waveform, the corresponding
amplitude spectrum (in dBm re 50 Ω 1 mW) was obtained.
From the waveforms in Fig. 4, the composite transducer
has a rise/fall time of ~50% shorter than the PZT transducer
due to its heavily damped nature as have been seen in its
lower Qm (Fig. 3). Importantly, shortening the rise time candirectly shorten the time required to form a bond. This isdefinitely an advantage for promoting high-speed bonding.
The amplitude spectra (the inserts of Fig. 4) reveal that
excitations of other frequency components (i.e. higher-orderharmonics and sub-harmonics to the fundamental) are ~10
dBm lower in the composite transducer, except thefundamental amplitude (16 dBm or 295 nm) – It is almost
equal to that of the PZT transducer (16.1 dBm or 297 nm).
Hence, the purity of axial excitation of ~95 % is obtained
for the composite transducer but only ~86 % for the PZTtransducer, reflecting again the low lateral coupling natureof the composite-based transducers.

100k

Voltage (V)quasi-axial mode. After that, no prominent resonances arefound. Consequently, the impedance spectrum is very purein that the axial modes are promoted by impeding non-axial
and spurious modes in the composite ring stack.
For the PZT transducer [Fig. 3(b)], the working axial
mode is found at 135.7 kHz (M3), while the two lower axial
modes, M1 and M2, and the three higher quasi-axial modes,
M4, M5 and M6, are located at 54.4, 105.1, 184.5, 195.1
and 227.6 kHz, respectively. Besides, very complicatedresonance characteristics covering two broad frequency
ranges of 250-350 kHz and 400-600 kHz originate from theradial [Fig. 2(b)] and wall-thickness [Fig. 2(c)] activities of
the PZT ring stack. Thus, when this highly laterally coupledtransducer is used in a bonding operation, the non-axial andspurious resonance modes, especially for those withfrequencies very close to the working axial mode or equal toits higher frequency harmonics, are likely to be excited
simultaneously, leading to a reduction in the purity of axial
motion and hence in the bonding stability and yield.
Due to the effect of epoxy damping, the measured Qm ofthe composite transducer at M3 is 296, corresponding to a
3-dB bandwidth of 456 Hz [the insert of Fig. 3(a)]. The PZTtransducer has ~2.4 times increased in Qm (= 710), thus
giving a smaller 3-dB bandwidth of 191 Hz [the insert of
Fig. 3(b)]. Under such a low Qm operation, the composite
transducer will be more adaptable to various types ofbonding surfaces so that a relaible working platform can becreated for high-frequency bonding.

30
0
-30
-60
-90

0 100 200 300 400 500 600 700 800 900 1000

Frequency (kHz)20-2Rise Time

-4

0

2

4

6

Effective Bonding Time

8

Fall Time10 12 14 16 18 20 22 24

Time (ms)Fig. 4. Axial vibration responses at the horn tip for the (a) composite
and (b) PZT transducers, where 1 V = 10 dBm = 147 nm.

superior characteristics of low lateral coupling and reduced
Qm, which made it a promising transducer for use inhigh-frequency wire bonding. Process tests have showedthat an ultrasonic wire bonder when equipped with thecomposite transducer possesses a wider operating window
and a shorter bonding time, thereby promoting the fine-pitchand high-speed capabilities of the machine.

Acknowledgments
This work was supported by the Innovation andTechnology Fund (UIM/117) of the HKSAR Government
and the Centre for Smart Materials of The Hong Kong
Polytechnic University.
References
1. S.W. Or: PhD Thesis (The HK Polytechnic University, 2001).
2. H.L.W. Chan, S.W. Or, K.C. Cheng and C. L. Choy: US Patent US
6,190,497 B1 (2001).
3. S.W. Or and H.L.W. Chan: J. Appl. Phys. 90 (2001) 4122.
4. ASM Assembly Automation Ltd.: Private Communications (2004).

- 202 -

F-1

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 203-204
16-18 November, 2005

A Study on the Behavior of Heating due toAbsorption of Ultrasound in Medium◎ Chiaki YAMAYA and Hiroshi INOUE (Akita Univ.)

1

Introduction

Recently, the ultrasound wave has been put to practicaluse not only the diagnostic equipment but also the thermotherapy that uses the effect of ultrasound wave in living
body. The analysis of temperature rise due to absorptionof ultrasound in the soft tissue medium are important analyzing object for the clarification of the affect in biologicaltissue[1, 2] and the estimation of medium constants. Three
dimensional simulation by FDTD(Finite-Difference TimeDomain) method has been executed by the equations thatconsiders the absorption attenuation based on the acoustic basic equations(ABE)[3–5] ．The Westervelt equation
exists as an equation that describes the absorption attenuation of the ultrasound, too[6–8]. Acoustic field distribution
is calculated respectively by these equations, internal heatgeneration in the medium is calculated from ultrasonic intensity and heat analysis is executed.

2

lossless wave propagation, the loss term and the nonlinear
term, respectively. In this simulation, the nonlinear term isignored in the Westervelt equation . The absorption attenuation is considered by eq.(4) and (5).
Heat conduction equation(HCE)(eq.(6)) and connection
equation(eq.(7)) are as the followings.










  Ú

(1)

  Ú(2) Æ   







(3)

The Westervelt equation cannot be derived from ABE. For
ABE,  is the sound pressure, Ú=(   ) is the particle
velocity,  is the volume elasticity,  is the density,  isthe attenuation coefficient assumed to be proportional toparticle velocity[3][4], where







(4)

is the wave number, is the absorption coefficient,  isthe sound velocity. For Westervelt equation(eq.(3)), Æ isthe acoustic diffusivity, where

Æ



(5)

is the angular frequency.    is the coefficient of nonlinearity with  being the nonlinearity parameter of the medium[9]. In eq.(3), The first and second
terms, the third term and the fourth term describe linear
yamaya@venus.ee.akita-u.ac.jp

 

(7)




(8)

 is the thermal conductivity. In eq.(7),  is the ultrasonicintensity.
ABE(eq.(1) and (2)) and Westervelt equation(eq.(3)) areas the followings.

  Ú

(6)

In HCE (eq.(6)),  is the temperature,  is the internal
heat generated by source,  is the specific heat,  is thethermal diffusivity, where

Analysis method







  





Ô


   Ô  

(9) is the intensity which passes a unit area perpendicular to the propagation direction of an ultrasound  at unit
time,  is the integral multiple of one cycle of an ultrasonic wave or time longer enough than a cycle. The spatial region for analysis in water and the value of medium
constants are shown in Fig.1 and Table 1, respectively.
The absorbing medium is placed at near field length[10]¾for sinusoidal wave,   = ÛÐ ( is the wave length,  isthe radius of sound source), from the sound source. The
size of the absorbing medium is 606060mm . The
sound source is assumed as the circular of 9mm in radius.
The sound source excites a sinusoidal wave of 1MHz and
100kPa.
The cell size is one edge 70 m as a cubic cell in both
ABE and Westervelt equations. The absorbing boundary
condition(ABC) of ABE and that of Westervelt equation
are first order Mur’s ABC and second order Mur’s ABC,
respectively. This reason is that the reflection from ABC
appeared remarkably in the Westervelt equation.

3

Result and discussion

Acoustic field distributions on the center axis of soundsource are shown in Fig. 2. The result of FDTD analysis
by Westervelt equation is in good agreement with that of
ABE. This shows that a similar analytical result of acoustic
field is obtained though these are different equations. It is- 203 -

Sound source a=9mm(1,1001,1001)

Near field lengthx

Normalized sound pressure(2295,1,1001)

z
y

1.2

Absorbingmedium(1,1,1001)(1,1,1)(1,1001,1)

Near field length(2295,1001,1)(2295,1,1)

Fig. 1 Analytic region
Table 1 medium constants





 
 



Absorbingmedium

2.22

2.23

1000

960.0

4186

2149

0.596

0.18

1.42

0.872



6.5

25.3 10 

4Conclusion

0.8
0.6
0.4
0.2

Acoustic basic equation
Westervelt equation

20
40
60
80
100
Distance from sound source [mm]Fig. 21

confirmed that the attenuation in the absorbing medium isthe diffusion attenuation when the absorption medium is
not put and the absorption attenuation with the medium.
Fig. 3 shows temperature distribution in the absorbing
medium on each irradiation time at center axis of soundsource. These results are normalized by each peak temperature on irradiation time 5s. Because similar acoustic
field distribution had been obtained from two equations, acorresponding result was obtained in the temperature distribution.
The Westervelt equation is calculated by one element(the sound pressure) though four elements (thesound pressure(scalar:one element) and the particle velocity(vector:three elements)) are needed in ABE. However,
it is necessary to store the sound pressure in front of 15
steps from the formulation of the Westervelt equation for
FDTD. The FDTD calculation memory in the propagationof ultrasound of ABE and Westervelt equation are 46.1GB
and 62.8GB, respectively. The calculation in Westervelt
equation needs more calculation memories.

10

Normalized temperature

Volume elasticity[ ½¼ N/m ]Density [kg/m ]Specific heat
[J/kg K]
Thermal conductivity[W/m K]
Thermal diffusivity[ ½¼  m /s]
Absorption coefficient[dB/m]

Water

Absorbing medium

0.8
0.6

120

Acoustic field distribution
Acoustic basic equation
Westervelt equation
Irradiation time1 - 5[s]

0.4
0.2
0

0
10
20
30
40
50
60
Distance from the surface of absorbing medium

Fig. 3 Heat distribution on each irradiation time
Acknowledgment This study were supported by a supercomputer resources at Information Synergy Center, Tohoku University, and General Information Processing Center, Akita University.
References[1] J. L. Drewniak and F. Dunn : J. Acoust. Soc. Jpn. 13(1992) 115.
[2] Y. Saito, T. Ishizeki, T. Tsuchiya and N. Endoh : Jpn.
J. Appl. Phys. 44 (2005) 4600.
[3] H. Inoue and F. Yoshida : Proc. Spring Meet. Acoust.
Soc. Jpn. (2002) 1057.
[4] H. Inoue and F. Yoshida : Tech. Report of IEICE,EMD2001-93, (2002) 13.
[5] C. Yamaya and H. Inoue : in Proc. Int. Fed. Med. Bio.
Eng. APCMBE2005, 8 (2005)[6] I.M. Hallaj and R.O. Cleveland : J. Acoust. Soc. Am.,
105 (1999) 7.
[7] I.M. Hallaj   : J. Acoust. Soc. Am., 109 (2001)
2245.
[8] J. Huang   : J. Acoust. Soc. Am., 116 (2004)
2451.
[9] T. Kamakura, “Fundamentals of Nonlinear Acoustics”，AICHI PUBLISHING CO.,LTD., TOKYO(1996)[10] Ultrasound handbook edit committee, “UltrasoundHandbook”, MARUZEN CO.,LTD., TOKYO (1999)

In this study, the temperature distribution was examinedby ABE and Westervelt equation that models absorption ofultrasound that becomes factor of temperature rise whenultrasound is irradiated. The calculation memory in Westervelt equation is larger than ABE though a result corresponding to ABE is obtained by only one element in Westervelt equation. In a future work, a detailed examination
for the heating behavior of ultrasound by the experimentis enumerated.
- 204 -

F-2

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 205-206
16-18 November, 2005

Brillouin Scattering Study of Liquid Glass Transitionin Lithium Borate Glass◎Yuji Ike∗, Seiji Kojima (PAS, Univ. Tsukuba, Japan)
1.

Introduction
Since the properties of alkali metal borate
glasses have anomalous behavior against molar
fraction, the composition dependence of structuraland physical properties are strongly interested.
The anomalous behavior has been deduced fromthe changes in B2O3 structural units. The random
network of B2O3 glass is mostly constructed by
BO3 planar triangle units. The addition of alkali
metal oxide to the B2O3 glass causes theformation of BO4 tetrahedral groups by crosslink
between the BO3 group units building the glass
network. Such properties have been studied byseveral experimental techniques (NMR, Raman
spectroscopy, etc.) [1, 2]. Also the elastic
parameters are sensitive and indicative to thestructure of glass network, thus it plays asignificant role in understanding the structuralcharacteristics of glass network. The change inborate units affects the B2O3 glass network and it
causes a stiffening of the structure. This tendency
is indicated by the increase of the elastic moduli[3]. However, the information of elastic properties
in a high temperature region is scarce. The elastic
properties such as velocity, elastic modulus andattenuation in high temperature, especially at theglass transition temperature have much attracted
in the study of glass transition. In the present
study, different compositions of lithium borate
glass system are investigated by micro-Brillouinscattering technique [4].
2.

Experimental
The composition formula of lithium borate
glass is denoted by x Li2O · (1-x) B2O3, where x
indicates the mole fraction of Li2O. All glasses
were prepared with high homogeneity in order toinvestigate the inherent nature of binary mixture
in lithium borate glasses. Details of highhomogeneity glass preparation method can beseen elsewhere [3].
One of the techniques measuring the elastic
property, Brillouin scattering is known to be a
powerful tool to examine the transparent
condensed materials not only the probe of high∗frequency elastic property, but also the ability to
explore in an extremely wide temperature range.
The experimental setup of a micro-Brillouinscattering apparatus is shown in Fig. 1. The
Brillouin scattering spectra were measured at
backward
scatteringgeometry.
The
Sandercock-type 3 + 3 passes tandem Fabry-Perot
Interferometer (FPI) was combined withmicroscope and operated to acquire the spectra of
the scattered light. A free spectral range of 60
GHz was applied to measure the Brillouin spectra
between ±50 GHz. The temperature wascontrolled and started from 25 ºC to 600 ºC.
Temperature stability was within ±0.1 ºC below
500 ºC.
3.

Results and Discussion
The acquired Brillouin spectra of the lithium
borate glasses showed one longitudinal acoustic(LA) mode. Brillouin shifts and full width at half
maximum (FWHM) have been obtained from thespectra by using the fitting function which takes
account the instrumental function. The
longitudinal velocity of sound Vl was obtained
from the Brillouin shift ν (=qVl/2π) and thescattering wave vector q (=2πnsin(θ/2)/λ), wheren, λ and θ are the refractive index of the sample,
the wavelength of the laser and scattering angle,
respectively. The longitudinal elastic modulus M
is derived by M = ρVl2, where ρ is the density.
The LA mode absorption coefficient α can be
determined by FWHM of Brillouin component
using the equation, α = πΓ/Vl where Γ is FWHM.
The temperature dependence of M are plotted asshown in Fig. 2. The figure indicates M becomeslarger with increasing mole fraction of Li2O.
Larger the elastic modulus indicates the stiffer the
network and this behavior is due to themodification in some amount of cross-linking
borate units from triangle BO3 to tetrahedral BO4
by addition of Li2O. Also we clearly observed thekink in the present temperature range. The
temperature dependence above and below thekink are significantly different. The marked
decrease above Tg is similar to B2O3 and othere-mail: s-ike@ims.tsukuba.ac.jp- 205 -

Fig. 1 Schematic diagram of the micro-Brillouinapparatus. The object lens focuses an incident
laser beam on to a sample and also corrects
scattered light from a sample. Collimatedscattered light passes through two Fabry-Perot
interferometers (FPI 1&2) and detected by aphotomultiplier tube.
longitudinal elastic modulus (N/m)

glasses [5, 6]. Above the temperature M decreases
with increasing temperature and it indicates thedecrease of the elasticity of glass network. The
kinked point temperature can be a glass transition
temperature, Tg, and this temperature shifted tohigher temperature by adding the Li2O. Thistendency is also related to the glass network
variation and higher Tg describes tighter the glass
network by the change of coordination number ofboron atoms. Above Tg, the slope of the curve
becomes steeper hence the glasses are softened byincreasing temperature, while the dependence
below Tg is small and nearly constant. Anotherelastic parameter of absorption coefficient α
exhibits the close relation to the temperaturedependence of M and shows a huge increasebeyond Tg, in a supercooled temperature region.
We observed this behavior is in common with alldifferent concentration of lithium borate glasses.
Further analysis to understand the relaxation
process of the present glass system, thetemperature limitation is the problem at the
moment.
4. Summary
We investigated the elastic properties of lithium
borate glasses in a wide temperature range. At theglass transition temperature the elastic modulus
and also absorption coefficient of LA mode
obviously show anomaly. The glass transition
temperatures increase remarkably with theincrease of the mole fraction of Li2O. These
results are well indicated by the change in someamount of cross-linking borate units of triangle
BO3 to tetrahedral BO4 units. This transformationof glass network leads to the glass more rigid incomparison with a pure borate glass and causes
the increase of glass transition temperatures.

Tg

10

6x10

Tg

10

4x10

x = 0.06
x = 0.14
x = 0.28

10

2x10

0

100

200

300

400

500

600

500

600

o

temperature ( C)-1absorption coefficient α (m )References[1] G. Carini, M. Cutroni, A. Fontana, G. Mariotto,
F. Rocca, Phys. Rev. B 29 (1984) 3567.
[2] J. Zhong, P. J. Bray, J. Non-Cryst. Solids 111(1989) 67.
[3] M. Kodama, T. Matsushita, S. Kojima, Jpn. J.
Appl. Phys. 34 (1994) 2570.
[4] F. Jiang and S. Kojima, Appl. Phys. Lett. 77(2000) 1271.
[5] R. E. Youngman, J. Kieffer, J. D. Bass, L.
Duffrène, J. Non-Cryst. Solids 222 (1997) 190.
[6] H. Anwar, S. Kojima, M. Kodama, B.
Whittington, M. Olesiak, M. Affatigato and S. A.
Feller, Jpn. J. Appl. Phys. (to be published).

Tg

10

8x10

6

2.5x10x = 0.14
6

2.0x10Tg

6

1.5x10

0

100

200

300

400
o

temperature ( C)

Fig. 2 Temperature dependence of longitudinalelastic modulus and absorption coefficient oflithium borate glasses.

- 206 -

INV-1

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 207-208
16-18 November, 2005

Discovery of ATP motor and its mechanism
Masasuke Yoshida (Chemical Resources laboratory, Tokyo Institute of Technology)

All living organisms on the earth rely solely on a single molecule as an energy currency, ATP(adenosine triphosphate). This small molecule supports nearly all the cellular activities that requireenergy, and our body synthesizes roughly as much ATP every day as our body weight. In the
biological world, ATP synthesis is certainly the most prevalent chemical reaction and the enzyme,
ATP synthase, responsible for most of this task, is one of the most ubiquitous, abundant proteins on
the earth. ATP synthase uses physical rotation of its own subunits as a step of catalysis－a novel
mechanism, different from any other known enzymes. Rotation is not a favourite motion in livingorganisms; there is no animal with wheels, no bird with a propeller, and no fish with a screw. On amolecular scale, besides ATP synthase only bacterial flagella are known as a rotary motor. The
crystal structures of the main part of ATP synthase show in atomic detail how the appearance of thisworld tiniest motor made of protein is remarkably reminiscent of the man-made motors. The driving
force that spins ATP synthase is gradient of hydrogen ion concentration across membranes that inturn is made by respiration (burning the food) or by sunshine. We have video-imaged the rotary
motion of ATP synthase that spins as fast as several hundreds revolutions per second. The
mechanism of the motor is completely different from the man-made motor. ATP synthase is
two-particle structure connected with a common rotary shaft. The flow of hydrogen ions through thelower particle drives the rotation of the central rotor that then forces upper particle to make thebending motion for synthesis of ATP. Imagine billions of billion rotary motors are spinning in our
body, day and night, without rest. When the motors stop, we die.

- 207 -

～ Memo ～- 208 -

INV-2

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 209-212
16-18 November, 2005

U

Simulation of Surface Acoustic Wave Devices, Review
Ken-ya Hashimoto
2005-2006 IEEE UFFC Society International Distinguished Lecturer
Faculty of Engineering, Chiba University
1. Introduction
Presently, surface acoustic wave (SAW) filters are
mass-produced and widely used in various consumer
products and communication equipment. For theirresearch and development, use of fast and precise
simulation and design tools is essential, and much effort
has been paid for their enhancement for many years.
Fortunately, recent rapid progress of computer
technologies has made it possible to deal with largescale problems using small personal computers. So as
for computers, anyone can set up the latest research
environment with small investment. The remaining task
is to establish simulation and design software tools.
This lecture reviews simulation technologies used in
the research and development of modern SAW devices
from various aspects.

2. Role and Flow of Numerical Simulation
For the research and development of SAW devices,
the numerical simulation has been widely used forvarious purposes such as (a) understanding underlying
physics, (b) reducing the number of trials, (c) improving
the production yield, and (d) optimizing the device
design.
Although a number of simulation techniques havebeen developed, none of them are perfect. So we mustselect appropriate one for each purpose with the tradeoff between computation speed and precision.
Theoretically, rigorous three-dimensional analysis isapplicable for the whole device structure. However, thisapproach is not realistic because of required computer
speed and memory size.
Fig. 1 shows a typical SAW device structure called
the one-port SAW resonator. The SAW is generated byan interdigital transducer (IDT), and resonates between
two Bragg reflectors. Usually the IDT and reflectors are
uniform laterally, and their aperture is much larger thanthe SAW wavelength. So lateral SAW propagation on
the surface is a minor effect, and its influence is often
ignored.
Now the problem under concern becomes twodimensional, and the rigorous analysis is possible for thewhole structure by the use of current computer power.

Grating (Bragg) Reflectors
VinPiezoelectric substrate
Interdigital Transducer (IDT)
Fig. 1 Typical SAW device structure

However, the simulation is still too slow to use inpractical device design. So further simplification isnecessary.
In the device design, rigorous SAW field is out ofinterest and only electrical properties are necessary to
know. So SAW propagation and excitation are modeled
in behavior basis rather than physical one, reducing to a
one-dimensional problem.
This type of simulators called the behavior model
analysis is mainly used in the optimization process. In
this case, the calculation speed is very crucial because
the same simulator will be executed for a huge numberof iterations to search the optimal solution.
Accuracy of the simulation is critically dependent
upon parameters used in the model. Two-dimensional
full-wave analyses are often used for the extraction of
the parameters. In this case, accuracy is more important
than the calculation speed.
Sometimes SAW oblique propagation on the substratesurface is taken into account after the one-dimensional
analysis. Again use of behavior models is essential, andsimplified two-dimensional analysis is applied for the
extraction of simulation parameters. Since the obliquepropagation is usually a minor effect, not so highaccuracy may not be required.
Finally, influence of parasitic impedances, such as
bonding pads, wires, package, etc., must be taken into
account. This is because even tiny wire inductances givesignificant influences to device performances in GHz
range. Parasitic impedances are estimated experimentally or by the use of commercial tools.

- 209 -

3. Simulation Based on Behavior Model[1]As the behavior models, the coupling-of-modes(COM), p-matrix and equivalent circuit models arewidely used. Difference in these models is how one
period of the IDT is expressed in a three-port element
shown in Fig. 2, where U±(x) is the SAW amplitude
propagating to the ±x direction, and V and I are theapplied voltage and current, respectively.
U+
UI
V
Fig. 2 Three-port element model for IDT
Once each period is modeled, the whole IDT isexpressed in the same three-port structure by themutual-connection.
It should be noted that short-circuited (SC) and opencircuited (OC) gratings are equivalent to IDTs where allelectrodes are short-circuited or isolated electricallyfrom bus-bars, respectively.
Here the COM model is introduced because of its
simplicity and existence of the number of references.
In the model, U±(x) and the current I(x) on the bus-bar
are assumed to be governed by the following
simultaneous linear equations:− jθ p x

+ a−e

IDT

Gap(ϕ)Iin

Gap (ϕ)Reflector

Vin

(4)

+ jθ p x+ jξ − V ,

(5)

(6)

4. Full-Wave Simulation[2]In this category, the finite element method (FEM), theboundary element method (BEM), the spectral domain

(2)

(3)

and
U − ( x ) = Γ+ a + e

Reflector∫+ jξ + V(1)

where V is the applied voltage, C, κ12 and ζ are thestatic capacitance, and the reflection and the excitation
coefficients per unit length, respectively. And θu (=βu2π/pI) is the detuning factor corresponding to deviationof the SAW wavenumber βu from the Bragg condition(βu=2π/pI) under a situation where κ12=0, namely the
SAW reflection is ignored. And pI is the IDT period.
When θu, κ12, ζ and C are uniform (independent of x),
the general solution of Eqs. (1) and (2) is given by+ Γ− a − e22+ jθ p x∂I ( x )= −4 jζ *U + ( x ) − 4 jζU − ( x ) + jωCV ,∂x− jθ p x)

Fig. 3 Behavior model for one-port SAW resonator
Then device characteristics can be readily calculated,
provided that parameters κ12, ζ, C and θu are determinedin advance.
Usually, κ12, ζ and C are almost constant in certain
wide range of frequencies. On the other hand, θuchanges almost linearly with the frequency. So θu isoften expressed in a following form:θ u = ω / Vg − 2π / p I + κ 11 ,(9)where Vg is the SAW group velocity and k11 is aconstant called the self-coupling coefficient.
As it will be shown later, this analysis offers fairly
good accuracy, especially for devices employing
conventional Rayleigh-type SAWs. Of course, accuracyof the analysis is critically dependent upon that of theparameters κ12, ζ, C, Vg and k11.
Here one question arises. How shall we determine
these parameters? Ideally, they are determined from
SAW propagation characteristics in infinitely long SC
and OC gratings. In former days, they were determined
experimentally. Presently, various rigorous numerical
techniques have been developed and are now widelyused for this purpose.

∂U + ( x )= − jθ u U + ( x ) − jκ 12U − ( x) + jζV ,∂x∂U − ( x )*= + jκ 12
U + ( x ) + jθ u U − ( x ) − jζ *V ,∂xandU + ( x) = a + e(

β p = 2π / pI + θu − 4 | ζ |2 / ωC − κ12 − 4ζ 2 / ωC
(7)from the condition of I(x)=0.
Total current I is calculated by the substitution of thegeneral solution into the following equation:∂I ( x)I=
dx = (−4 jζ *U + − 4 jζU − + jωCV )dx. (8)∂xSimple rearrangement of these results and application ofboundary conditions give the three-port model shown in
Fig. 2 for the uniform IDT.
When the IDT is not uniform and the simulation
parameters are dependent on x, Eqs. (4)-(6) can besolved numerically.
Then the whole structure of a one-port SAW
resonator is expressed as an array of the three-port
model as shown in Fig. 3. In the figure, Gaps represent
structural discontinuities between the IDT and thereflector, and are usually regarded as a simple delay line.

∫

U+
U-

[P]

In the equation, βp is the SAW wavenumber in the IDT
with V=0, which is equivalent to the SC grating.
Note that βp for the OC grating is given by

where Γ+=(θu- θu)/κ12, Γ-=(θu- θu)/κ12*,
ξ+=(ζθu- ζ*κ12)/θp2, ξ+=(ζ*θu- ζκ12*)/θ p2, and2

θ p = β p − 2π / p I = θ u2 − κ 12 .

- 210 -

Piece-wise expression
with x-0.5 dependence

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1
Inverse of number of subdivision, 1/(N-1)
Fig. 4 Static capacitance calculated by BEM.

The figure also shows another calculation where thefollowing expansion functions are used:

q( x) =

N −1

anTn (2x / w)

∑ 1 − (2x / w) ,
n =0

2

(10)

where Tn(x) is the n-th order Chebyshev function, w is
the electrode width, and an is the coefficients to be
determined. In this case, convergence is very fast. This
indicates that the charge divergence must be taken into
account appropriately for fast calculation.
It should be noted that conventional FEM tools are
based on continuity of field variables and exhibit slow
convergence for this kind of problems.
The author and coworkers developed a simulation
tool called the FEMSDA for the calculation of SAW
excitation and propagation properties of infinitely long
gratings with finite electrode thickness h. The SDA isapplied for the substrate region. This allows to speed up
the calculation, while the substrate surface is required to
be flat. On the other hand, the FEM is employed only
for the electrode region because it allows arbitraryelectrode cross-section and/or structure.
From the FEM/SDA analysis, an effective permittivity is calculated for the structure. Then it is used todetermine electrical properties by applying a method
proposed Blφtekær, et al., which allows to take the

4000
3980
3960

0.7
h=2%λ
OC

3940

0.3

3920
3900

0.6
0.5
0.4
0.2

SC

Attenuation (dB/λ)

1.16
1.14
1.12
1.1
1.08
1.06
1.04
1.02
1

influence of the charge concentration into accountskillfully for infinitely long gratings.
Fig. 5 shows, as an example, calculated SAW phasevelocity Vp (=ω/Re(βp)) and attenuation α(=−40πloge×Im(βp)/Re(βp)) with propagation in infinite
gratings on a 128oYX-LiNbO3 substrate. As an electrodefilm, Al was chosen. There is a frequency region where
Vp changes linearly and α becomes large. This region iscalled the stopband where the Bragg reflection occurs.

Phase velocity (m/s)

Normalized capacitanceanalysis (SDA) and their combinations are representative.
Attention should be paid for their use in thesimulation of SAW devices. It is known that charge q(x)on IDT fingers concentrates at the electrode edges anddiverges in a form of x-0.5. Since conventional simulation software assumes continuity of field variables suchas the electric flux, the divergence results in very slow
convergence of the calculation accuracy with respect tothe number of variables.
Fig. 4 shows, as an example, the static capacitance C0of the IDT calculated by the BEM. When a piece-wisefunction is employed to express q(x), the conversion isvery bad, and the calculation error decreases in a formof N-1, where N is the number of variables.

0.1
3880
0.45 0.46 0.47 0.48 0.49 0.5 0.51 0.520
Relative frequency, fp/VB
Fig. 5 Calculated SAW phase velocity and attenuation on
Al/128oYX-LiNbO3. Thin lines: FEMSDA, and Boldlines: COM. In the figure, VB=4,025 m/s.
The COM parameters are determined by fitting
calculated βp with that given Eqs. (6) or (7). The figure
also shows Vp and α calculated by using Eqs. (6) and (7)with the determined COM parameters. It is seen thatthese two calculations agree extremely well. This
indicates that the COM analysis described in 3 iseffective for the simulation of Rayleigh-type SAWs.
So we can construct practical simulators provided that
the COM parameters are evaluated as a function ofdesign parameters such as film thickness and width inadvance, and the data are stored in an appropriate form
for the access at the simulation.
Attention should be paid to the fact that the full-wave
analyses require material constants for the electrode film,
and they deviate considerably with deposition andetching techniques, equipment and conditions. Material
constants for the substrate also change with the wafer
supplier and lot. So the estimated COM parameters must
be corrected for individual fabrication process by thefitting with experiments.

5. Simulation of SH-SAW Devices[3]Presently, shear-horizontal (SH) type SAWs arewidely used in low-loss RF filters instead of theRayleigh-type ones. Although behavior-models
described in 3 are also applicable to the SH-type
SAWs, influence of the bulk acoustic waves (BAWs)
must be taken into account for the precise simulation.
Fig. 6 shows, as an example, calculated Vp and α ininfinite gratings on a 42oYX-LiNbO3 substrate. As an- 211 -

2.5
2
1.5
1
0.5

0
0.44 0.48 0.52 0.56
0.6
Relative frequency, fp/VB
Fig. 6 Calculated SAW phase velocity and attenuation on
Al (h=10%λ) /42oYX-LiTaO3. Bold lines: FEMSDA, andthin lines: COM. In the figure, VB=4,226 m/s.
The figure also shows Vp and α calculated by using
Eqs. (6) and (7) with the COM parameters determinedby the method described above. In this case, theagreement is poor. This is due to significant reduction of
the SAW velocity in the stopband. That is, below thecutoff frequency of the BAW reflection, the scattered
BAWs cancel each other, and the cancellation results in
reduction in the SAW velocity (energy storing effect).
On the other hand, this effect diminishes above thecutoff frequency due to the BAW radiation. Note thatscattering amplitude takes a maximum near the cutoff
frequency, resulting in maximum velocity reduction.
Various techniques were proposed for taking thiseffect into account in the behavior model, and arewidely used in the design of current RF SAW devices.
It was also shown that the electromechanical couplingstrength for the SAW varies significantly near the cutoff
frequency of the BAW reflection. Namely, there is asituation where portion of energy radiated as the SAW
is converted as the BAW at the grating fingers and is
detected by the IDT or vice versa.
Fig. 7 shows the admittance of a one-port resonatoron 42oYX-LiTaO3 calculated by taking these effects into
account. In the figure, experimental one is also shown. It
is seen that their agreement is excellent.
Although the simulation seems perfect, there still
remains a room for further improvement. First, thecalculation is too slow to use in the computer
optimization. Second, it cannot simulate behavior at
structural discontinuities properly where the energy
storing effect changes abruptly. So our goal is far
beyond.

25
20
15
10
5
0
-5
-10
-15
865

Relative admittance Y/G 0

4200
4150
4100
4050
4000
3950
3900
3850
0.4

Attenuation (dB/λ)Phase velocity (m/s)

electrode film, Al was chosen. At first look, theirvariation with the frequency seems similar to that shownin Fig. 5 at frequencies lower than the upper edge of the
stopband. It is seen that α remains relatively large at
frequencies higher than the stopband. This attenuation isdue to the Bragg reflection (backscattering) of theincident SAW to BAWs.
StopbandBAW
backscattering
34250

conductancesusceptance870

875 880 885 890 895
Frequency (MHz)Fig. 7 Admittance of one-port resonator on 42oYXLiTaO3.Thick lines: experiments, and thin lines: theory.

6. Remaining Tasks
In this paper, the simulation technologies for the RF
SAW devices were reviewed.
Once simulation tools are ready, it is a starting pointof a trial road. This is because minor effects in formerdays become obvious after evolution, and further
improvement is always necessary.
For example, a little discrepancy can be seen in the
conductance just above the resonance. in Fig. 7, this isdue to the lateral SAW leakage, and its suppression is
one of the hottest topics in the simulation of RF SAW
devices.
In addition, although certain amount of additional
SAW propagation loss was assumed in the simulation,
there is no clear answer where and how the loss occurs.
Once the origin is revealed, one may find propercountermeasure which results in further reduction of thedevice insertion loss. Precise simulators may be also
powerful for such investigation.
Acknowledgement
The author thanks to Prof. Yamaguchi, Dr. Omori andall lab members for their cooperation. This talk is given
as a part of the international distinguished lecturer
program sponsored by the Ultrasonics, Ferroelecctrics
and Frequency Control (UFFC) society, the Institute ofElectrical and Electronics Engineers (IEEE), Inc.
References[1] K.Hashimoto: “Surface Acoustic Wave Devices in
Telecommunications” (Springer-Verlag, 2000)[2] K.Hashimoto, T.Omori and M.Yamaguchi: “Analysis of
SAW Excitation and Propagation Under Periodic
Metallic Grating Structures”, in Advances in Surface
Acoustic Wave Technology, Systems and Applications,Vol. 1, edited by C.C.W.Ruppel and T.A.Fjeldly, (World
Scientific, 2000) pp. 133-182.
[3] K.Hashimoto, T.Omori and M.Yamaguchi: “Modellingof Shear-Horizontal-Type Surface Acoustic Waves andits Application to COM-Based Device Simulation”,
Proc. IEEE Ultrason. Symp, (2001) pp. 127-132.

- 212 -

P2-1

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 213-214
16-18 November, 2005

Probabilistic Superposition of Energy Modes for Treating
2n-Layered Mechanical Impedance Mismatch System
Michio Ohki, Natl. Def. Acad., Japan

1. Introduction
D1

The existence of mechanical impedancemismatch layers inside an elastic resonance systemcauses the shift of resonance frequencies and makes
the analysis of the system complicated. In this study,a systematic analysis of this problem is performedby considering 2n energy modes or complex
dynamical variables, η1, η2, …, η2n, on the circuit
structure representing 2n-layered system, as shownin Figs. 1(a) and 1(b).

(a) Z 0

Z2

Z1
κ1

ξ1

ξ0
= −1
η1

(b)

κ4

κ3

ξ2

ξ3

η2

ξ4
= +1

η3

+

η4

+

d i-1
di
t i-1
r i-1
r' i-1d i-1 t' i-1
d
−

+

−

O i-1
b

P bi

O ib

P bi+1
Z i+1
κ i+1

Behavior of ηi. Terminals Pa , Oa , Obi and
i

η2

η4

η3

η1

η2

η3

η4

η1

η2

η3

η4

η1

η2

η3

η4

η1

η2

η3

η4

Fig. 3 Operation of matrices D1, D2, S1 and S2, bywhich η = (η1, η2, η3, η4)T is multiplied.

iFig. 2
Pbi are located on the i-th circuit block and used for
inputting and outputting ηi in which ηi is multipliedby +κi at Oai and Obi, and by −κi at Pai and Pbi. Whenηi flows in the circuit, it is also multiplied by di, ri, ti,
r’i-1, t’i-1 that represent energy transfer and exchange.

andThe behavior of ηi is illustrated in Fig. 2, in
which ηi is multiplied by appropriate multiplication
factors. Let us consider a column vectorη = (η1, η2, …, η2n)T (T: transpose)as a collection of 2n energy modes.
We define the matrices D1, D2, S1 and S2, bywhich η is multiplied, to deal with the behavior ofenergy modes collectively. The operation of thesematrices is illustrated in Fig. 3 in the case of
4-layered system, and entries of those matrices are
given by
⎛ d1
⎜
d2
⎜
D1 = ⎜
⎜
⎜
⎝

d i+1

+

Zi
κi

d i+1

r' i

t' i

i

Z i-1
κ i-1

−

ti

ri

η4

2. Formulation of Energy Transfer
Exchange in 2n-Layered System

O ia P ai+1

−

S2

η3

η2

η1

S1

ξ

Fig. 1 (a) Resonance system with 2n (n =1,2,3,..)mechanical impedance mismatch layers. The i-th (i=1,2,…,2n) domain has the mechanical impedanceZi and the drive ability κi. (See ref. 1 with regard
to the drive ability.) ξ is a propagation-time-based
coordinate. The case of 4-layered system is
illustrated. (b) Circuit structure representing (a), in
which the energy mode ηi flows clockwise in thei-th block.

O i-1
P ai
a

D2

Z5

Z4

Z3

κ2

η1

d3

⎞
⎟
⎟ , D2 = D1⎟⎟
d 4 ⎟⎠

whered i = exp( − j ω(ξi − ξi −1 ) π /( 2ω0 ) − α i ) ,

(1)

(2)

(ω0 is the characteristic frequency, and αi is thedissipation factor that is regarded as being in
proportion to ξi – ξi−1 for the present), andE-mail: michio_ohki@ieee.org- 213 -

⎛ r1 t1′⎜
⎜ t r′
S1 = ⎜ 1 1
r3
⎜
⎜
t3
⎝

⎞
⎟
⎟,
t3′ ⎟
⎟
r3′ ⎟⎠

⎛ r0′
⎜
r2
⎜
S2 = ⎜
t2
⎜
⎜
⎝

t 2′
r2′

⎞
⎟
⎟,
⎟
⎟
r4 ⎟⎠

Table I 16 cases of matrix parameters in eq. (6)considered using Figs. 2 and 3. R1 = D1S2D2S1, R2 =
D2S1D1S2, R3 = S1D1S2D2, and R4= S2D2S1D1 below.

(3)

j

1

(4)

Since {η1, η3} and {η2, η4} are considered in a“spatially alternating” manner in the input andoutput processes, we prepare matrices K1 and K2:
⎞
⎞
⎛−κ1
⎛+ κ1
⎟
⎟
⎜
⎜
+ κ2
− κ2
⎟,
⎟.
⎜
⎜
K1 = ⎜⎟ K2 = ⎜
⎟
−κ3
+ κ3
⎟
⎟
⎜
⎜
⎟
⎜
⎜
+ κ4 ⎠− κ4 ⎟⎠⎝⎝(5)

The extension of the above formulation into thecase of a general 2n-layered system is easy.
An appropriate matrix product that operates on η
is expected to represent the “path” of η in aprobabilistic manner, and we regard thesuperposition of those energy modes, eachundergoing one of all possible paths, as acharacteristic function of the system, which can beformulated by the infinite geometric series ofmatrix, Neumann series, asη(ω) =∑ K out ( j ) ( A0( j ) + R( j ) A0( j ) + R( j ) 2 A0( j ) + L) K in ( j ) η0

(Kin(j),

R(j)

j

A0(j)

(Kin(j),R(j)Kout(j))with

Z − Z i +12
ri = i, ri′ = −ri , ti = ti′ = 1 − ri .
Z i + Z i +1

A0(j)

(6)

j

as a function of ω, where j is an index to classify
the path, A0(j) and R(j) are the initial term and thecommon ratio, respectively, of the infinite
geometric series that represents the superposition of“possible paths,” Kin(j) and Kout(j) are multiplication
factors for input and output processes, respectively,and η0 = (1, 1,…,1)T is an initial value of η.
Consideration of all possible paths, by inspecting
Figs. 2 and 3, leads to the construction of 16 casesof A0(j), R(j), Kin(j) and Kout(j), as listed in Table I.
3. Calculation Results and Their Interpretation
We investigate a few examples of 4-layered
system, #1, #2 and #3, illustrated in Fig. 4. The #2
and #3 are physically identical. The resonancefrequencies and the resonance intensities(normalized appropriately) of the system arecalculated from eq. (6), which is summarized inTable II. The summation of ηi over i = 1 to 4, Σηi,
is also calculated for each case.
Comparing these results with those from

Kout(j))

D1

R1(K1, K2)

9

S2D2S1

R4(K2, K1)

2

D2

R2

(K2, K1)

10

S1D1S2

R3(K1, K2)

3

S1D1

R3

(K1, K2)

11

R4

R4

(K1, K1)

4

S2D2

R4

(K2, K1)

12

R3

R3

(K2, K2)

5

D2S1

R2

(K2, K1)

13

R1

R1

(K2, K2)

6

D1S2

R1

(K1, K2)

14

R2

R2

(K1, K1)

7

D2S1D1

R2(K1, K1)

15

R3S1

R3

(K2, K2)

8

D1S2D2

R1(K2, K2)

16

R4S2

R4

(K1, K1)#1
Z0 = 0Z1 = 1κ1 = 1Z2 = 1κ2 = 1Z3 = 1
κ3 = 1Z4 = 1
κ4 = 1Z5 = 0#2
Z0 = 0Z1 = 2
κ1 = 0#3
Z0 = 0Z2 = 1κ2 = 1Z5 = 0Z3 = 2 Z4 = 2
κ3 = 0 κ4 = 0

Z1 = 2
κ1 = 0−1

Z2 = 1κ2 = 1−0.5

Z3 = 1
κ3 = 10

Z4 = 2
κ4 = 0Z5 = 0ξ

0.5 0.75 1

Fig. 4 Boundary conditions of calculation examples.
Table II Calculation results for examples #1-#3 of
4-layered system. ωres is the resonance frequency of thesystem, normalized as ω0 = 1.
#

ωres

η1

η2

η3

η4

Σηi

1

1.000

0.144

0.356

0.356

0.144

1.000

3.000

0.856

-0.356-0.356

0.856

1.000

0.7836

0

0.333

0

0

0.333

3.2164

0

0.333

0

0

0.333

0.7836

0

0.167

0.167

0

0.333

3.2164

0

0.167

0.167

0

0.333

2
3

Mason’s equivalent circuit for 3-layered system,2)
we find that both the resonance frequencies and theresonance intensities are correctly obtained, if weregard Σηi as the resonance intensity of the system.
Each ηi indicates the “spatial distribution” of η.
When layers are such that 2n = 2,4,8,16,…, eachwith 2n equal intervals, the methodology ofmulti-resolution (wavelet) analysis of level
0,1,2,3,…, respectively, for treating spatiallyinhomogeneous structure1) can be applied.

- 214 -

1) M. Ohki: Jpn. J. Appl. Phys. 44 (2005) 275.
2) M. Ohki and K. Toda: J. J. A. P. 41 (2002) 3413.

P2-2

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 215-216
16-18 November, 2005

Control of acoustic streaming induced by a focusing
source with two coaxially arranged transducers ∗Kazuhisa Matsuda (Sunagawa High School)
Tomoo Kamakura ( University of Electro-Communications)
Miyuki Maezawa (Olympus Co. LTD)
1

Introduction

2

Experiments and discussion

An ultrasound transducer as a source consistsof two coaxially arranged confocal piezoelectricelements of an inner disc and an outer ring.
Their active surfaces are almost the same: thediameter of the disc element is 27.2 mm, and theinner and outer diameters of the ring elementare 29.2 mm and 40 mm, respectively. Both theelements have the same resonance frequencies of
3.45 MHz and the same focal lengths of 60 mm.
The transducer is flush-mounted on a side wallof a tank filled with fresh water.
Two electric tone-burst signals of 3.45 MHz
and of about 50 cycles duration fed from a function generator are power-amplified individually
and are applied to the transducer. The amplitudes and phases of the signals can be varied precisely. Let the phase difference between
the two signals be θ. When the phase angles
are equal: i.e., θ = 0, the signals are in-phase.
The signals are out of phase when θ = 180◦ . A
needle-type hydrophone with an active diameter of 0.5 mm that picks up local sound pres∗

Tachikawa-shi, 190-8583, Tokyo.

- 215 -(a) In−phase

Fundamental (3.45 MHz)
Second harmonic(6.9MHz)

220

200
40

240
Sound pressure level [dB]

Acoustic streaming occurs because of momentum being transferred from the wave to
fluid through energy dissipation in the wave
caused by sound absorption. It is evident thatthe characteristics of acoustic streaming depend
temporally and spatially on the characteristics
of sound field. This means that it is possible, forexample, to control the streaming by changing
sound fields appropriately. The present report,
therefore, is primarily concerned with some experimenal demonstrations on a simple control
method of the streaming by varying the phasedifference of two initial ultrasound waves emitted from a focused transducer.

Sound pressure level [dB]

240

60
80
Axial distance [mm](b) out−of−phase

Fundamental (3.45 MHz)
Second harmonic(6.9MHz)

220

200
40

60

80

Axial distance [mm]Fig. 1 Fundamental and second harmonic soundpressure levels on the beam axis. Above: in-phasecase (θ = 0), bellow: out of phase case (θ = 180◦ ).

sure is mounted on a translation stage driven
by stepping motors.
Figure 1(a) shows on-axis sound pressure amplitudes of fundamental frequency 3.45 MHz
and second harmonic frequency 6.9 MHz for thecase where the two electric signals are in-phase.
Circle symbols in the figure are all measured
data, and solid curves the theoretical prediction
based on the spheroidal beam equation (SBE)[1]. As can be seen, the theory agrees very wellwith the experiment.
Acoustic streaming was observed in the samewater tank using laser Doppler velocimetry.
Axial distribution of the streaming velocity isshown in Fig. 2. Circles in black and whiteFig. 2 Axial profiles of streaming speed for θ = 0
and θ = 180◦ .

are all the measured data, and solid and dotted curves the theoretical data that are numerically predicted by the stream-function vorticity
method [2]. Overall, the streaming speed is accurately predicted by the theory. The in-phase
excitation accelerates the flow near the focus
and attain a maximum speed of 6.3 cm/s ataround 62mm just behind the focus. After that,the speed is gradually decreased. The speed
profile for the out-of-phase excitation is different: two distinct peaks of the speed appear at
56 mm and 68 mm, and the streaming is slightly
faster at the latter point by 5 mm/s.
Axial speed profiles in the plane perpendicular to the axis are given in Fig. 3, where theplane is located at z = 61.5 mm, just behindthe focal plane. Symbols and notations are thesame as those in Fig. 2. In the paraxial region
less than 2 mm from the axis, the theoreticalpredictions accord with the experimental data
in a relatively good fashion. Away from thepoint, however, the measured speeds are overall larger the computed data.

Fig. 4 Measured data of the sound pressure(above) and streaming speed (below) on the axis.
The angle is changed 0, 60◦ , 120◦ , and 180◦ .

Experimental data are shown for the soundpressure and speed profiles only when the phasedifference θ is changed from 0 to 180◦ by anincrement of 60◦ . Further measurements werecarried out for θ = 240◦ , 300◦ , and 360◦ . However, the data are not shown here because of
their similar profiles as Fig. 4. Evidently, theseexperiments suggest that it is feasible to control
the field of acoustic streaming by changing the
distribution of sound pressure.
This work is partially supported by the Ministryof Education, Science, Sports and Culture, Grantin-Aid for Exploratory Research, 16656062.

References[1] T. Kamakura, T. Ishiwata, and K. Matsuda,“Model equation for strongly focused finiteamplitude sound beams,” J. Acoust. Soc. Am.
107, 3035-3046 (2000).
[2] T. Kamakura, K. Matsuda, Y. Kumamoto, and
M. A. Breazeale “Acoustic streaming inducedin focused Gaussian beams,” J. Acoust. Soc.
Am. 97, 2740-2746 (1995).

Fig. 3

Axial streaming speed at z = 61.5 mm.

- 216 -

P2-3

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 217-218
16-18 November, 2005

Brillouin Scattering Spectroscopy on α–β Phase Transition of Quartz Using
an Angular Dispersion-Type Fabry-Perot Interferometer◎Shinya Tsukada†, Yuji Ike, Jun Kano and Seiji Kojima (Univ. Tsukuba, PAS, Japan)
1. Introduction
In material sciences, it is important to measure thephysical properties of materials without contact andnondestructively. Light scattering measurement is
one of them. Brillouin scattering spectroscopy is an
optical method to measure the elastic property. In
this report, temperature dependence of Brillouinscattering of a well-known quartz crystal wasmeasured in two different geometries using an
Angular Dispersion Fabry-Perot Interferometer(ADFPI). ADFPI is the very powerful tool tomeasure Brillouin scattering [1]. ADFPI has beensolved some problems that scanning type FPI has.
So far ADFPI has been applied to the measurement of
the liquid-glass transition and has shown its excellent
ability [2]. For the measurement of solid materials,
Brillouin scattering of a quartz crystal, a quartz glassand lithium niobate crystal were measured at room
temperature. While no structural phase transition has
not yet observed [3]. In this report, we havesucceeded the observation of structural phasetransition using ADFPI for the first time.
A quartz crystal is one of the most popular crystals
as a jewel, piezoelectric device and optical elements
and so on. In the point of physics, some unique
phenomena are known and they have attracted many
physicists. The α-β phase transition is one of theunique phenomena of quartz crystals. It undergoes aphase transition at 573 oC from its low temperatureα−phase with point symmetry D3 to its hightemperature β−phase with point symmetry D6 [4].

light was collected with the scattering angle of 90°.
The interference fringes can be observed after passing
through a FPI. ADFPI is composed of a solid etalon
which is mechanically stable. Very high reflectivity
coating was used for an etalon to obtain high contrast
and high resolution. The transmitted light through an
etalon was focused onto a CCD detector. The
exposure time of CCD detector was 120 seconds
which became much shorter than that of theconventional scanning method.
The sample
temperature was controlled by the tube-shaped
furnace.
The combination of a CCD detector, a solid etalon
and a light source with high stability of wavelength
enables to make acquisition time much shorter thanthat of a scanning FPI and to make resolution higher.
The synthetic quartz crystal with 6×7×8 mm along
the X, Y and Z axes respectively was used as a sample
of which six faces were polished to optical grade. A
crystal of high quality makes the Rayleigh scattering
small and enables to measure lower frequency
scattering.

2. Experimental
The schematic diagram of the ADFPI is shown in
Fig.1. The light source was a diode-pumped
solid-state laser (DPSS, Coherent) with a ring cavity
operating at 532 nm with 100 mW. The scattered†Fig. 1.

s-tsukada@ims.tsukuba.ac.jp- 217 -

Schematic diagram of experimental setup of
ADFPI.

3. Results and discussion
Acoustic phonons propagating in the [100], [010]
directions were studied in this report. Fig. 2 shows
the spectra with the [010] direction at several
temperatures.
2 1. 1 ℃

L

L

L

T1

L

T1

T1

R

T2 T2

Intensity (arb. units)

5 7 5. 2 ℃

T2 T2

L

T1

L

T1

L

R

T2

T2

T1

L

T2

T2

5 7 6. 4 ℃

L

L

L

R

L
T1

T1

6 0 2. 7 ℃

L

L

L

L
R

T1

T1

-20

0

20

B r i l l o u i n s h i f t ( GH z )Brillouin shift νB (GHz)

Fig. 2. Brillouin spectra of the phonon propagating
the [010] axis of a quartz crystal. L, T1 and T2
denote the Brillouin components scattered by thelongitudinal acoustic phonon, transverse acoustic
phonons, respectively.
R denotes the Rayleighscattering.
30
28
26
24
22
20
18
16
14

Shapiro et al, already measured Brillouin scatteringof quartz [5,6,7] and showed their spectrum at room
temperature [6,7]. Compared with their spectrum atroom temperature, our spectra had better resolution,
because finesse of ADFPI was around 80.
The temperature dependence of Brillouin shift andfull width at half maximum (FWHM) of the [010]phonon were obtained from the Lorentzian fitting as
shown in Fig. 3. In the vicinity of the α−β phasetransition temperature, the anomaly of Brillouin shift
was clearly observed. In particular, the anomaly of
longitudinal wave was apparent. One transverse
mode disappeared above the phase transition
temperature. It is attributed to the change of thecrystal structure into D6. The transition temperatureobserved was 574 oC that is 1 oC higher than thetemperature reported in [4,5]. The FWHM shows apeak at the phase transition temperature. FWHM isrelated with the damping constant [8], the attenuationof the sound wave became larger above the phasetransition temperature.
It is known that
incommensurate phase exists in the range of 1.4 oC
around the phase transition temperature [9]. Elastic
anomaly related to an incommensurate phase between
α and β phases could not be obtained in the present
experiments.
In summary, we applied ADFPI to the study ofstructural phase transition for the first time andsuccessfully obtained spectra with the variation oftemperatures. In the vicinity of the α-β phasetransition temperature, the anomaly of Brillouin shift
and FWHM were clearly observed.

LA

TA 2
TA 1
0 100 200 300 400 500 600 700

FWHM (GHz)2■ LA○ TA 1△ TA 2
1

0References
Jae-Hyeon Ko and Seiji Kojima: Rev. of Sci.
Inst. 73 (2002) 4390
[2] Yuji Ike and Seiji Kojima: J. Mol. Struct. 744(2005) 521[3] Yuji Ike and Seiji Kojima: J. Kor. phys. soc. 46(2005) 90[4]
Robert B. Sosman: The properties of silica,(The Chemical Catalogu Company, Inc. New
York, 1927)[5]
S. M. Shapiro and H. Z. Cummins: Phys. Rev.
Lett. 21 (1968) 1578
[6]
Stephen M. Shapiro, Robert W. Gammon andHerman Z. Cummins: Appl. Phys. Lett. 9 (1966)
15[7]
Stephen M. Shapiro and Herman Z. Cummins:
Proc. the International Conference on Light
Scattering Spectra of Solids (1968) 705[8] Terutaro Nakamura: Ferroelectricity involvedin Structual Phase Transition (Shokabo Tokyo,
1998)
[9]
G. Dolino and P. Bastie: J. Phys. Condens.
Matter 13 (2001) 11485[1]

0

100 200 300 400 500 600 700
Temperature (℃)

Fig. 3. Temperature dependences of Brillouin shift
and full width at half maximum of a quartz crystals.
- 218 -

P2-4

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 219-220
16-18 November, 2005

Analysis of stress-induced ferroelectric domain structure
by ultrasonic atomic force microscopy◎Seishiro Ide, Toshihiro Tsuji, Kazushi Yamanaka (Tohoku University)

1. Introduction
Ferroelectric materials applied to piezo-actuators
are expected to experience combined electrical andmechanical loads. However, only few studies onstress-induced ferroelectric domain structure havebeen made [1]. In this paper, we report an attempt toanalyze ferroelectric domain structure induced using
nanoindentation and compare the elastic andpiezoelectric properties on the nanoscale.

frequency and the amplitude at resonance [3,4]. We
observed that both frequency and amplitude were highin the center of the indentation, and low in the
periphery of indentation. We consider that the highfrequency probably includes an artifact due to thesteep slope in the center of indentation, but the low
frequency showed elastic property because of gentle
(a)
(a)

(b)
[001]

2. Experimental method

-105

500nm
0.0
(c)

[001]
105(nm)Lateral crack

6

Node

[100]

81
0]
[10 (nm)0

We applied a nanoindentor attached to an atomic
force microscope (AFM) with cube cornered diamond
tip on polished ferroelectric samples. We prepared
PMN-PT single crystal (sample 1) and soft PZTwithout poling (NEC Tokin, N-21) (sample 2), andindented up to maximum load 1500µN and 1000µN,
respectively.
We used ultrasonic atomic force
microscopy (UAFM) [2-4] and piezo-response force
microscopy (PFM) [5] to observe before and after
indentations, where we applied 2nd deflection mode
in UAFM, which is effective in stiff materials.

0

)
(n m

81 6

(d)

Ultrasonic vibration
Base

Tip

500nm

Cantilever

537.0

SampleFig.1. Schematic illustration of UAFM at the 2nd
deflection mode.

500nm
541.7(kHz)

0.13

(e)

(f)

500nm

500nm

0.31(nm)

3. Result and discussion
3.1. PMN-PT single crystal
Figure 2 shows a result of the observation of
nanoindentation on sample 1. Fig.2 (a) is AFM
topography. The directions of arrows show
crystallographic orientations. In spite of using acube cornered tip, the upheaval around the indentation
was of rectangular shape. It may be related to
crystallographic structure. Fig.2 (b) is a 3D viewaround indentation, where grayscale is inverted tohighlight the shape of the indentation. Fig.2 (c) and(d) are UAFM images representing the resonance

0.013

0.091(nm)0.0112.1(o)Fig.2. Nanoindentation on PMN-PT single crystal. (a)Topography. (b) 3D view around indentation.
Grayscale is inverted. UAFM image representing (c)resonance frequency and (d) amplitude at resonance.
PFM image representing (e) amplitude and (f) phaseshift.

- 219 -

slope around indentation.
The latter can be
explained by the decrease of effective stiffness due tolateral cracks [6].
On the other hand, PFM images showed bothincrease and decrease of amplitude (Fig.2 (e)) anddelayed phase (Fig.2 (f)) in the periphery of theindentation.
Around the indentation, we observed the decreaseof amplitude and delayed phase expanding towards[100] and [001] directions. Since they had not been
observed before indentation, we consider that acrystallographic reorientation caused by the stressfield decreased the piezo-response in these regions.
They were observed as a slight increase of frequency
in UAFM images.
3.2. Soft PZT ceramics
Figure 3 shows the observation of nanoindentation
on sample 2, where solid squares show common area.
Fig.3 (a) and (b) are AFM topography and 3D view(a)

300 nm

(b)

0.0
(c)

529.3
(e)

（nm
）

5 55

67.0(nm)
300 nm(d)

300 nm

536.7(kHz) 0.26(f)
300 nm

3.3. Common structure in both samples
In both samples, we observed stress induced
structures around indentations. In PMN-PT, thesestructures were induced along (100) directions. In
PZT, these structures located similarly to PMN-PT.
We consider that they are related both to the stressfield and to the crystallographic orientation.
4. Conclusions

0

0

（ nm
）

5 55

67

around indentation. There was not a steep edge dueto the corner of indentor. Fig.3 (c) and (d) show
UAFM images representing resonance frequency
and amplitude.
Within the indentation, there was a high frequency
region surrounded by a sharp boundary (Fig.3 (c)). In
Fig.3 (d), this region showed increase of amplitude.
Fig.3 (e) and (f) show PFM images representing
amplitude and phase shift. In Fig.3 (e), there wasdecrease of amplitude within the indentation (Fig.3(a)). This may be related to the high frequency
region in Fig.3 (c). In Fig.3 (f), phase was advancedin the center of the indentation. In the regions
around the indentation indicated by broken lines, weobserved cyclic changes of frequency (Fig.3 (c)).
Since they were not observed in topography (Fig.3(a)) and UAFM image before indentation, we consider
that they are structures induced by stress field duringnanoindentation.

0.99(nm)

We compared the elastic and piezoelectric
properties of ferroelectric materials on the nanoscaleby simultaneous observation within and aroundnanoindentations using ultrasonic and piezo-response
atomic force microscopy. Then a sign of the stressinduced ferroelectric domain structures was obtained.
We believe that this observation method will be usefulfor comprehensive evaluation of these materials
applied to actuators.
Acknowledgements
This work was supported by the Ministry of Education,
Culture, Sports, Science and Technology.

300 nmReferences

0.63

0.81(nm)0.0

20.4(o)Fig.3. Nanoindentation on soft PZT ceramics. (a)Topography. (b) 3D view around indentation.
Grayscale is inverted. UAFM image representing (c)resonance frequency and (d) amplitude at resonance.
PFM image representing (e) amplitude and (f) phaseshift.

[1] D. Liu and K. W. White: Appl. Phys. Lett., 85(2004) 3459.
[2] K. Yamanaka and S. Nakano: Jpn. J. Appl. Phys.,
Part 1, 35 (1996) 3738.
[3] T. Tsuji, H. Ogiso, J. Akedo, S. Saito, K. Fukuda
and K. Yamanaka: Jpn. J. Appl. Phys., Part 1, 43(2004) 2907.
[4] T. Tsuji, S. Saito, K. Fukuda, K. Yamanaka, H.
Ogiso, J. Akedo and Y. Kawakami: Appl. Phys. Lett.,
87 (2005) 071909.
[5] A. Gruverman, O. Auciello and H. Tokumoto:Appl. Phys. Lett., 69 (1996) 3191.
[6] K. Yamanaka, Y. Enomoto and Y. Tsuya: IEEE.
Trans. Sonics and Ultrasonics, 32 (1985) 313.

- 220 -

P2-5

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 221-222
16-18 November, 2005

Study on Layer Mode Device on GaN/Al2O3
Manabu Yokota, Chinami Kaneshiro, Kazumi Nishimura*, Naoteru Sigekawa*, and Kohji Hokawa
Kanagawa Institute of Technology, ＊) NTT Photonics Laboratories

I.Introduction
GaN film has many features as semiconductor
devices, such as blue light emitting diodes, refractive
transistors and ultra violet wave sensors, etc. The
system also has piezoelectricity and expected for new
semiconductor coupled acoustic wave functional
devices in higher frequency range in a future. Layer
modes are effective for fabricating higher frequency
devices under conditions of the limited fabrication
technology for realizing small size pattern. Based on aview point of device design, we studied layer modes
on GaN/Al2O3 system using result of mode
calculations and experiments. The layer mode has
shown a relatively wide band characteristics and aninteresting response on delay.

substrate is faster than that in GaN. On the other hand,devices with propagation direction <1-100>, energy
confinement in layer is relatively low. The layer mode
does not exist under the same wave number at which
the wave along <11-20> has multi mode transmission
ode condition.

II.Experimental conditions of layer mode devices
In order to study actual potential of device
performance, we have fabricated delay lines on GaN
films (including a commercially available film) andmeasured device characteristics. Parameters used in
the experiment are listed in Table I.
Table I

Experimental device parameters

40 I.L.[dB]

SAW

50

layer

60
70
80

1

Freq. [GHz]2a) Path length=2.5mm LS=2um, kH=1.57
40 I.L.[dB]

layer

SAW

60
80
100
0

1.0

Frec. [GHz]2.0b) Path length=5mm LS=0.8um

GaN films

4 kinds of films, includingcommercially available.
Film thickness 0.15um, 2 um
Substrate
Sapphire (c-plane)
Prop. direction <1-100> and <11-20>
IDT
50 pairs uniform overlap
N. of pairs
25 pairs for low freq. device。
Line & space
0.8um, 2um, 4um
Path length
2.5 mm, 5 mm

kH=3.93
Figure 1 Response for <11-20> prop. direction
I.L. [dB]
30

SAW
50

layer；leaky wave

70

In this experiment, we adopted film thickness
smaller than wave length (4 times of line width).

90
0

1.0

2.0

Freq. [GHz]a) Path length=2.5mm LS=2um, kH=1.57

20 I.L.[dB]

III. Experimental result
A. Effect of propagation direction
The layer mode was sensitive to the orientation ofsapphire substrate, namely, propagation direction ofwaves. Figure 1 and 2 compares dependence of
frequency responses on propagation directions. In thecase of device with the propagation direction along<11-20>, the layer waves are relatively well confinedin a film, because the effective velocity in sapphire- 221 -

layer

SAW

40
60
80
0

1.0

2.0

Freq. [GHz]b) Path length=5mm LS=0.8um kH=3.93
Figure 2 Response for <1-100> prop. direction

C. Multi-mode condition
Figure 4 shows pass band characteristics of layer
mode device with the propagation direction of<11-20>. The devices have relatively wide pass bandcharacteristics and group delay characteristics withnegative dispersion (delay decrease with frequency
increase). These effects could not be estimated by thesimple mode calculation. We can explain responses
reasonably, if we would define that two layer modes
propagates along the layer with different propagation
velocity and it couples on the propagation path. In thisoperation, the frequency response detected by thesame output transducer varies depending on the
propagation path length from input transducer. The
center frequency, at which maximum output signal is
obtained, also varies, as shown in the figure 3. In
addition, negative dispersion appears by the effective
propagation path reduction caused by the interference
effects of two modes. By controlling propagation path
length and mode interaction, which depends on film
thickness, we can design devices with a wideband
frequency response.

I.L.[dB]

40

4.8

3.2
60
1.6

80
0.8

0.85

0.9

0.95

1

GroupDelay[us]6.4a) Path length=2.5mm0
Freq. [GHz]I.L.[dB]-40b) Path length
=5mm

1.0

-50
0.8
-60
0.6

-70

-80
0.8

0.9

1

GroupDelay[us]B. Single mode condition
Experimental result have shown that the single
mode operation was stably obtained at the conditionsof kH values 1 and 4 for propagation direction of<11-20> and <1-100>, respectively. This condition
agrees well with calculation values. At these single
mode conditions, the frequency responses are
relatively simple. Therefore, we can design by asimplemethodjust
includingdispersion
characteristics on conventional design method used
for SAW devices.

4.0
Freq. [GHz]

L&S=2um, Thickness 2um, 50-50 pairs IDT
Figure 3 Multimode responses along <11-20>
20 I.L.[dB]4060fIV. Conclusion
We have experimentally studied layer mode devices
on GaN/sapphire substrates. The results clarified thatthe device is useful for various high frequency devices
at the condition of single mode operation. At themultimode condition, layer mode devices showed arelatively flat and negative dispersion characteristic.
The devices also have a mode coupling effect, which
would be effective for realizing high speed and wideband devices. Future works should be focused on the
improving reproducibility.
Acknowledgement
The authors wish to thank Associate professor.
Koh for his valuable discussion. This work is
supported by the fund for Advance Technology Centerfrom the Ministry of Education, Culture, Sports,
Science and Technology.

- 222 -

80
0.8

0.9

1

Freq. [dB]

Figure 4 Response of multi mode device withwide bandwidth
References[1] Y. J. Kim, G. H. Lim etc. “SAW Characteristics
of GaN Epitaxial Films Deposited on Different
Plane,” Proc. IEEE Ultrasonic Symp. 1 427-430(2002).
[2] K. Hohkawa, C. Kaneshiro, K. Koh, K.
Nishimuru and N. Shigekawa, “Study on
Photo-induced Acoustic Charge Transport Effectin GaN Film” IEEE MTS Symp. WE1B, June 15,
2005.
[3] N. Shigekawa, K. Nishimura, H. Yokoyama,“Side gate effect GaN SAW device” USE 2004
Proc. 2-17, pp. 191-192 (2004).

P2-6

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 223-224
16-18 November, 2005

Acoustic Wave Device Using GaN film with n+ Conduction Layer
Takahiro Mizusawa, Masaya Wada, Keishin Koh, Kazumi Nishimura*, Naoteru Sigekawa*, and Kohji Hokawa
Kanagawa Institute of Technology, ＊) NTT Photonics Laboratories
1. Introduction
By using epitaxially grown highly conductive
semiconductor layer, deposited on the insulating
semiconductor film, such as GaAs, and GaN, as
electrode elements, it is possible to fabricate IDT’s
without metal electrodes [1], [2]. In addition, if weplace the conductive film under layer in multiple
epitaxial films, we could obtain new types offunctional devices. This paper studies SAW devices
having electrode elements consisting of epitaxiallygrown GaN film with highly conductive layers, andclarify the potentials of this structure.

We fabricated two kinds of device; namely, 1) n+layer on the surface; IDT is formed by etching n+layer as shown in Figure 1, and 2) n+ layer under the
piezoelectric layer; metal electrodes are formed on thesurface of piezoelectric layer. In this structure, we canestimate two kinds of devices, namely single and dual
phase metal electrodes as shown in Figure 2.
n+ GaNGaN
Al2O3 substrate

Fig.1 ＩＤＴ with surface n+ layer

2. SAW devices with n+ layer electrodes

Electrode

Since n+ layer has the same nature as that of
conductive metals, its role in SAW device is to form
electrodes pairs and some kinds of control electrodes.
There are two basic structures for applying electrodesof transducer. One is a surface electrode, which is thesame as that of conventional metal electrodes of IDT.
The other is buried electrode, which would be difficultforconventional metal electrodes to growpiezoelectric film with single crystal semiconductor
film with a crystallography good enough to form highperformance semiconductor devices. This element
makes it possible to control and detect electric field.
This nature would lead on SAW devices many
improvements stated as follows.

GaN
Al2O3 substratea) Dual electrode
ElectrodeGaN
GGa
Al2O3 substrate

Table I. Parameter of test device
Substrate
GaN on Al2O3
Thickness of GaN film 2 µm
Thickness of n+ layer
400 nm
IDT line &space
2µm:2µm / 0.8µm:0.8µm
Overlap. length
800 µm / 400 µm
Number of pairs
50-50 pairs
Prop. length
2.5 mm2) This structure allows monolithic fabrication oftransducer with transistors using the same layer
structure (co-integration).
3) Co-integration would leads new functional device
utilizing semiconductor.

genni@ele.kanagawa-it.ac.jp

n+ GaNb) Single phase typeFig.2 Test devices with buried n+ layer

1) We can improve performance of semiconductor
SAW coupled devices, such as convolver andprogrammable devices.

4) This enables transducer with lateral field excitation,
which extends the fabrication limit of pattern size.
Here, we report some experimental results using
epitaxially grown n+ GaN film.

n+ GaN

3．Experimental results
Table I lists device parameters used for theexperiments. We have fabricated a device with IDTs
consisting of n+ layer with carrier densities larger than
1018/cm3. We have used Cl2 gas dry etching process- 223 -

S11[dB]

‑1

‑2

‑3
0

1
Frequency[GHz]2

S12[dB]

‑20

SAW

‑40

Layer

‑60
‑80

0

1
Frequency[GHz]2a) Line & space 2µm
S11[dB]

0
‑1
‑2
‑3

0

1
Frequency[GHz]

‑20

2

SAW

‑40

S12[dB]for patterning. Compared to the metal electrode, theresistance of semiconductor electrodes is slightly large,
but response was almost same as the device with usual
metal electrodes. The precise result is reported inreference 2. Figure 3 illustrates experimental result ofa device with buried n+ layer device. Compared to the
responses of conventional (without n+ layer) device,this device is a comparable insertion loss to that ofconventional device at a high frequency range.
However, at a lower frequency the insertion loss was
relatively large. This difference is mainly caused by
the difference in amplitude distribution in the depth
direction and also the conductance of layer mode. We
also observed the direct feed through can not be
reduced even if we connect n+ layer to the groundlevel. This is also mainly caused by the direct feedthrough signal path through the n+ layer. This problem
is also solved if we decrease resistance of the layer.
The energy loss in lower frequency device is mainly
caused by the resistance of n+ layer which closes
critical conductance of propagation SAW. The loss
would be reduced, if we could enhance conductivityof n+ layer. For the higher frequency device, since thewavelength is comparable with the film thickness andthe large fraction of electric field is effectively applied
to the film surface and the fraction of electric field inn+ layer become to be weak. (ref. Fig.3 (e) and Fig.3
(d)).

Layer

‑60
‑80

0

1
Frequency[GHz]2b) Line & space 0.8µm

4. CONCLUSION
Experimental result confirms the effectiveness ofn+ layer for transducer electrodes. For buried structure
the layer, it would be required much more highly
conductive layer to isolate from electromagnetic
coupling. We have now fabricating a device withcircuit structure as shown in Figure 2 (b). The results
would be presented in the symposium.

Current[μA]

1
0
‑1
‑2
‑3

‑4

0
Voltage[V]4(c) I-V property of IDT with n+ layer (λ=3.2µm)

REFERENCES
1）N. Shigekawa, K. Nishimura, H. Yokoyama and K.
Hohkawa; Appl. Phys. Lett. 87 084-102 (2005).
2) N. Shigekawa, K. Nishimura, H. Yokoyama and K.
Hohkawa; Technical Report of IEICE Oct. (2005).

- 224 -

Current[μA]1

ACKNOWLEDGEMENT
The authors wish to thank Dr. Yokoyama for help
with growth of GaN film and Dr. Kaneshiro for her
discussion. This work is supported by the fund for
Advance Technology Center from the Ministry of
Education, Culture, Sports, Science and Technology.

0
‑1
‑2
‑3

‑4

0
Voltage[V]4(d) I-V property of IDT with n+ layer λ=8µm)
Fig.3 Frequency characteristics of device withstructure as Fig. 2 a)P2-7

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 225-226
16-18 November, 2005

Evolutional Research on the Second Harmonicin 50MHz Band Nonlinear Surface Acoustic Wave

Yoshiaki Tokunaga, ○Atsushi Yasuno, Masaki Suzuki (OEDS R&D Center of Kanazawa Inst. Tech.)

1. Introduction
Recently, a barrier among cognitive science, life
science, human-social science and nature science has
been lowered. Besides, the new academic region called
new science (chaos, fractal, nonlinearity, and so on) has
attracted lots of interest. In this trend, some researchersattempt to reconsider phenomena in a piezoelectric
crystal from a point of view of the complexity1). As tothe surface acoustic wave generated on the piezoelectric
substrate, its behavior easily shifts from a linear region to
Fig. 1a nonlinear region because most of its energy is trapped

Diagram of the experimental sample.

in a depth of a wavelength from the surface of thesubstrate.

Therefore,technologicallyinteresting

been observed. We have studied the complexity of the

3. Results and Discussions
First, the measurement results of velocities of thenonlinear surface acoustic wave (NLSAW) from a point

fundamental and the 2nd harmonic are described. Figureof view of technological application2)-3). Through these2 shows the relationship between the propagation timendstudies, we discovered that the second harmonic (2and the propagation distance of SAW (LSAW) in theharmonic) in 50MHz band NLSAW had soliton-likecharacteristics. In this paper, we comprehensively reportlinear region. The optical probe was irradiated on thesubstrate surface at 500μm intervals in a row, and theexperimental results which we have found out so far.

temporal process of LSAW was measured. Obtained data

phenomena which do not emerge in a linear region havewere collinear approximated using least–square method,
2. Sample and Measuring Method
128 ° Y cut X propagation crystals were used as

and its velocity was estimated 3996m/s by the slope ofline. This velocity agreed fairly with the theoretical value.

LiNbO3 substrate. IDTs were regular type with 20 pairs.
20

because the dissipation between the wave and the18thermal phonon bath in the substrate was negligible, andits system could be treated as a conservative system.
Figure 1 shows a diagram of the experimental sample. A
20 mm free space was interspaced between IDT (I) andIDT (II) in order to do light scattering measuring and toevaporate a metal film.

Dista nce From IDT (I) [mm]The center frequency of the IDT was around 50MHz

04 Y-0 5

16
14
12
10
8
6
4
2
0
0

39 96 [ m/s]
1
2
3
4
5
Prop a gation Time [µs]6-----------------------------------------------------------E-mail: toklab@mlist.kanazawa-it.ac.jp

Fig. 2- 225 -

Velocity of SAW in the linear region.

Secondly, when the input signal was increased, NLSAW

metal thin film was few% bigger than the case thatwas generated, and its velocity was measured in the same

NLSAW did not experience the metal thin film.

way. Figure 3(a) and (b) show velocities of theCorrespondingly, the 2nd harmonic was not affected byfundamental and the 2nd harmonic, respectively. The

the metal thin film. In other word, it was conservative.

velocity of the fundamental in NLSAW was 3994m/s,

According to these facts, we predicted that the 2nd

and it was few m/s slower than that of LSAW.

harmonic might be a kind of soliton. In order to confirm

Correspondingly, the velocity of the 2nd harmonic wasthis prediction, a collision experiment of two 2nd

4000m/s, and it was about 4m/s faster than the velocity

harmonics was attempted. Figure 6(a) and (b) showof the fundamental in LSAW. This tendency was

intensity of observed scattering lights when two

observed in many measured samples.

NLSAWs which were launched from two spatially

14
12
10
8
6
4
2

1
2
3
4
5
Pr op agation Tim e [µs](a) Fundamental.

collision, the difference of intensities could not be

14

estimated. On the 2nd harmonic, it was confirmed that its

12
10

intensity was algebraic sum of two waves at the collision8

point.
gold th in film

64
2

39 94 [ m/s ]00

IDTs. Because the fundamental lose its shape by the

16

6

0
0

40 00 [ m/s]
1
2
3
4
5
Pr op a gation Tim e [µs ]6(b) 2nd harmonic.

Velocities of the fundamental (a) and the 2nd

Fig. 3

harmonic (b).

gold th in film

10 -2
Collision

10 -3

10 -4

10 -5
IDT (I)
IDT (II)

10 -6
0

4

8

12

16

IDT(I)

Figure 4 shows a comparison between full width at

IDT (II)Distance From ID T (I) [m m ](a) Fundamental.

ndhalf-maximum (FWHM) of the fundamental and the 2harmonic in NLSAW. The FWHM of the fundamental
was about 1000μm, on the other hand, the FWHM of

20

Normalized D iffracted Light Intensity

Distace From IDT (I) [mm]181604 Y-0 6

Normalized Diffracted Light Intensity

04 Y-0 6

18

Fig. 6

10 - 2Collision

10 - 310 - 410 - 5
IDT (I)
IDT (II)

10 - 6
0

4

8

12

16

IDT (I)

20IDT (II)Distance From ID T (I) [mm]nd(b) 2 harmonic.

Collisions of two fundamentals (a) and two 2nd

harmonics (b).

the 2nd harmonic was about 500μm, and it was about
Conclusion
Characteristics of the fundamental and the 2nd harmonic

half of the fundamental’s.
2 0 00FW HM [ µ m]

Distance From IDT (I) [mm]opposite IDTs collided head-on at the middle of two

20

20

in 50MHz band NLSAW on LiNbO3 substrate were

Fu n .
2n d.

investigated by the optical probing method.

1500

It was found that the 2nd harmonic in NLSAW
substantially exhibited soliton-like characteristic.

10 00

Acknowledgement
This work was supported by “High-Tech Research

500

0
0

Fig. 4Center” Project for Private Universities: matching fund2

4
6
8 10 12 14 16 18 20
Dista nce From IDT (I) [ mm]subsidy from MEXT, 2000-2005.

FWHM of the fundamental and the 2nd

harmonic.

References
1) Y. Tokunaga and Y. Kugo: Jpn. J. Appl. Phys. 39(2000) 2933.

By the way, what will happen if NLSAW experiences ametal thin film? We compared FWHM between the case2) Y. Tokunaga, T. Okayama, et al.: Proc. Symp. Ultrason.

that NLSAW experienced the metal thin film and theElectron., 24, (2003) 131.

3)

case that NLSAW did not experienced it. The value of3) Y. Tokunaga, T. Nakaya, et al.: Proc. Symp. Ultrason.

fundamental in the case that NLSAW experienced theElectron., 25, (2004) 263.

- 226 -

P2-8

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 227-228
16-18 November, 2005

Application Limitations of Mason Equivalent Circuit as a
Linear Coupling Circuit
Michio Ohki, Natl. Def. Acad., Japan

1. Introduction
Mason’s equivalent circuit has a kind ofsymmetric structure for representing two different
coupling effects, longitudinal (L-) effect andtransverse (T-) effect, and it is expected that thisstructure works successfully when the piezoelectric
coupling coefficient is not so large; that is, in thelinear region. This study indicates the application
limitations of Mason’s equivalent circuit as a linear
coupling circuit in a semi-quantitative manner.
2. Symmetrical behavior of −C0 and +C0
In Mason's circuit, two kinds of electrical
capacitance components, −C0 and +C0, are included,
as shown in Fig. 1. Two switches SW− and SW+ are
included in the circuit to explain the behavior of−C0 and +C0 in a symmetric manner. There are fourcases with regard to the behaviour of SW− and SW+:(i) (SW−, SW+) = (open, open), (ii) (open, short), (iii)(short, short), and (iv) (short, open).
Cases (i) and (ii) correspond to the L-effect, andcases (iii) and (iv) correspond to the T-effect inwhich –C0 is shorted out. Cases (i) and (iv)represents the situation of infinitely large internal
resistence in electric source, and cases (ii) and (iii),
that of infinitely small internal resistence.

SW−
Y'

Y

In case (i), the resonance frequencies of the total
system, ωres, equal those of the mechanical part ofthe system, the admittance of which is denoted by Y,
and those values correspond to the anti-resonancefrequencies, ωA, of the electrically observed
admittance denoted by Y’. In this case, −C0
apparently disappears due to the offset to +C0 on
the circuit, and the coupling between –C0 and C
does not occur, which is the intrinsic state for
L-effect. On the other hand, in case (ii) −C0
emerges and is coupled with C, causing the changein ωres from ωA to ωR that is the resonancefrequencies of Y’. In both cases, the behavior of −C0
determines the resonance of the total system, while+C0 does not work.
The cases (iii) and (iv) in the T-effect correspondto cases (i) and (ii) in the L-effect, respectively, but
the behavior of –C0 and +C0 is interchanged; that is,the case (iii) represents the intrinsic state of T-effect,in which the coupling between +C0 and C does notoccur, since +C0 is now shorted out, and ωres in thiscase corresponds to ωR of Y’; on the other hand, incase (iv), +C0 emerges and is coupled with C,
causing the change in ωres from ωR to ωA. In both of(iii) and (iv), the behavior of +C0 determines theresonance of the total system, while −C0 does notwork.
The symmetrical coexistence of −C0 and +C0 is
inevitable in the framework of Mason's circuit.
3. Behavior of Linear Coupling Circuit

− C0
SW+

+C0

We investigate the shift of the resonancefrequency of the total system due to the linear
coupling between −C0 and C, or between +C0 and C,
where the value of C can be calculated by1∂Y .
C = limj ω→0 ∂ωC

Fig. 1 Framework of Mason’s equivalent circuit with
dielectric components −C0 and +C0, and withequivalent elastic component C at the frequency ω0.
E-mail: michio_ohki@ieee.org

Let us consider the type of piezoelectrictransducer that is shown in Fig. 2.
Figure 3 shows the dependence of the difference- 227 -Electrode(a)

B' Z 1

Electrode

ZD

Z2

B"

Z2

B"

Electrode(b)

B' Z 1

ZD
Electrodev
v=1Fig. 3 Relationship between δω, which is thefrequency shift due to the linear coupling at thefundamental mode, and C/(C0 + C), which is a kindof linear coupling parameter, under the condition ofv = 1 and (B’, B”) = (free, free). (a) δω = ωR – ωAwith ωA = 1 for L-effect, and (b) δω = ωA – ωR with
ωR = 1 for T-effect.

Fig. 2 Piezoelectric transducer for (a) L-effect (b)T-effect. v is the volume ratio of the drive domain to
the whole transducer. B’ and B” are mechanicalboundaries set free or clamped. Z1, ZD, Z2 aremechanical impedance of the respective domains.

between ωR and ωA, δω, on C/(C0 + C) under theboundary condition of v = 1 with (B’, B”) = (free,free) for the fundamental mode, based on thecalculation using Mason’s circuit, in which (a) and(b) are the results for L-effect and T-effect,
respectively. In the case of L-effect, the value ofC/(C0+C) cannot exceed 0.5; that is, the value of C
never exceeds that of C0, although material
constants composing C and C0 are thought of as
being independent of each other. This unnatural
result is caused by the framework of Mason’s
circuit as a linear coupling circuit in which C and–C0 are linearly coupled. In this case, the total
capacitance for the case (ii), (C−1 – C0−1)−1, must be
finite and larger than zero. In a high coupling regionin which the satisfaction of the linearity is doubtful,
the mathematical result does not reflect the physicalsituation naturally. Therefore, Mason’s circuit isexpected to be completely reliable in the region of
C/C0 ` 1.
Since the important and essential factor for this
discussion is the ratio of C to C0, we investigate thedependence of the (normalized) value of C onvarious boundary conditions, which is shown in Fig.
4, in the cases of (a) (B’, B”) = (free, free) or (free,clamped) or (clamped, free) and (b) (B’, B”) =(clamped, clamped), both for Z1 = ZD = Z2
(impedance matching inside the transducer). In eachcase, the value of C only depends on the value of v;
that is, the value of C does not depend on thelocation of the drive domain in the transducer aslong as the volume ratio v is fixed.
In the case of (a) in Fig. 4, the value of C is in

Fig. 4 Dependence of the value of C (normalized)on the volume ratio v. (a) (B’, B”) = (free, free) or(free, clamped) or (clamped, free), (b) (B’, B”) =(clamped, clamped) for Z1 = ZD = Z2. (Thedependence of C on v is just identical with thedependence of the “density of resonance pattern”
defined in ref. 1 on v.)

proportion to the value of v, which means that thecondition of C/C0 ` 1 is satisfied more easily as v
becomes smaller, providing the value of C0 is fixed.
The value of C0 is actually in inverse proportion tov for the L-effect, as long as the material constant isfixed, while the tendency becomes the inverse for
the T-effect. In the case of (b), the situation is thesame as in the case of (a) in the region of v ` 1.
Therefore, we can expect that the application
limitations of Mason’s circuit as a linear coupling
circuit for L-effect can be removed completely inthe situation of v 0. For T-effect, nothing can beasserted with regard to the present problem in theframework of Mason’s circuit.

- 228 -

1) M. Ohki: Jpn. J. Appl. Phys. 44(2005) 4509.

P2-9

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 229-230
16-18 November, 2005

Effects of RF Filters on Performance of Power Amplifiers
J.G. Wen, H.L. Li, T.Omori, K. Hashimoto, and M.Yamaguchi (Chiba University)

2. Effects of RF filters on PA performanceFig. 1 shows a circuit used for the discussion. A
common emitter Class AB PA was designed to offer
optimal output power (Pout) and PAE at a frequncyof 1950MHz with 50Ω termination. The simulation
was carried out under Ansoft Designer version 1.1.
Vb

0

S 11

S (dB)RF power amplifiers (PAs) have widely beenused in various communication systems, and more
and more emphasis has been laid down on highpower-added efficiency (PAE) for extending abattery life and reducing heat generation in mobile
communication systems.
When RF PAs are operated near a saturation state
to get a better PAE, the nonlearity is not neglible,
i.e., harmonics may also affect PAE. RF filters are
employed for the suppression of unwanted EM
radiation and applied to the output network of PAs.
For this purpose, the filters based on surface
acoustic wave (SAW) and film bulk acoustic
resonator (FBAR) technologies are widely used.
Independent of the technologies employed, however,the filter function is realised by the frequency
dependence of the impedance mismatch with
peripheral circuits. This suggests that the input
impedance of the filters may badly affect PAE aswell as the transmission characteristics of the output
network.
From this point of view, this paper discusses
how the impedance characteristics of an RF filter
affects the PA performance when it is inserted
between the PA output and load.

is: G=12 dB and P1dB=12.7 dBm at Pin=1.5 dBm.
An RF FBAR filter for the W-CDMA Tx band
(1920-1980MHz) developed by Fujitsu ResearchLabs was used in the simulation. Fig. 2 shows thefrequency response of the filter, where the insertion
loss and reflection coefficient are 1.7 dB and -8.0
dB, respectively, at 1,950 MHz.

-10

-20

S 21
-30

-40

-50
1.6

1.7

1.8

1.9

2.1f

2.0

(G H z) 2.2

Fig. 2 Frequency response of FBAR filter used in
the disucssion
2.1 Effects of S11
Fig. 3 shows Pout as a function of the input power
Pin. When the FBAR filter is short-circuited (CaseA), Pout is almost proportional to Pin when Pin is lessthan –2 dBm. For Pin>–2 dBm, Pout increases more
rapidly and saturates at Pin≈ –1 dBm.
14

Pout (dBm)

1. Introduction

13
12
11

without filter

10Vc

with filter S11 suppression

9

with filter

8matching

matching

7
6

Pin (dBm)

5

Zeff

-6

Fig.3
Fig. 1 Circuit topology
As a transistor, NEC 2SC5751 was employed andbiased at 3 V and 10 mA. Note in the following
simulation that matching elements are assumed to
be lossless to clarify the effect of the filter on the PA
performance. The performance of the designed PA

-5

-4

-3

-2

-1

0

1

2

3

Pout as a function of Pin

If the filter is inserted (Case B), Pout becomes
smaller than that in Case A. For Pin <–2 dBm, thedecrease in Pout is mainly caused by S21 of the filter,and relatively small (≈ –1.1 dBm). However, thedecrease becomes much larger than that expected only
by S21 when Pin>–2 dBm.
The figure also shows the calculated Pout when S11- 229 -

significantly improved by the harmonic tuning at
the region where the transistor is almost saturated.
The series of the calculation indicate that this PAE
enhancement is due to the lack of power
consumption for the harmonic components.
14

Pout (dBm)of the filter is set at zero (Case C). Compared Case C
with Case A, it is seen that the decrease in Pout isalmost independent of Pin, and that the decrease seems
to be simply determined by S21. This suggtests that S11.
should significantly affect Pout near a saturation stateof PA.
To investigate the result shown in Fig. 3 in detail,the dependence of Pout on S11 was calculated.
Although the filter performance used in thecalculation is not exactly identical with that shownin Fig. 2, Fig. 4 shows how S11 affects Pout when
S11=–5, –15 and –25 dB.

13
12

(1)
(2)

11
10
9

13
8

Pout (dBm)

12
11
10
9

45.2-j2.3
735.9-j5.9

Pin (dBm)
6-6

14.7-j10.8

8

-5

-4

-3

-2

-1

0

1

2

3

4

Fig. 4 Pin dependence of Pout. (1) designed underharmonic-tuning,and(2)
designedto50Ω termination.

7
6
5
4

0.8

PAE

3
2

0.7

1

Pin (dBm)

0
-10

-9

Fig. 3

-8

-7

-6

-5

-4

-3

-2

-1

0

1

0.6

2

Influence of ZLof the filter to Pout

(1)
(2)

0.5

The result clearly shows that S11 considerably
reduce Pout near a saturation state. Hence, avariation in S11 of RF filters must be suppressedsufficiently when they are used with RF PAs.

0.4

0.3

0.2

2.2 Influence of harmonics

Pin (dBm)-6

It is known that a proper harmonic tuning canincrease the output power[2]. From this, it isexpected that S11 at the harmonic frequencies also
affects the PA performance.
Usually, RF filters exhibit reactive input
impedances at the rejection bands including
harmonic frequencies. Then the real part of the load
impedance Zeff(ω) for the transistor looking toward
the output matching circuit (see Fig. 1) can beideally zero at harmonic frequencies even when the
RF filter is connected. In contrast, it is usuallynonzero at harmonic frequencies when the matchingcircuit is simply terminated by a resistive load suchas 50Ω.
Fig. 5 shows the calculated Pin dependence of Pout
when Zeff(ω) for second and third harmonicfrequencies are set at a suitable pure imaginary
value. For comparison, the result for a 50Ω load isalso shown. The result shows that the difference
between the two cases is small.
Fig. 6 shows the calculated PAE as a function of
Pin for these two cases. In contrast to the result for
the Pout characteristics shown in Fig. 5, PAE is

-5

-4

-3

-2

-1

0

1

2

3

4

5

Fig. 5 Pin dependence of PAE. (1) designed underharmonic-tuning, and (2) designed to 50Ωtermination.
3. Conclusions
It was shown that S11 of an RF filter has asignificant
impacton
power
amplifier’s
performance, in particular, near a satuaration state.
The effect of S11 upon PA performance at harmonicfrequencies was also discussed.
Acknowledgment
This work was partially supported by the
Grant-in-Aids for Fundamental Scientific Researchfrom the Ministry of Education, Sports, Science andCulture.
References[1] S.C.Cripps: RF Power Amplifiers for WirelessCommunications, Chapter 2, Altech House (1999)[2] P.Colantonio,P1639, IEEE MTT-S Digest, (2002).

- 230 -

P2-10

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 231-232
16-18 November, 2005

SAW-Semiconductor UV Sensor Using GaN Film
Keishin Koh, Chinami Kanashiro, and Kohji Hohkawa, Kanagawa Institute of Technology,
Kazumi Nishimura, Naoteru Shigekawa NTT Photonics Laboratories
1. Introduction
GaN material has good semiconductor nature; it not
only is used to realize various optical devices such as LD,
LED and photodiode, but also it is used to make various
electronic devices with high temperature and with highspeed. On the other hand, GaN material also haspiezoelectric nature; surface acoustic wave (SAW) devices
such as filter, oscillator have been studied [1]. Comparing
with Si, GaAs material, GaN material has many functional
properties and it is expected to realize novel highly
functional devices with monolithic structure. In thispaper, we will focus on the charge transfer effect in GaN
film and investigate basic characteristics of GaN
SAW-semiconductor UV sensor based on the transfer
effect of charge induced by UV light.
2. SAW-semiconductor UV sensor
Recently, it is strongly required to develop highperformances UV sensor. Comparing with UV sensor withvacuum tube structure, the semiconductor UV sensors
have better feature such as low cost, low noise and small
size. Among them, GaN material is used as material ofsemiconductor UV sensor since it is wide band energy gap
and the UV sensor with MSM structure has been studied[2]. On the other hand, the relationship between traveling
SAW and induced carriers in GaN film also has beenreported [3]. We also confirmed transports effect of
induced carriers by traveling SAW and investigated basic
phenomenon. These experimental results indicate
possibility of realizing high functional UV sensor using
GaN film. Figure 1 shows a cross sectional view of
SAW-semiconductor UV sensor device. The GaN film isdeposited on a non-piezoelectric substrate (Al2O3). When
UV light is irradiated on propagation path and carriers(electron and hole) are generated in GaN film, if SAW is
not generated and there is not the electric filed in the film,

Input SAWUV light

OutputGaN
Al2O3 Substrate
Fig. 1 SAW-semiconductor UV sensor..

so no carriers can transfer. When SAW is generated and theelectric filed be formed in the GaN film. The induced
carriers are transferred by propagation of potential wellcaused by traveling SAW and the output signal is detected
at output IDT as MSM photodiode with Schottky junction.
This UV sensor has feature as following.
1) It has high sensitivity by traveling SAW and has
selectivity of UV wavelength2) The output signal could be selected from DC, amplitude
or phase RF signal. Moreover the output signal could beremote controlled.
Here, we report experimental results of the SAW-semiconductor UV sensor using insulating GaN film.
3. Basic experimental results
We deposited the GaN film on the Al2O3 substrategrown by MOCVD and fabricated test devices on GaN
films. Table 1 shows parameters of test devices. Figure 2(a) shows the frequency response of the test device, there
are two peaks corresponding to SAW mode and Sezawa
mode respectively. Figure 2(b) shows I-V characteristics ofoutprt IDT as a MSM diode, light current caused byinduced carriers is observed under UV irradiation.
Table I. Parameter of test device
Substrate
GaN on Al2O3
Thickness of GaN film
2 µm insulating film
IDT line &space
2 µm: 2µm / 0.8µm:0.8µm
Overlap. length
800 µm / 400 µm
Number of pairs
50-50 pairs
Propagation length
2.5 mmElectrodeAl, 100 nm--------------------------------------------koh@ele.kanagawa-it.ac.jp- 231 -

Variable Att.

-20
Wafer2 2.5mm

Layer

Insertion Loss [dB]

-30
-40

MirrorSAW

UV
source-50

Output
A/V

RF

-70

Test device-80Fig. 3 Experimental set-up

-90
0

200

400

200

600 800 1000 1200 1400
Frequency [MHz]
(a)

1000

f
f
o
V
U

100

-15dBm

2

µ

0

0dBm
-5dBm

m
c
/
W

4
.
2
V
U

100

-25dBm10

f =579.4 [MHz]

-100

0

1
0.1

-10

-5

)
V
(
e
g
a
t
l
o
V

-200

Bias

Slit

-60

0

5

1

10

(b)

4. Conclusion
We confirmed the charge transfer effect on the insulating
GaN film under UV irradiation and investigated basic
characteristics of SAW-semiconductor UV sensor. From aview point of the device application, it is necessary toinvestigate other basic characteristics such as frequency

100

2

(a)

250

Fig. 2 Basic characteristics of test device. (a) Frequency
response, (b) I-V characteristics of IDT.

Figure 3 shows configuration of experimental setup. The
UV lamp with 360nm peak was used. We measured thecharge function of GaN SAW filter under UV irradiationon the propagation path. Figure 4(a) shows relationshipbetween DC output signal and UV irradiation intensity
under different input SAW power. The DC output almostis proportional to UV light power and weak UV light withmicrowatt order could be detected. It indicates that the
charge transfer efficiency is relatively higher than the
conductive GaN film. In additional, as the SAW powerincrease, the DC output level also increases. Fig. 4(b)shows relationship between the DC output and IDT bias
voltage, we can find it is effective to enlarge DC output
signal.
In experiment, output signal is not stability due tocharging up effect. We measured DC output signal at thecondition which transferred carriers is completely emitted
In future, it is important to design a optimum device
structure for emitting carrier.

10

UV power ( µW/cm )Input SAW power 5dBm

200
150

0dBm

100

-15dBmf =579.4[M Hz]

50
0

0

UV power: 0.24 µW/cm
4

5

6

7
8
Bias Voltage ( V)

9

2

10

(b)

Fig. 4 Basic characteristics of UV sensor.
dependence of film, dependence of UV wavelength.
Moreover, we have to find way exactly to control charges
in film and study optimization of device structure.
Acknoledgement
The authors wish to thank Dr. Yokoyama for help withgrowth of GaN film and Mr. Mizusawa for theircontributions on measurements. This work is supported by
the fund for Advance Technology Center from the Ministryof Education, Culture, Sports, Science and Technology.
References
1) 1. O. Ambacher ， Appl. Phys., vol. 31, pp.
2653-2710, 1998.
2) D. Ciplys, R.Rimeika, A. Sereika, R. Gaska, M. S.
Shur, J. W. Yang and M. A. Khan, Electron. Lett., vol.
37, no. 8, pp. 545-546, April 2001.
3) T. Palacios, F. Calle, J. Grajal, E. Monroy, M.
Eickhoff, O. Ambacher, and F. Omnès, Proc IEEE
Ultrason Symp, pp. 57-60 2002.
4) K. Hohkawa, C. Kaneshiro, K. Koh, K. Nishimuru
and N. Shigekawa, IEEE MTS 2005, to be pubrised.
- 232 -

P2-11

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 233-234
16-18 November, 2005

Simple Ultrasonic Anemometer
Using a Bended Sound Probe◎Kosuke KUDO, Koichi MIZUTANI1 and Masahisa ISHII2
1
2

Graduate School of Systems and Information Engineering, University of Tsukuba
National Institute for Rural Engineering

1. Introduction
A heat exchange between some spaces isperformed via air usually. Winds are ones of theimportant parameters for advanced environmental
controls(1-2). A whole motion of the wind is estimatedsimply. However, the winds change fluxionally. An
adequate value should be estimated from several
measurement results at several points during a brief
period. Ultrasonic anemometers rapidly measuremean winds in an area(3-6). Conventional ultrasonicanemometers have applied two pairs of transducers
at least. Sound reflections have been applied toreduce the transducers7-8). The sound reflection
allows the ultrasonic anemometer to reduce thetransducers(9). However, the ultrasonic anemometerwith the sound reflection was a symmetric shape.
The symmetric shape caused some problem tomeasurement.
This paper described a simple ultrasonicanemometer using a bended sound probe. The
bended sound probe allowed the ultrasonicanemometer to measure the wind with a pair of loudspeaker and microphone.

2. Principle
A schematic diagram of a simple ultrasonicanemometer is shown in Fig. 1. Two sound probes

consist of two propagation paths by a loud speaker(SP), microphone (MIC) and acoustic reflector. A
bended sound probe is composed by the acoustic
reflector. Each device is installed on the apexes of an
inequilateral triangle with 30o and 60o. The MIC
receives two times of flights (TOF) of the sounds
from the SP. The wind velocity is decomposed to x-ycomponents; wx and wy. Using the wx and wy, theTOFs along the direct path, t1, and reflected path, t2,
are expressed as follows:
L1
,
(1)
t1 =
c + wx
L3
L2
, (2)
t2 =
+
c + w x 2 + 3w y 2 c + 3w x 2 − w y 2
where L1, L2 and L3 denote lengths of propagation
paths, c sound velocity, respectively. Two wy arederived from eq. (2). Figure 2 shows typical tracesof eq. (2) with L1 = 2.0 m, L2 = 1.0 m, L3 = 1.7 m, w
= 5.0 m/s and c = 347.3 m/s. θ are 20, 40 and 60
degrees. The two wy from one TOF indicate different
values by biases of central axes. A correct wychanges by the wind direction. Estimations of wy are
difficult in the space with extreme changes of thewind. However, the wind direction changes
continuously. A correct wy is estimated using theprevious wind velocity and direction in the rapid
measurement. Wind velocity, w, and direction, θ, are
L1

wx

θ

wy
w

x

wx t1

ct1
30

o

60

y

ο

wx’ t22ct21x’
wx’ = wx - 1 wy
2
2
3

ct22

L2

L3
wy’ t21w
wy’ = 1 wx + 3 wy
2
2
y’

: SP
: MIC
: Acoustic reflector
velocity vector: Soundunder wind

Fig. 1 Parameters for wind measurements. Wind vector is decomposed to two patterns. Apparent sound velocities are composed of soundvelocities and wind velocities along propagation paths.

- 233 -

3. Simulation
Simulations were performed. Parameters were L1= 2.0 m, L2 = 1.0 m, L3 = 1.7 m, w = 5.0 m/s and c =347.3 m/s. A sampling frequency of the TOF was
100.0 MHz (0.01 µs). Incident angles of the winds
were changed from 0o to 360o. Figure 4 shows tracesof ideal and simulated wx and wy. wx was measured
with slight errors. wy indicated some errors from theideal wy. The maximum error was 0.2 m/s at 20o.
Figure 5 shows the wind velocity and direction. w
indicated some errors from ideal w. The maximum
error was 0.1 m/s. The used TOF were digital values.
Half adjusts of ideal TOF caused the errors. In thissituation, 0.01 µs difference of TOF corresponded to
0.1 m/s difference of wy.

1.6

Time of flight through reflected path ( µ s)

derived from the wx and wy.

20 degrees
40 degrees
60 degrees

0.4

-0.8
-5.0

-2.5

6.0

Wind velocity (m/s)

wx
wy

Ideal wx
Ideal wy

180

270

3.0

0.0

-3.0

-6.0

0

90

Fig. 3 Simulated wind velocities of x-y component (wx and wy). A
sampling frequency of TOF was 100.0 MHz (0.01 µs). Ideal
wx and wy are indicated also.

6.0w
Ideal w

5.0

4.0
360
θ
Ideal θ

180

0

0

90

180

270

Incident angle of wind (deg)Fig. 4 Simulated wind velocity and direction.

- 234 -

360

Incident angle of wind (deg)Wind velocity (m/s)

1) S. Wang, M. Yernaux and J. Deltour: J. Agric. Engeng Res.,73 (1999) 189.
2) E. Shilo, M. Teitel, Y. Mahrer and T. Boulard: Agric. Forest
Meteorol., 122 (2004) 3.
3) R. E. Corby: Electronics, 23 (1950) 88.
4) G. S. Campbell and M H. Unsworth: J. Appl. Meteoril., 18(1979) 1072.
5) K. Hamotani and N. Monji: J. Agric. Meteorol., 55 (1999)
109.
6) K. Mizutani, K. Taruishi, Y. Hachisuka, K. Kudo and M.
Ishii: Jpn. J. Appl. Phys., 43 (2004) 3099.
7) J. S. Kim and A. Yamada: Jpn. J. Appl. Phys., 41 (2002)
3592.
8) K. Kudo and K. Mizutani: Jpn. J. Appl. Phys., 43 (2004)
3095.
9) K. Kudo, K. Mizutani and K. Itoga: Jpn. J. Appl. Phys., 44(2005) 4407.

5.0

Fig. 2 Typical traces of eq. (2). L1 = 2.0 (m), L2 = 1.0 (m), L3 = 1.7
m and c = 347.3 (m/s). Wind velocity and direction were 5.0
m/s and 20, 40 and 60 degrees.

θ (deg)Acknowledgment
This work has been supported in part by aGrant-in-Aid for Scientific-Research (B) (No.
17380149) from the Japan Society for the Promotionof Science.

2.5

wy (m/s)

4. Conclusion
A simple ultrasonic anemometer using a bended
sound probe was described. A SP, MIC and acoustic
reflector were composed two propagation paths. In
the simulation, good agreement was indicated except
to errors caused from digital values. The windvelocity and direction would be measured with thesimple ultrasonic anemometer.

0360

P2-12

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 235-236
16-18 November, 2005

Measurement of Temperature Distribution
Using Acoustic Reflector Array◎Satoshi KAWABE and Koichi MIZUTANI (Univ. Tsukuba)
1. Introduction
In recent years, it has become necessary to control
temperature distribution of living space such as office
appropriately (1) from viewpoint of energy
conservation. Measurement of distribution ofenvironmental parameters using sound wave issuperior to contact-type sensors in the point that it canmeasure large area and it is nondestructive andnoncontact method. Therefore, many studies on themethod have been made (2-9). In case it is hard to
install a lot of acoustic sensors in measurement space,
the method using acoustic reflectors (hereinafter:
reflector) is very effective (9).
In this study, we constructed acoustic reflector
array and measured spatial mean temperature ofdivided areas by receiving the reflected waves. Further,
we compared the result of measurement with that ofthermocouples to evaluate the system.
2. Principle of measurement
Relationship between sound velocity and airtemperature is expressed as follows:T=Τ0c 2 − Τ0 .
331.32 2H

c1 T1

Area 2c2 T2

Area 3

c3 T3

L3
L2

L

(2)

2 L2 2 L 2
+
.
c1
c2

(3)

Eq. (3) is arranged as follows:

L1
Reflector1

2 L1
.
t1

where L1 is the distance between the center of SP and
the center of reflector1. Substituting c1 for eq. (1), thespatial mean temperature, T1, is obtained. In addition,using c1, t2 is expressed as follows:
t2 =where T (℃) is the spatial mean temperature along thesound propagation path ,c(m/s) is the sound velocity.
Measuring the propagated distance and the time offlight (TOF) of the sound, the sound velocity is
obtained. Thus the spatial mean temperature is
calculated from eq. (1).
Figure 1 shows the schematic diagram of measureArea 1

c1 =

(1)

Τ0 = 273.15Κ

Speaker andMicrophonement of temperature distribution using reflector array.
The unit with a loud speaker (SP) and a microphone(MIC) is installed at the higher position than reflector
array. Reflectors are arrayed at even interval and theheight of the center of each reflector is equal .The
angles of the reflectors are set so that theperpendicular of each reflector would intersect with
SP. The measurement space is divided into several
areas on the boundary of reflectors. It is supposed thatthe temperature is uniform in each area and the soundvelocity is uniform too. A torn burst is transmitted
from the SP and reflected by each reflector, andreceived by the MIC. Analyzing the cross correlation
between the transmitted signal and the received signal,
the TOF of the reflected signals, t1～ ti , are obtained.
The sound velocity in the Area1, c1, is expressed asfollows:

Reflector 2
L

Reflector3
L

Fig. 1 Schematic of a measurement of temperature
distribution using acoustic reflector array.
Temperatures and sound velocities are uniform
in each area.

−1

⎛ t
1⎞
c 2 = ⎜⎜ 2 − ⎟⎟ .
⎝ 2 L2 c1 ⎠(4)

In eq. (4), values included in right side are known.
Therefore, c2 is obtained and the spatial mean
temperature in area2 is calculated. Thus the spatial
mean temperature of each area is obtained
sequentially.
3. Experiment
Figure 2 shows an experimental setup. The heightof center of the SP is 450 mm higher than those of thereflectors. And three reflectors are arrayed at intervalof 900 mm. The sound wave used for measurement
was frequency modulated with 10± 2 kHz in 10
waves. Transmitting signal was generated by apersonal computer (PC) and was radiated from the SP
via an amplifier and a digital to analog (D-A)
converter (DAQ Card-6062E / National Instruments).

- 235 -

Normalized Amplitude

2.4m
Area1

2m

4m

Heat source

Area 2

Area 3

Thermocouples

4.2m
D/A-A/D
Converter

Data
LoggerPC

Fig. 2 Experimental setup.

4. Conclusion
The measurement system of the temperature
distribution using acoustic reflector array has been
proven to be valid for detecting the area of heatconcentration. Amount of heat measured by the soundprobe and that of the thermocouples have error within
2.5 – 6.5 %, and the experimental result caught the(a) Trnsmitted signal

0-11(b) Received signal

0-11(c) Cross correlation result
between (a) and (b)

0
-1
0

5

10

Time (ms)15

20

Fig. 3 Measured signals. (a) transmitted signal. (b)received signal. (c) cross correlation result
between (a) and (b).
35

Thermocouple
Sound Probe30

Temperature ( C)

The sound wave reflected by the reflector array isdigitized by the PC via the MIC, amplifier and ananalog to digital (A-D) converter (DAQ Card-6062E).
Figure 3 shows the waveforms experimentallymeasured by the system. Figure 3 (a), (b) and (c) aretransmitted signal, received signal, and crosscorrelation result between Fig. 3 (a) and Fig. 3 (b),respectively. In Fig.3 (c), three peaks show the times
when the transmitted signal reflected by the threereflectors reached the MIC.
Temperature distribution was experimentallymeasured in four cases. Figure 4 shows experimental
results. Figure 4 (a), (b), (c) and (d) were without heatsource, heat source was installed in the Area1, Area2
and area3, respectively. Rectangles are temperatures
measured with the sound wave and black dots showtemperatures measured with nine thermocouples.
Figure 4 shows that the temperature of the area where
the heat source was installed was higher than those ofother areas. To evaluate the measurement with soundwave, the product of the area length and thetemperature of the area was calculated. In this study,
the product is defined as ‘amount of heat’. For
example, amount of heat of Area1 in Fig. 4 (a) is
calculated with 0.9×22.7 = 20.43 (m･℃). Amount ofheat measured by the sound probe and those of thethermocouples were corresponding well in the area
where no heat source was installed. In the areas where
the heat source was installed, the differences were
2.5 % (Area1) 4.2 % (Area2) 6.5 % (Area3). These
differences were caused because the thermocouples
could not catch precipitous gradient of the temperature.
And the differences were caused because the soundvelocities were not completely uniform in
consequence of a vertical gradient of temperature.
However, the measurement system detected the areaof heat concentration adequately.

1

(a)

(b)

25
Heat Source

20
35

(c)

(d)

30
25
Heat Source

20

0

0.9

1.8

Heat Source

2.7 0

0.9

1.8

2.7

x (m)

Fig. 4 Temperature distributions measured by soundprobe and nine thermocouples.
actual temperature distribution well. The system has
the advantage that it can be used in life space easily,
and the data obtained with this system is useful forenergy conservation.
Acknowledgement
This work has been supported in part by aGrant-in-Aid for Scientific Research (B) (No.
17380149) from the Japan Society for the Promotionof Science.
References
1)
2)
3)
4)
5)
6)
7)
8)
9)

- 236 -

F. Engdahl and D. Johansson: Energy and Buildings, 36(2004) 205.
E. C. Shang and Y. Y. Wang : Computer Physics
Communications. 65 (1991) 238.
I. Basarab-Horwath and M. M. Dorozhevets: Proc. IEEE
Ultrasonics Symp (1994) 1891.
M. Fujii and X. Zhang: Proc. 4th World. Conf.
Experimental Heat transfer, Fluid mechanics andThermodynamics. Brussels, 1997, (1997)169.
K. Mizutani, K. Nishizaki, K. Nagai and K. Harakawa:
Jpn. J. Appl. Phys. 36 (1997) 3176.
A. Funakoshi, K. Mizutani, K. Nagai, and K. Harakawa:
Jpn. J. Appl. Phys. 39 (2000) 3107.
K. Ikeda: Jpn. J. Appl. Phys. 40 (2001) 3566.
E. Ishikawa and K. Mizutani: Jpn. J.Appl. Phys. 42 (2003)
5372.
K. Kudo, K. Mizutani, T. Akagami and R. Murayama: Jpn.
J. Appl. Phys. 42 (2003) 3189.

P2-13

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 237-238
16-18 November, 2005

Temperature Measurement Using Network Controlled
Acoustic Sensors◎Kyohei SAWAMURA, Koichi MIZUTANI, Kenichi KASHIWAZAKI and Ikuo

ODANAKA (Graduate School of Systems and Information Engineering, University of Tsukuba)
1. Introduction
In recent years, a thermal environmental
measurement is required in various fields to maintainan optimum environment for life space, farm,large-scale structure and so on1,2). The automatedmeteorological data acquisition system(AMeDAS)
given from the Meteorological Agency is not suitable
for the local meteorological data, so the measurement
space of the AMeDAS is too large and coarse.
Thermometer from the satellite can measuretemperature distributions widely, but it is influenced
by radiative heat. And the result of measurement isground temperature and too coarse for local data3,4).
Therefore, the acoustic sensor has been applied to themeasurements of distributions with a small number of
sensors without influence of radiative heat5-7). As
low-frequency can propagate over along distance,measurement of acoustic sensors can apply to large
space 8).
In this paper, the purpose is to form the network by
many acoustic sensors in the large space and to gettemperature data and distributions locally.
2. Principle
Schematic diagram of an acoustic sensor is shownin Fig. 1. In the measurement space, two acoustic
sensors are set facing each other at the both corner ofthe baseline. The measurement space along thebaseline is equally divided into N cells. In a unit cell,
Tn is an air temperature and sound speed cn isdetermined by air temperature distribution Tn.
ΔL=L/N are defined. The time of flight (TOF) of
the sound wave is described as

∆L L
= .
c
n =1 c nN

TOF = ∑

(1)

Thus, average of sound speed in all section is found.
Spatial mean temperature is obtained eqs. (1) and (2).

T =Τ0c 2 − Τ0 .
2
331.32(2)Here, T0=273.15K is absolute temperature. An
influence of wind is able to remove when sound wave
is transmitted two-way at the same time. In thisexperiment, the TOF is measured by cross-correlation
between the received signal and the transmitted signal.

L

Acoustic
Sensor 1

Acoustic
Sensor 2ǍLacoustic wave

c1

T1

c2

T2

cn

Tn

cN
TN

Fig. 1 Schematic diagram of an acoustic sensor

Fig. 2 Layout of the network measurement system
3. Experiment
3.1 Layout of acoustic sensors
Figure 2 shows the future model of network
sensing formed by acoustic sensors. As for each sensor,
the functions such as signal generating, recording andprocessing will be installed. And the acoustic sensors
are connected by wireless network. The wirelessnetwork is IEEE802.11b, which has the advantage ofbeing able to transmit over a long distance. Therefore,
measurement function is dispersed.
3.2 Experiment Setup
This time acoustic sensors had not equipped thefunction of signal generation, recording, andprocessing yet. So sensors were connected to thePC via an analog-to-digital (A-D) converter and adigital-to-analog (D-A) converter (DAQCard-6062E /National Instruments). Various functions were given
to the sensors and the network by wireless LAN(IEEE802.11b) was able to be formed by connecting
PC. As the first experiment, four acoustic sensors are
set at the corner of the 10×10 ㎡ square in the open
air and are formed wireless network. Figure 3 is thecomposition of the experiment. For separating fromthe signal received on each acoustic sensor, two kindsof signals were used. One signal was a ten-wavelength
burst sound modulated from 4.5 to 1.5 kHz in linear
frequency at sensor 1 and 2, the other was used aten-wavelength burst sound modulated from 1.5 to 4.5- 237 -

kHz in linear frequency at sensor 3 and 4. The burst
wave was transmitted ten times at the repetition
frequency of 2.5Hz in one measurement. TOFs were
measured by cross-correlation after averaging thereceived signal with the number of measurement times.
To synchronize the acoustic sensors distributed on the
network, trigger lines was set. Because the sampling
rate of the A-D/D-A was 250 kHz, the time resolution
was 4 μs. This line was wired line. The experiment
was carried out at 60s intervals from local time of
23:50 to 0:50 at the University of Tsukuba.

D

Acknowledgement
This work has been supported in part by aGrant-in-Aid for Scientific Research (B) (No.
17380149) from the Japan Society for the Promotionof Science.

Path F

Path Ath
Path B

Fig. 3 is the composition of the experiment. The
height of baseline was 2.15 m above the ground.
25

ultrasonicanemometer

Acoustic Sensor

Path A

24
23

Temperature( C)22Path B

25
24
23

25

Path C

24

22

23
22

Path D

25
24
23

25

Path E

24

22

23
22

Path F

25
24
23

4. Conclusions
In order to measure the temperature distribution of
the large space, we formed the wireless network andsix acoustic paths by four acoustic sensors in the 10×
10 ㎡ square in the open air, as the first experiment.
The monitoring system of the temperature wasenabled. Applying this system, wind measurements
would be realized by using methods of ultrasonicanemometers. In a future work, remote controlling
would be required. The remote controlling would be
realized via the Internet.

Pa

th

C

Path EPa

3.3 Measurement results and discussion
Figure 4 shows the measurement results of acoustic
path of A, B, C, D, E and F. The solid line represents
the time change in the mean spatial temperature. The
broken line represents the average of the temperaturepoint measured by ultrasonic anemometer (Model
81000 /Yong).
So the transmitted signal of adjoining sensors weremixed in the received signal, the TOF don’t turn out
separately. However, the TOF could measure bycross-correlation. This time, by using two kinds ofsignals and taking cross-correlation between thereceived signal and the signal transmitted from eachsensor, we could get a clear peak that shows the
TOF.As for the path A ,the time change of temperatureby acoustic sensors and ultrasonic anemometer has alonger periodicity with a maximum range of ±0.4℃
around 23.3 ℃.However, as for the path E, the timechange of temperature by acoustic sensors has ashort periodicity with a maximum range of ±0.6℃
around 23.3 ℃ .This difference is caused by thedifference of the spatial mean and the point
measurement.

reference

23:50

0:00

0:10

0:20

0:30

0:40

22
0:50

Time (10 min/div)
Fig.4 shows the measurement results of acoustic path
A, B, C, D, E, F.
1) J. E. Son, J. S. Park and H. Lee: Acta Hort. 578 (2002) 257.
2) D. J. Tanner and N. D. Amos: Acta Hort. 599 (2002) 193.
3) T. Honjo: J. Agric. Meteorol.56 (2000) 227.
4) X. Wang, T. Kawasumi, H. Tani: J. Agric. Meteorol.56(2000) 283.
5) G. Hofelmann, H. O. Luck and P. Beckord: J. Aerosol. Sci.
23 (1992) 51.
6) K. Mizutani, K. Nishizaki, K. Nagai and K. Harakawa: Jpn.
J. Appl. Phys.36 (1997) 3176.
7) E. Ishikawa, K.Mizutani and K. Nagai: Jpn. J. Appl. Phys.
40 (2001) 5446.
8) K. Attenborough et al.: J. Acoust. Soc. Am. 97 (1995) 1- 238 -

P2-14

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 239-240
16-18 November, 2005

Measurement of the Sound Speed in a Thread〇Shigemi Saito, Yasuhiro Shibata and Akira Ichiki (Tokai University)
1. Introduction
A sound speed and elastic constant are importantparameters to describe the mechanical property of amaterial. They are often used for identifying
materials. Although a certain volume is usuallyemployed for the measurement, some materials
such as natural fiber and human hair are found only
in a shape of thread. The data of sound speed
measured for various materials including biologicaltissues have been accumulated.[1,2] However thesound speed has scarcely been measured for suchmaterials. Different from a method to use laser forsound excitation that has previously been conducted
for ceramic fibers[3], a simple measuring method
for the sound speed in a more flexible thread suchas human hair is presented in this paper.
2. Transmission and Reception of UltrasoundThe experimental setup for measuring the soundspeed in a thread is shown in Fig.1. A glass beadwith a hole is firmly bonded on each surface of a
300 kHz cylindrical PZT transducer A with adiameter, thickness and weight of 10 mm, 6.5 mm
and 3 g so that the hole of each glass bead isparallel to the transducer surface. A thread sample
is fixed to one of beads with a thief knot, and astring to hang a plumb is connected to another bead.
A bead to fix the thread sample is bonded also to atransducer B of the same type. The transducer B isbacked with a cylinder of epoxy-resin and ironpowder mixture (ITW Industry, Devcon B), and thethread sample is hanged underneath. Both theplumb and the weight of the transducer A apply atensile force to the thread. One cycle 300 kHz
voltage with an amplitude from 5 to 60 V is applied
to the transducer A. An induced ultrasoundpropagates in the thread and then generates thevoltage across the transducer B. A distance between
two transducers, namely a thread length l, ismeasured with a reading microscope.
3. Measurement of Sound Speed
The waveform of the received voltage of thetransducer B for the case that a thread is a humanhair about 50 mm and 60 µm in length and diameteris shown in Fig.2. Although the waveform becomes
lengthy with slow transient responses due to thetransducer resonance, the propagation of ultrasoundis confirmed. The delay time τ from the rising
portion S of the driving voltage to the position R of
the first trough in the received voltage is measured
on a digital oscilloscope.

RS

τ
Fig.2. Example of observed waveforms.
80

Transducer B

70

Oscilloscope

Thread Sample
Reading
MicroscopelFunction
Generator

Transducer A

τ (µs)

60
50
40
30
20
10

Plumb

0
0

Fig.1. Experimental arrangement.

20

40

60

80

100

120

140

l (mm)

Fig.3. Measured delay for a 60 µm diameter hair.

ssaito@scc.u-tokai.ac.jp- 239 -

An example of measured delay time τ under a
15 gf load including the weight of the transducer A
is shown in Fig.3. The measured plots are almost on
a line. Extrapolation of this line yields a finite valueof τ even at l=0. This must be caused by a neglected
diameter of glass bead, a small local elasticity ofthe thief knot and the delay due to reading the firsttrough. From the reciprocal dl/dτ of the slope of theline, the sound speed is obtained as 2160 m/s.
Using various weights up to 100 g, the delay timeτ was measured for a 70 mm long copper wire 130µm in diameter. The result is shown in Fig.4 with ameasured l. Although the propagation speed oftransversal wave which we call string vibration is
proportional to the square root of tension, themeasured speed does not change much with thetension. Hence the detected wave is not atransversal wave but a longitudinal wave. In thisthread, the load of 100 gf corresponds to a tensilestress of 70 MPa. This is relatively large, but thethread is still within elastic limit.[4]nearly equal to polystyrene. A trend of high soundspeed in damaged hair is found for tested samples.
5. Conclusions
The longitudinal sound speed in a thread wasmeasured with a simple manner. The sound speed ina human hair showed individual differences.
References
1. P.-K. Choi: Ultrasonics Handbook (Maruzen,
Tokyo, 1992) p.714.
2. S. A. Goss, R. L. Johnston and F. Dunn: J.
Acoust. Soc. Am. 64 (1978) 423.
3. E. A. Lindgren, M. Rosen and K. E. Amin:
Ultrasonics 32 (1994) 411.
4. T. Iwatani et al.: Fundamental Material Strength.
(Maki Publishing, Tokyo, 1996) p.14.
60(a) Cu wire50

133 µm(3000 m/s)

40
τ (µs)

4. Measured Example

100 µm177 µm(3470 m/s)
(3190 m/s)

20
10
0
0

20

40

60

80

100

120

140

l (mm)

τ (µs)The l-τ relations measured for copper wires withvarious diameters, a gold wire with 30 µm diameterand a constantan wire with a diameter 100 µm
under a load of 25 gf, 3 gf and 25 gf, respectively,
are shown in Fig. 5. Obtained sound speeds are also
indicated in the figures. These values differ fromthe literature values, which may be calculated with
the Young’s modulus measured for block specimen.
The effect of processing to a thread is expected.
The l-τ relation was measured under a 4.5 gf
loading for human hair samples offered from overtwenty people. The examples are shown in Fig.6.
The sound speed disperses in a relatively wide
range between 2000 m/s and 2500 m/s due toindividual difference. Combining the specificgravity of a hair sample for 45 year old woman
measured as 1.22, the Young’s modulus is estimatedto be ρ0c02=4.98×109 Pa that means a hardness

30

90
80
70
60
50
40
30
20
10
0

(b) Au wire, Constantan wire
Au (2270 m/s)

Constantan (4100 m/s)
0

20

40

60

80 100 120 140 160 180
l (mm)

Fig.5. Measured result for metallic wires.
72

33

70

l

50

69

τ(µs)

32

45 year female(100µm, 2020m/s)

6071

l (mm)

τ (µs)

τ

70

40
30

22 year female(60µm, 2230m/s)

20
23 year male
(70µm, 2500m/s)

10
31

68
0

20

40

60

80

0

100

0

20

40

60

80

100

Load (gf)l (mm)

Fig.4. Dependence of τ on the tension in a copper wire.

Fig.6. Measured result for human hairs.

- 240 -

120

P2-15

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 241-242
16-18 November, 2005

Correlation Characteristic Improvement of Sound
Probe Using Inverse Problem◎Ikuo ODANAKA, Koichi MIZUTANI and Kyohei SAWAMURA
Graduate School of Systems and Information Engineering, University of Tsukuba

1. Introduction

Loud speaker(SP)Recently, the importance of a mean temperature
measurement of wide area is pointed out from theviewpoint of energy conservation1-3). However, tomeasure the spatial mean temperature, a lot ofelements are needed in the temperature measurement
element such as thermo-couples of the contact type.
They occupy working space and cause heat exchange.
For these reasons, acoustical temperature
measurement is suggested. Acoustical temperature
measurement uses a sound probe. Advantage of using
sound probe is that it has no environmental impact, beable to obtain the spatial mean temperature, withquick response, and without occupy working space4).
Time of flight (TOF) between two transducers (a loudspeaker – microphone) is obtained by using a crosscorrelation. An impulse response of acoustical system
is added to the signal and waveform distortion has
occurred. So measurement error in TOF has been
finally produced in a cross correlation result5). On theother hand, in the field of optical imaging andultrasonic imaging, a inverse problem is used andimaging accuracy improvement is put into practice by
the deconvolution of the impulse response fromreceived signal 6-7). In this paper, this method isapplied to the cross correlation characteristic
improvement of sound probe.

2. Time of Flight (TOF) Measurement
Figure 1 shows schematic diagram of a soundprobe. The sound probe is consisted by a couple of aloud speaker and a microphone. All of themeasurements in this paper are performed in an
anechoic room. A distance between transducers L is
fixed at 2.1 m. The TOF of the sound probe τ is
obtained based on a maximum value of the crosscorrelation value calculated by

φ = limT →∞1s(t )x(t + τ )dt ,2T ∫−TT(1)

Relationship between spatial mean temperature T and-----------------------------------------------------------E-mail: odanaka@aclab.esys.tsukuba.ac.jp

dis

tan

ce:
L

Transmitted signal:s(t)Microphone(MIC)

PC

Received singnal:x(t+τ)

Fig. 1 Schematic diagram of sound probe

the TOF τ is expressed as2

Τ0 ⎛ L ⎞T=⎜ ⎟ − Τ0 .
331.32 2 ⎝ τ ⎠(2)

T0 is 273.15 K. A spatial mean temperature T can bemeasured by using Eq. (2).

3. Correlation Characteristic Improvement
3.1 Waveform distortion
We should change a condition of a transmitted
signal such as a frequency by the change of conditionof measuring object. If the sound source of the
speaker were point sound source, a precise calculationof the TOF is performed at any frequency. In fact, inan acoustical measurement, an impulse response of
the acoustical system is added in the received signal
and a waveform distortion has occurred8). It causes
TOF measuring error.
3.2 Iimprovement
In this paper, sine wave with frequency of 3kHz isused as a transmitted signal. A relational expression of
I/O signals of an acoustical system is expressed as∞x(t ) = ∫ s (t )h(t − k )dk = s (t ) ∗ h(t ) .
−∞(3)

In this equation, s(t) is a transmitted signal, x(t) is areceived signal and h(t) is an impulse response of theacoustical system. Many researches for the method ofacquiring the impulse response have executed because- 241 -

S (ω ) =X (ω ).
H (ω )

1(a)0Relative Amplitude

the impulse signal could not be generated in a usual
method. Crossing spectrum method, maximum length
sequence (MLS) method, and time stretched pulse(TSP) method are typical method9-11). In this paper, theimpulse response is acquired by using the pulse wave
with limited amplitude and short continuance time.
This pulse wave has a flat frequency response at leastin the demanded range. Compute the Fouriertransform of both sides of Eq. (3) and transform it into
the shape as

TOF = 6.312 ms
-1
1

(b)

0
TOF = 6.296 ms
-1
1

(4)

(c)

So we can reduce the impulse response of theacoustical system by using Eq. (4).
Figure 2 (a) shows simulation of cross correlation.
This figure indicates that an ideal cross correlation
result has completely symmetric shape, Fig. 2 (b)shows the cross correlation result of transmitted
signal: s(t) and received signal x(t+τ) and Fig. 2 (c)shows the cross correlation result of transmitted
signal: s(t) and signal s(t+τ) acquired by thedeconvolution of the impulse response of an
acoustical system: h(t) from the received signal: x(t+τ).
Shape of Fig. 2 (b) is clearly not symmetric shape, andshape of Fig. 2 (c) is almost symmetric shape.

3.3 Application to TOF measurement
Using Eq. (1), TOF τ can obtained. An actual
spatial mean temperature T is 25.03 degrees Celsius(deg C), and the theory value of the TOF at 25.03 deg
C: τ is 6. 312 ms. The TOF derived from Fig. 2 (b): τ1
is 6.296 ms. To compare with theory value, measuring
error is 0.016 ms and calculated spatial mean
temperature: T1 is 26.54 deg C. The temperature
measuring error is 1.51 deg C. This result suggeststhat a small number of TOF measuring error causes a
lot of temperature measuring error. The TOF derived
from Fig. 2 (c): τ2 is 6.316 ms and calculated spatial
mean temperature: T2 is 24.65 deg C. When τ1 iscompared with τ2, the accuracy of τ2 is 0.012 ms
better than that of τ1 and accuracy of T2 is 1.13 deg C
better than that of T1. From these result, it can be said
that accuracy improvement of measuring TOF
achieves by using deconvolution.
4. Conclusions
Cross correlation characteristic improvement has
performed with inverse problem in this paper. TOF of0
TOF = 6.316 ms
-1
0

2

4
6
8
TOF(ms)Fig. 2 Results of the cross correlation. (a): The
simulation of cross correlation result (b): Thecross correlation result of transmitted signal: s(t)and received signal x(t+τ). (c): The crosscorrelation result of transmitted signal: s(t) andsignal s(t+τ) acquired by the deconvolution of theimpulse response from the received signal

sound probe has been calculated based on maximum
value of cross correlation result, so waveform
distortion causes measurement error. This error could
have been reduced by the elimination of the impulse
response of an acoustical system with deconvolution.
And a pulse wave with limited amplitude and shortcontinuance time can be used as an impulse signal.

Acknowledgment
This work has been supported in part by aGrant-in-Aid for Scientific Research (B) (No.
17380149) from the Japan Society for the Promotionof Science.
References(1) C. Boonen, K. Janessens, and D. Berckmans: trans. ASAE 45,
1077-1082 (2002).
(2) T. Boulard, S. Wang, R. Haxaire, C. Kittas, G. Papadakis, and M.
Mermier: Acta Hort. 507, 51-59 (1999).
(3) F. L. K. Kempkes, J. C. Bkker, and N. J. van de Braak: Acta Hort.
456, 363-370 (1998).
(4) S. Wang, M. Yernaux and J. Deltour: Journal of AgriculturalEngineering Research, Vol. 73, 189-197 (1999).
(5) E. Ishikawa and K. Mizutani: Japanese Journal of Applied Physics,
Vol. 42, No. 1, 8, 5372-5373 (2003).
(6) J. Biemond, R. L. Lagendijk and R. M. Mersereau: Proc. IEEE, 78,
5, 856-883 (1990).
(7) B. Haider, P. A. Lewin and K. E. Thomeuus: IEEE Trans. Ultrason.
Ferro. & Ferq. Control, 45, 1, 98-113 (1998).
(8) G. H. Glover: Neuroimage, vol. 9, 416-429 (1999).
(9) F. Satou: Proc. Of Forum Acusticum Sevillia 2002 (2002).
(10) N. Aoshima: J. Acoust. Soc. Am., 69, 5, 1485-1488 (1981).
(11) Sinha: D. Reidel Publishing Company (1986).

- 242 -

P2-16

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 243-244
16-18 November, 2005

Suppression of spurious vibration of cantilever in ultrasonic atomic force microscopy◎Toshihiro Tsuji, Seishiro Ide, Kazushi Yamanaka (Tohoku University)

k n* k1 = rne + rnv (Ω Q1 )i

1. Introduction
Ultrasonic atomic force microscopy (UAFM) is
now being accepted as a practical tool for
non-destructive evaluation of mechanical properties
and subsurface defects in MEMS devices [1-3].
Since in UAFM and non-contact AFM with atomic
resolution [4] resonance vibration of cantilever ismeasured, spurious vibration (SV) could degrade
precision of the measurement. Although it has beenrecognized as a serious problem in UAFM, origin of
the SV has not been clarified so far. In this paper, we
propose an effective method for suppression of SV.

Ω =ω

(k r (Ω Q )) ,k1 m1 , Q1 = m1 k1

v
1 1

1

where r , r are relative stiffness and viscosity, Ω
is frequency ω normalized by the resonancefrequency of the cantilever, and Q1 is the quality
e
n

v
n

factor related to the cantilever, respectively. r0e iscalled as relative contact stiffness. Position x1 andx 2 are expressed by⎛ x1 ⎞⎛ A B ⎞⎛ D ⎞1⎜⎜ ⎟⎟ =⎜⎟⎟⎜⎜ ⎟⎟
2 ⎜
x
−
AC
B
⎝ B C ⎠⎝ E ⎠
⎝ 2⎠

(

(2)

)

A = 1 + r0e − Ω 2 + 1 + r0v (Ω Q1 )i , B = 1 + (Ω Q1 )i ,2. Spurious vibration
Figure 1 shows amplitude spectra of deflection
vibration at the top of the cantilever. Excitation
power of piezoelectric vibrator was -35dBm.
Fig.1(a) is a non-contact spectrum. Although NCD1, 2
and 3 are the 1st, 2nd and 3rd deflection modes,
respectively, identified by cantilever vibration theory,
there are many SVs. Fig.1(b) is a spectrum where
the tip is in contact with (100) surface of Si wafer at
the load of 500nN. Although CD1 and 2 are the 1st
and 2nd modes, they were interfered by SV1 and 2.

( )D = (r + r (Ω Q )i )x , E = (r + r (Ω Q )i )x ,C = 1 + r − MΩ + 1 + r2v (Ω Q1 )i , M = m 2 m1 ,
2

e
2

e
0

v
0

1

SV1
NCD1

0

e
2

1

3

x0 = 0x 3 = aecorresponds to the situation of the UAFM.
On the other hand, the boundary condition ofx 0 = ae iωt and x 3 = 0 corresponds to the situation of
atomic force acoustic microscopy [5].
(a) Clamping wire

Piezoelectric
vibrator

(b)

x3
x2

x1
CD1

SV1

SV2

CD2

Base

(b)
(–50dB shifted)
0

k 2*
m2

SV2

(a)
–100

and

iω t

NCD3

NCD2

v
2

where the boundary condition of0
Amplitude (dB)

(1),

200
400
Frequency (kHz)L

Cantilever

x0

k1*
m1

k 0*

Fig.2 Model of spurious vibration. (a)Schematics of
cantilever holder. (b)Mass-spring model.

600

Fig.1 Resonance vibration affected by spurious
vibration. (a)Non-contact (b)Contact on Si wafer.

Figure 3 shows amplitude spectra calculated by
Eq.2 with the boundary condition of UAFM, whereparameters are r0e = 10 , Q1 = 100 , M = 1 × 104 ,r0v = 1 are common. Fig.3(a) shows the dependence

3. Suppression of spurious vibration
First, we propose and verify that SV is originatedfrom bending vibration of the locally clamped base as
shown in Fig.2(a). We model this situation as shownby Fig.2(b), where m n , k n* , x n represent mass,
stiffness and position, respectively, and subscriptn = 0,1,2,3 represent sample stage, cantilever, locallyclamped base, and cantilever holder, respectively.
Stiffness normalized by k1 is expressed byofspectra

on

r2e ,

where

r2v = 1 .

Whenr = 1 × 10 (solid curve), SVI interferes the resonancee2

5

due to r0e , represented by CR. When r2e is 1/8
times lower (chained curve) or 8 times higher (dashedcurve), the interference can be avoided (SVII, SVIII).
Because r2e is approximately proportional to thethird power of the thickness of the base, the increase- 243 -

of r2e can be realized easily. Moreover, thefabrication of the holder that clamps whole part of thebase is effective because r2e is enhanced.
Figure 3(b) shows the dependence of spectra
around CR on r2v , where r2e = 1× 105 . Solid curve

10

4

10

2

10

0

NCD2
(a) SVII SVI

SVIII

r2e = 8×105
r = 1× 10–2

r = (1 8) × 10
e
2

–4

10

0

10

4

10

2

10

0

5

–50

10

(a)
CD1

0

SVI , r = 1
v
2

(b)

r = 1× 10
v
2

3.2
3.4
Normalized frequency Ω

3.6

Fig.3 Variation of spurious vibration in resonancespectra. Dependence on (a) relative stiffness and (b)relative viscosity of the base..
4. Experimental verification
When the length of base L is changed (Fig.2(a)),
we can decrease r2e . Dashed lines in Fig.4 represent
contact resonance frequency (CD1,2). As L
increased, SV was shifted to lower frequency (SV1,2).
0

L=
1.98
L=
2.24
–100
L=
2.61
L=
2.97
(mm)
0

200
400
Frequency (kHz)600

5. Conclusions

r2v = 1× 10 6

SV1

CD2

Fig.5. Suppression of spurious vibration shown in
Fig.1 by modified holder. (a)Non-contact and (b)Contact on Si wafer.

CR

5

NCD3

(b)
–150 (–50dB shifted)

5

Normalized frequency Ω3Amplitude (dB) (shifted by –30 dB.)

NCD1

–100

e
2

10

Amplitude |x1/x3|

CR

Amplitude (dB)

Amplitude |x1/x3|

shows the same case as Fig.3(a). When r2v is in therange of 10e5 to, 10e6, the SV is almost damped(dashed and chained curves, respectively). In order torealize large value of r2v , not only the selection of thebase material but also the effective design of thecantilever holder is important.

The behavior of SV1 can be explained by analyticalsolution of deflection vibration of thick plate (Dashedcircles) [6]. However, other SVs may affect CD1,2.
Aiming to remove all affection, we fabricatedprototypical holder that clamps whole part of the baseby a stainless steel cover (modified holder). Fig. 5
shows a result. Most of SVs shown in Fig.1 were
suppressed, which indicates that the cover is effective
to the suppression of not only deflection modes but
also other vibration modes, such as torsional andlateral modes.

SV2

CD1
200
400
Frequency (kHz)

CD2
600

Fig.4. Dependence of spectra on base length. Contact
on Si wafer.

We showed the effective method for the suppressionof spurious vibration in UAFM, based on the analysis
of coupled vibration system. The holder that clamps
whole part of the base and the increase of the basethickness are promising method. We verified that themodified holder effectively suppressed the spurious
vibration.
Acknowledgement
This work was supported by the Ministry of
Education, Culture, Sports, Science and Technology.
References
1. K. Yamanaka and S. Nakano: Jpn. J. Appl. Phys.,
Part 1, 35 (1996) 3787.
2. T. Tsuji, H. Ogiso, J. Akedo, S. Saito, K. Fukuda
and K. Yamanaka: Jpn. J. Appl. Phys., Part 1, 43(2004) 2907.
3. T. Tsuji, S. Saito, K. Fukuda, K. Yamanaka, H.
Ogiso, J. Akedo, and Y. Kawakami: Appl. Phys. Lett.
87 (2005) 071909.
4. Y. Sugawara, M. Ohota, H. Ueyama and S. Morita:
Science 270 (1995) 1646.
5. U. Rabe, K. Janser and W. Arnold: Rev. Sci.
Instrum. 67 (1996) 3281.
6. S. P. Timoshenko, D. H. Young and W. Weaver, Jr.:
Vibration Problems in Engineering (Willey, New York1974) p. 492.

- 244 -

P2-17

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 245-246
16-18 November, 2005

Effect of temperature on the response of ball SAW hydrogen gas sensor
Kazushi Yamanaka1, ○ Takuji Abe1, Naoya Iwata1, Toshihiro Tsuji1, Tsuyoshi Mihara1,
Kanwar Jit Singh1, Shingo Akao2, Noritaka Nakaso2, Tsuneo Ohgi2, Dong Youn Sim3 , Yusuke
Ebi 4, Takeshi Fukiura4, Hidekazu Tanaka4
1. Tohoku University, 2. Toppan Printing, 3. Ball Semiconductor, 4. Yamatake
1. Introduction

constructed for an acoustic microscopy [6].

Hydrogen fuel cells are expected as themost promising clean energy source.
However, since hydrogen is an explosive gas,
security of fuel cells requires hydrogensensor with wide sensing range, fast
response. Hence, we have developed a single
element hydrogen sensor using surface
acoustic wave (SAW) on a sphere withmultiple roundtrips of collimate beam [1,2] It
realized wider sensing range than other SAW
sensors [3,4] or FET sensor[5]. In this paper,
we report effect of temperature on the
response of the ball SAW hydrogen sensor.

3. Experimental results
Figure. 2 shows an envelop signal of 34 MHz
tone burst at (a) 1 to 73 turns, (b) 60 to 68
turns and (c) 65 turns. The vertical axis is 1.6
log10 (V), where V is the received signal.
From this figure, the Q factor was evaluatedas ~37,000. In (c), five curves were obtained,
just before and after introduction of 1.8 % H2
at 23 ℃ . It is noted that the amplitudedecrease was completed within the samplinginterval (2s).

2. Equipment for sensor evaluation

2

Figure 1 shows a setup. We deposited a
40nm-thick Pd-Ni alloy thin film on a
Langasite ball of 10 mm diameter. The ballwas exposed to hydrogen at concentrations
from 0.1 % to 3% by using a flow cell usingmass flow controllers (MFCs). Since the MFC
does not tolerate high temperature (>100 ℃), the gas was rapidly heated after
passing through the MFC. The heater consistof a densely wire together with atemperature regulator.

10

Log V a.u.

0

1000

0.4
0.2
0
800

900

TC
Regulator

0.3
HeaterN2 lineH2 linePulser

MFC
TC
MFC
10 L/minDigitalVoltmeter

Computer

0.25

Tone burst
transmitter/receiver862

Oscilloscope

864

Time (μs)

Fig.1. Setup of Ball SAW sensor evaluation
The SAW was excited and detected by an
IDT. The measurement was performed either
by a short pulse excitation (dotted box) [1,2]or by a tone burst transmitter and receiverwith a quadrature and envelop demodulator,

Fig. 2. Amplitude change of the tone burst envelop
signal at room temperature (23 ℃).
Figure 3 shows an amplitude response
obtained with the digital signal analysis ofpulse waveform [1,2] for H2 concentrations
between 0.1 % and 3 %. The magnitude of the- 245 -

270

250
–40

320

310

–40

40

0

Time (s)40(d) 124℃ 1.8%

1.2％
1.8%

0.028

270

260

–40

0.026
100

320

310

(b) 23℃ 1.8%

2.4%

0

0

Time (s)Amplitude (mV)

0.03(c) 124℃ 0.6%

260

0.1％
0.1％
0.6％
Amplitude (mV)

Amplitude (V)0.032(a) 23℃ 0.6%
Amplitude (mV)

0.1% 0.1% 0.6% 1.2% 1.8% 2.4% 3.0%
H2
H2
H2
H2
H2
H2
H2
N2
N2
N2
N2
N2
N2
N2
N2

sensor protecting membrane.

Amplitude (mV)

response (decrease of amplitude) increased
with the H2 concentration. Since the response
was not saturated at 3% H2, there is apossibility to differentiate H2 concentration
above 3 %, in contrast to our previous
experiment [2].

0Time (s)

40

–40

0

Time (s)3.0%Time ( min )Fig. 3. Amplitude response of the sensor in H2 from
0.1 to 3 %.
Figure 4 shows the effect of temperatureon the response for low (0.6%) and medium(1.8%) concentration H2, introduced at timet=0. The response time is defined as a timerequired for 90 % change of the amplitude. At
23 ℃, the response time was around 4 s at
0.6 % H2 (a), whereas it was shorter (~2s) at
1.8 % H2 (b). However, when the sensor washeated (124 ℃ ), the response time wasshorter than at room temperature. It was 2 s
at 0.6 % H2 (c), whereas it was less than 2s at
1.8 % H2 (d). This result indicates thatheating to 124 ℃ reduces the response time.
We also note an overshoot of the signal atroom temperature where the signal firstdecreased rapidly and then increased slowly
to the equilibrium. Such an overshoot was
not present at 124 ℃, which is favorable forreliable sensor operation. Since we observed
these distinct changes within a samplinginterval, the flow cell response within 2s was
confirmed.
As a drawback of the sensor at its present
form, a base line drift was observed in Figs. 3
and 4. The possible reason may be either
variation of temperature, pressure or gas
atmosphere such as humidity. The drift may
be eliminated by using a reference sensor
without sensitive film [3,4] and by using a

Fig. 4 Effect of temperature on the responsetime at room temperature (23 ℃) ((a) and(b)) and at 124 ℃ ((c) and (d)).
4. Conclusions
We investigated the effect of temperatureon the response of the ball SAW H2 sensor
and found that the response time is reducedby heating to 124 ℃. The overshoot of thesignal observed at room temperature wassuppressed by heating.
Acknowledgement
This work was supported by the Ministryof Education, Culture, Sports, Science andTechnology.
Reference[1] K. Yamanaka, S. Ishikawa, N. Nakaso, N.
Takeda, T. Mihara and Y. Tsukahara, Proc.
2003 IEEE Ultrason. Symp. 299 (2004).
[2] A. Mizukami, D. Y. Sim, I. Sato, T.
Miyagishi, H. Tanaka, T. Fukiura, H. Kazato,
S. Akao, T. Ohgi, N. Nakaso, H. Watanabe, T.
Mihara and K. Yamanaka, Proc. Symp.
Ultrason. Electron., 25, 301 (2004).
[3] A. D’Amico, A. Palma, E. Verona, 1982 IEEE
Ultrasonics Symp. 308 (1982).
[4] W. P. Jakubik, M. W. Urbanczyk, S.
Kochowski and J. Bodzenta, Sens. Actuators B
82, 265 (2002).
[5] R. C. Hughes and W. K. Schubert, J. Appl.
Phys. 71, 542 (1992).
[6] K. Yamanaka, Y. Nagata and T. Koda
Ultrasonics, 29, 159 (1991).

- 246 -

40

P2-18Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 247-248
16-18 November, 2005

Electromagnetic Acoustic Resonance to Assess
Creep Damage in a Cr-Mo-V Steel◎Toshihiro Ohtani (Ebara Research Co., LTD.), Hirotsugu Ogi and Masahiko Hirao(Osaka University)
1. Introduction
1,2)Electromagnetic acoustic resonance (EMAR)
is an emerging technigue for measuring theultrasonic velocity and attenuation of the metallic
materials. It incorporated an electromagnetic
acoustic transducer (EMAT) 1,2) in the resonant
technique, realizing a non-contact and highly
accurate measurements. In this study, the EMAR
was applied to detect the creep damage process in a
Cr-Mo-V steel, which is an important structual
material for the thermal energy plants. The material
was exposed to the temperature of 923K at various
stresses. Two types of EMATs were used: thebulk-wave EMAT for plate samples and theaxial-shear-wave EMAT for cylindrical samples.
We measured ultrasonic attenuation in thefrequency range between 1 and 7 MHz as the creep
advanced. The attenuation coefficient exhibits muchlarger sensitivity to the damage accumulation thanthe velocity. It shows a peak at around 30% and aminimum value at 50 % of the creep life, being
independent of the applied stress and type of EMAT.
The EMAR has a potential to assess the damage
advance and to predict the creep life of metals.
2. Samples and EMATs
We obtained the creep samples from two
hot-rolled commercial round-bars of a Cr-Mo-V
steel (JIS-SNB16). They were heated at 1283 K for
2 h, air-cooled, heated at 1223 K for 2 h, oilquenched, and heated at 963 K for 6 h, air-cooled.
One was used to make the plate specimen; its gauge
section was 5 mm thick, 18 mm wide and 35 mm
long. The other was cylindrical specimen; its gauge
section was 14 mm in diameter and 60 mm long.
Their mechanical properties are as follows; 0.2%
proof stresses is 778 MPa, ultimate tensile strength
834 MPa and elongation 22 %.
We have built two types of EMATs which generate
and receive the shear waves with themagnetostrictive effect of ferromagnetic materials

by a non-contacting manner. One is the bulk-wave
EMAT1,2) for the plate samples, which generates thepolarized shear waves traveling along the direction
normal to the surfaces. This EMAT is used to excite
and detect the through-thickness resonance for plategeometries. The resonant peaks appear at an equal
frequency interval and at each resonant frequency;
the attenuation coefficient is measured by thefree-decay method, leading to the frequency
dependence of attenuation. The other EMAT is forthe cylindrical samples and generate the axial
shear-wave; a surface shear horizontal wave, which
propagates in the circumferential direction with anaxial polarization 2,3). The resonant mode of theaxial shear-wave has own propagation region in the
cross-section. The simultaneous measurement of theresonant frequency and attenuation of the several
modes then enable us to evaluate the material
inhomogeneity in radial direction 2).
3.

Experiment

Creep tests were carried out at 923 K in air forvarious stresses between 25 and 55 MPa. Alongside
the samples, the unstressed samples were placed toinvestigate only the effect of thermal history. We
measured the resonant frequencies and theattenuation coefficients using the EMAR techniquewith the analog superheterodyne spectroscopy. The
details of measurements and electronics can befound elsewhere 1, 4).
In creep test, we interrupted loading andfurnace-cooled the samples in a furnace. Aftermeasuring the frequencies and attenuation
coefficients at room temperature, we restarted thecreep test. We repeated this procedure for every 20,
30 or 100 h until the rupture.

--------------------------------------------------------------------E-mail: ohtani.toshihiro@er.ebara.com- 247 -

Figure 1 shows the creep life fraction, t/tr (thecreeping time/ the rupture life), versus theattenuation coefficient, α of the 11th resonantmodes (around 3.5 MHz) under four different
applied stresses. It was the results for the platesamples. Rupture lives were 1018.1 h, 759.1 h ,292.0 h and 258.5 h for 25, 35, 45 and55MPa,
respectively. α initially increases as creep
progresses and shows a peak around t/tr=0.3. After
that, it decreases to t/tr=0.5 and increases to therupture. This trend is independent of the appliedstress. No remarkable change occurred in the
unstressed samples.

in cell and subgrains. They vibrate to absorb theultrasonic-wave energy. However, they areabsorbed into the cell walls and sub-boundaries ascreep progresses to cause the minimum attenuation.

0.03
923K in air
First resonant mode

-1

Results

Attenuation coefficient [μs ]4.

:35MPa:45MPa:55MPa

0.02

0.01

0.00

0.0

0.5

1.0

Life fraction, t/tr
-1

Attenuation coefficient [μs ]0.03

Fig.2. Relationship between the attenuation coefficient of
the 1st resonant modes and life fraction with an axial
shear-wave EMAT (35, 45 and 55 MPa, 923K).

:25MPa:35MPa:45MPa:55MPa923K in air
11th resonant mode
0.02

0.01

0.00

0.0

0.5

1.0

Life fraction, t/tr

Fig.1. Relationship between the attenuation coefficient of
the 11th resonant modes and life fraction with a bulk
shear-wave EMAT (25, 35, 45 and 55 MPa, 923K).

Figure 2 shows the relationship between α of thefirst resonant mode (around 3.8 MHz) and t/tr under
three different stresses for the rod samples. Rupture
lives were 774.3 h, 401.5 h and 185.0 h for 35, 45
and 55MPa, respectively. The trend of α is very
similar with that observed in plate samples by thebulk-wave EMAT. Namely, α experiences the peakat 30% and the minimum at 50 % of the creep life,
being independent of the applied stress and type of
EMAT.
Furthermore, we observed the change of
microstructure; grain size, precipitation density anddislocation structures with optical microscope,
SEM and TEM. The evolution of α as creep
progress is related to the microstructure change,
especially, dislocation structure. At attenuation peak,many long and isolated dislocations are generated

4. Conclusion
Creep damage in Cr-Mo-V steel (JIS-SNB16) at
923 K in air was evaluated through ultrasonicattenuation measured with the EMAR method.
Attenuation showed a peak at around 30% and aminimum value at 50 % of the creep life, being
independent of the applied stress a type of EMAT,
which is interpreted as resulting frommicrostructural changes, especially, dislocation
mobility. This is supported by TEM observations.
This technique has a potential to assess the damage
advance and to predict the creep life of metals.
References
1. M. Hirao, H. Ogi and Fukuoka: Rev. Sci. Instrum.
64 (1993) 3198.
2. M. Hirao and H. Ogi: EMATs for Science andIndustry, (Kluwer Academic Publishers, Boston,2003), 2.
3. H. Ogi, M. Hirao and K. Minoura: J. Appl. Phys.
81 (1997) 3677.
4. G.L. Petersen, B. B. Chick, C.M. Fortunko and M.
Hirao: Rev. Sci. Instrum. 64 (1993) 3198.

- 248 -

P2-19Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 249-250
16-18 November, 2005

Surface Roughness Characterization by Angular Distributionof Scattered Waves Using Air-Coupled Ultrasonic Technique◎Deden Dian Sukmana and Ikuo IHARA (Nagaoka Univ. of Technology, Mechanical Eng.)

1. Introduction
Surface roughness is an important parameter forevaluating surface properties of materials and istypically measured by a stylus instrument that
amplifies and records the vertical motions of astylus as it moves across the surface. For soft anddelicate surfaces, however, non-contact techniquessuch as optical methods are preferable1). In thiswork, a novel non-contact technique using
air-coupled ultrasound has been used to characterizesurface roughness of the materials. The technique isbased on scattering phenomenon. In our previous
work 2, 3), the scattered waves from material surfaces
were measured at various reflection angles θ2 on thexz-plane shown in Fig. 1, and the correlation
between the scattered waves and the surface
roughness was obtained. In this work, the scatteredwaves at a certain reflected angle are measured as afunction of azimuthal angle θ3 in xy-plane. The
relationship between the intensity distribution of thescattered wave and the surface roughness is
examined for various θ3. The Kirchhoff-based
scattering model is used to verify the validity of theexperimental results.

surface profiles of each specimen obtained using astylus profilometer. The height and lateralroughness parameters used for evaluating surface
profile are the root-mean-square roughness Rq andthe surface correlation length λo, respectively4). The
Rq and λo of the specimens are deviated from 0.04
to 115 µm and from 83 to 250 µm, respectively. The
ultrasonic wave is incident onto the specimen
surfaces at incident angle θ1=30o and the scatteredwave is measured at scattering angle θ2=30o andθ3=-90o~90o.
z

Incident wave

Scattered wave

θ1

θ2

y

θ3

x

Fig. 1 The scattering geometry used in theexperiment.

Rq (µm)-115.4

2. Experimental- 84.2The measurement system is mainly consisted of apair of broadband air coupled capacitancetransducers (Micro Acoustic Co., BAT) having
center frequency of 0.5 MHz, a square wave pulser(Ritec, SP-801A) and a broadband receiver (Ritec,BR-640A) 2, 3). The working distance between thetransducer and specimen surface is 35 mm. Figure 1
shows the scattering geometry used in theexperiment. Seven kinds of sandpaper sheets having
different roughness were used for the specimens
and were bonded on a steel plate to make flatness.
A polished stainless plate was used as a reference
specimen of smooth surface. Figure 2 shows the- 52.3- 42.6- 25.5- 18.7- 9.8- 0.04
0

10

20

30

40

Length (mm)Fig. 2 Surface profiles of each specimen measured
by a stylus method.

Email: ihara@mech.nagaokaut.ac.jp- 249 -

3. Results and discussion

Amplitude (arb. units)

0.010

Amplitude (arb. units)

0.35
Rq (µm)
0.04
9.8
18.7
25.5
42.6
52.3
84.2
115.4

0.30
0.25
0.20
0.15
0.10
0.05

Experiment
Theory0.008

θ3=70oRq=
84.2µm

0.006
0.004
0.002
0.000

Rq=
42.6µm

50

0

20

40

60

Rq=
52.3µm

Rq=
9.8µm
Rq=
Rq=
25.5µm
18.7µm0.00-80 -60 -40 -20

Rq=
115.4µm

80

100

150

200

250

Correlation length (µm)

Scattering angle, θ3 (degree)
Fig. 3 The intensity distribution of the scatteredwaves from the surfaces having different
roughness.
Figure 3 shows the change of the amplitudes ofthe scattered waves with the scattering angle in theazimuthal direction θ3 for various roughnesses. At
the specular reflection angle in which θ1=θ2=30o
and θ3=0o, the coherent component is dominant in
the scattered waves. We can see that the amplitudedecreases as roughness increases. The relationshipbetween such the coherent component and thesurface roughness was discussed in details in ref. 5.
In this work, we focus on the azimuthal angle
dependence of incoherent component. For θ3>30o,the incoherent component is considered to be
dominant in the measured waves. For the scatteringgeometry shown in Fig.1, the incoherent component
can be estimated using the Kirchhoff-based
scattering model and written as6)

(

)

2
2
2
2
2 2 2 -gk A + B λo
∞ gm
πλ F k eo∑=exp(),
I
A
incoh
M
2
m!m
4m
( 4πR )
m =1…………(1)
where g=k Rq (cos θ1+cos θ2) , k is the wave
number, AM is insonified area, R is the distance
between the specimen surface and the transducer, F,
A and B are the angular factors defined by θ1, θ2and θ3, respectively. It is noted that the incoherent
component depends on not only the roughness Rq
but also the surface correlation length λo of thesurface profile. Therefore, the correlation length λocan be estimated from the amplitude of incoherent
component when the roughness Rq is known.
2

2

Fig. 4 Relationship between the amplitude ofincoherent component and surface correlation
length λo.
Figure 4 shows the relationship between theamplitude of the incoherent component and thecorrelation length λo obtained experimentally andtheoretically. From this result, however, we havefound that there is no significant difference from theprevious results3) obtained with θ3=0o. This isbecause the surface characteristic of the sandpaper
is an isotropic and homogenized. The result may
differ if the specimen surface being used has adirectional roughness, as usually found in thesurface produced by machining process in which
the surface profile strongly depends on theazimuthal angle. In this case, the measurement of
the scattered wave in the different azimuthal angle
may give any useful information about thedirectional roughness.
Acknowledgment
The financial supports from Grant-In-Aid for
Scientific Researches (B17360351) by JSPS are
appreciated.

2

References
1.
2.
3.
4.
5.
6.

- 250 -

DJ. White House: Meas. Sci. Technol. 8 (1997)
pp.955-972.
Deden DS, et al.: Proc. Symp. Ultrason. Electron. 24(2003) pp.101-102.
Deden DS, et al.: Proc. Symp. Ultrason. Electron. 25(2004) pp.239-240.
Y.P Zhao: A Simple Introduction to Rough Surface
Characterization, 2000.
Deden DS, et al.: Jpn. J. Appl. Phys. 44 (2005) pp.
4417-4420.
J.A. Ogilvy: Theory of Scattering from Random RoughSurfaces (Bristol, IOP, 1991).

P2-20

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 251-252
16-18 November, 2005

Electromagnetic acoustic transducer (EMAT)
for generation and detection of guided wave
SongSong Li1, Toshimi Okada1 and Xiaoming Chen2
1.Toyama Prefectural University, Electromagnetic Wave Laboratory, 2.Toyama University

1. Introduction
A variety of guided waves techniques havebeen developed and their applications in NDE
have been discussed in the published articles[1]. The research effort into guided waves inplates has developed from the classicalRayleigh-Lamb wave [2]. For a given thickness
and frequency, there may exist many different
modes. This tends to make the received signals
complex and so difficult to interpret. For NDE,
it is desirable to excite a pure single mode ofpropagation because this will make theanalysis of data become easier. In this paper, aLorentz Force EMAT system is proposed for theexcitation and detection of single Lamb wave
mode and Rayleigh wave by controlling thecenter frequency of narrowband signals.
2. Dispersion curves of Lamb wave and EMAT
designFigure 1. Lamb wave dispersion curves for steel
plate: (a) phase velocity and (b) group velocity
Figure 1 (a) and (b) show the phase velocity
and group velocity dispersion curves of Lambwave in a steel plate, and the maximum limit of
frequency-thickness
productrange
is4MHz-mm. The wavelength of Lamb wave, λ ,1 isgiven by λ =cpf, where c p is the phase

1 Email: s-li@rdw.pu-toyama.ac.jp

velocity of the mode and f is the frequency. By
the dispersion curves, the phase velocity is thefunction of f * d . Eq. (1) is also written as

λ
d

=

cp
fd

, where d is the plate thickness. For agiven plate thickness and wavelength, theconstant λ / d is the tilt of the straight line
though the origin on the coordinate plane of thephase velocity dispersion curves, and the points
at which the line with tilt λ / d intersect thedispersion curves of the different modes are thepoints at which the modes can be excited.

Figure 2. Configuration of Lorentz ForceEMAT for generation of guided wave
The wavelength of Lamb wave is controlled
by the configuration of EMAT. A Lorentz EMAT
consists of a meander coil and a static
permanent magnet, which is shown in figure 2,
where the bias field is perpendicular to theplate and the magnet is over the meander coilthat is exerted an alternating current. The
meander coil creates eddy current in the plate,which then create the force in the plate by the
Lorentz interaction with the biased magnetic
field and hence give rise to the Lamb wave andRayleigh wave. The process can also be used todetect Lamb wave and Rayleigh wave. In themeander coil EMAT construction, thegeneration of the Lamb wave and Rayleighwave is realized by the spatially periodic stressradiating energy at an oblique angle. When the
period (2D) of the meander coil line acts as anultrasonic wavelength λ , the efficiency should
be greatest, thus, the optimized Lamb wave
mode can be excited. Therefore, in our
experiments, the spacing size (D) of meander
coil is selected as half a wavelength.
In Figure 2, the size of the magnet is 25mm x
25mm x 10mm, thus, the meander coil is- 251 -

designed as 6 fingers, and the spacing D
between connected fingers is 3.5mm (half awavelength). The wavelength of the Lamb wave
is 7mm.For a given plate thickness, theexcitation frequency of each mode is found. For
example, for 1.5mm thick steel plate, the valueof λ / d is about 4.7. The line with tilt of 4.7
intersects the dispersion curves of A0 and S0
mode at points A and B, indicating that the A0
mode and S0 mode can be excited at points A
and B, respectively, as shown in figure 1 (a).
Thus, the excitation frequency of each modecan be calculated. On the dispersion curves ofphase velocity, it is known that thefrequency-thickness product for the A0 mode is
about
0.36MHz-mm,
andthefrequency-thickness product for S0 mode isabout 1.02MHz-mm. Therefore, the excitation
frequency for A0 is 0.24MHz, and the frequency
for S0 is 0.68MHz. Similarly, for 1mm and 2mm
thick steel plate, the value of λ / d is 7 and 3.5,
respectively, and the excitation frequencies ofeach mode can be fond.
On the other hand, by the dispersion curvesof group velocity, the theoretical group velocityof the corresponding mode can be known. For
example, for 1.5mm thick steel plate, the groupvelocities of A0 mode and S0 mode are 2520m/s
and 5200m/s, respectively, as shown in the
figure 1 (b). These velocities can be used tocompare with the real velocity of the single
mode, and confirm if the desired single mode
has been excited.
The operation process of the experimental
setup used to excite and receive the Lamb wave
is: the tone-burst pulse generated by the pulse
generator, after being amplified by the power
amplifier, is sent to the T-EMAT (Transmitter-EMAT). The excited Lamb wave propagatesalong the steel plate. The transmission wave
and the reflection wave are received by the
R-EMAT (Receiver -EMAT). After being
amplified, the received waveforms are recordedin the digital oscilloscope, and are finally stored
in computer for the data analysis.
4. Experimental results
A series of tests were performed in the
different thickness steel plate. The transmitter
EMAT is placed on the left side of plate to avoid
the bidirectional signal, and the receiver EMAT
is placed on the position that is 50cm awayfrom the left side of plate. In order to observe
easily the position where Lamb modes appear,
the distance between the T-EMAT and R-EMAT
is 50cm.
It is known that Lamb wave modes are
successfully excited in different plate

specimens, and with increasing of plate
thickness, the excited Lamb modes also is
increased. For example, in 1mm and 1.5mm
steel plates, only A0 mode and S0 mode are
excited, while in 2mm steel thick plate, A1
mode also is excited. At the same time, it is foundthat this EMAT configuration also excited theRayleigh wave.
Figure 3 is a typical result using EMAT togenerate each wave mode in the 1.5mm steel
plate: (a) shows the Lamb wave A0 mode that it
is excited at 0.24MHz, (b) shows a Rayleighwave that is excited at 0.46MHz, and (c) shows
the Lamb wave mode S0 that is excited at
0.64MHz. A0(1), R(1) and S0(1) indicate thedirect wave signal, and R(2) and S0(2) indicate
the reflected wave signal from the right end ofplate. S(3) indicate the reflected wave from theleft edge of plate, because the attenuation of
the S0 mode is slow so that the reflected wave
signal from the right end of plate is again
reflected from the left edge of plate. From theexcitation result of each mode, it is found thatwhen one mode is excited, other modes are notexcited at this frequency.

Figure 3. Waveforms of the modes excited in
1.5mm steel plate. (a) A0 mode, (b) Rayleighwave and (c) S0 mode
5. Conclusion
In this paper, a narrowband signal EMAT
system that can be used to generate and detect
the pure modes is described. It has been shownthat this EMAT configuration can be usedsuccessfully to excite and receive the Lambwaves and Rayleigh wave, and the results are
consistent with the theoretical predictions ofdispersion curve.
References[1] Joseph L. Rose: Journal of Pressure Vessel
Technology. 124 (2002) 273[2] Viktorov I A: Rayleigh and Lambwaves-Physical Theory and Applications.
(New York: Plenum) 1967

- 252 -

P2-21

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 253-254
16-18 November, 2005

Synthesis of the rejection band profile
in ultrasonically induced optical fiber long-period grating◎Norimichi Fukuma, Kentaro Nakamura, Sadayuki Ueha (P&I Lab., Tokyo Inst. of Tech.)
1. Introduction
Among various kinds of acousto-optic devices,
fiber optic configuration has an advantage in
insertion loss and coupling optics. An opticalfiber long-period grating (LPG) is obtained byultrasonic bending vibration propagating along thefiber [1][2]. The ultrasonic vibration generates a
temporary periodic modulation in the refractive
index of fiber, which causes mode couplings
between the fundamental transmission mode of thefiber and the decaying modes at a certain optical
wavelength determined by the spatial period of themodulation. The ultrasonically induced grating
has an advantage in tunability. The rejection
optical wavelength is tunable with the ultrasonicfrequency. Authors experimentally investigatedthe vibrations excited on the fiber, as well as the
optical transmission characteristics of theultrasonically induced LPG [3]. A method tosynthesize the rejection band shape is proposed inthis report.

well agree with the Timoshenko’s beam theory.
The optical transmission characteristics were
measured using a super luminescent diode as abroadband light source and an optical spectrum
analyzer. Transmission characteristics with theultrasonic frequencies around 330 kHz are shown in
Fig. 2. The optical center wavelength of the dip is
shifted by the ultrasonic frequency. Transition of
the optical center wavelength of the dips is shownin Fig. 3 as functions of the ultrasonic driving
frequency.
Four series of rejection center
wavelength are observed and they are changed
monotonously by the ultrasonic driving frequency.

Fig. 1 Experimental setup for the ultrasonicallyinduced optical fiber long-period grating.

2. Experimental setup
A single mode fiber was used in our experiments.
A part of the coating is removed, and the bare fiber
is bonded at the top of a piezoelectric transducer asshown in Fig. 1. The transducer is driven atseveral hundred kHz: much higher than thefundamental resonance frequency of the transducer.
Then, many longitudinal resonance modes can beutilized over the frequency range for theexperiments to excite vibrations on the fiber atvarious frequencies. The silicon rubber coating of
400 µm in diameter on the opposite of the fiber side
acts as an acoustic absorber.

Fig. 2 Transmission characteristics at different
ultrasonic frequencies.

3. Optical transmission characteristics
The ultrasonic vibration decays rapidly in thecoated region, and a good traveling wave is excitedon the bare fiber [3]. The standing wave ratio was
1.4. Flexural waves are excited along the fiber inour experiment, since the measured wavelengths-----------------------------------------------------------e-mail : fukuma@sonic.pi.titech.ac.jp

Fig. 3
Optical center wavelengths vs. theultrasonic frequency.

- 253 -

4. Temporal change of the rejection bandIn the previous measurements, the opticaltransmission characteristics were observed by an
optical spectrum analyzer as time averaged results.
In this section, the temporal change of thetransmitted light intensity is observed. A tunable
semiconductor laser is used as a light source and thetransient of transmitted light intensity is recorded
by a photo detector at every wavelength just after
the voltage is applied to the transducer. The
results are shown in Fig. 4, where the excitation of
the transducer starts at t = 0. Much longer time isneeded for building up the rejection band than thetime for the propagation of the ultrasonic wave
from the driving point to the end. The rise time of
the transducer vibration is dominant, since theultrasonic transducer is driven at its resonant
frequencies.

Fig. 4 Rise of the rejection band just after
applying voltage to the transducer

5. Synthesis of the rejection band shape
A method to synthesize the arbitrary band shape is
proposed in this section. The driving frequency isswitched alternatively at a certain period so that two
different vibrations exist on the fiber at the sametime as illustrated in Fig. 5. The switching period
is so carefully chosen that the wave should travel ahalf of the fiber length for the given duration,
considering the propagation speed of the wave. In
this experiment, vibrations at 342.4 kHz and 344.1
kHz were excited intermittently on the fiber. As aresult, optical transmission characteristics shown in
Fig. 6 were obtained. The rejection band was
widened in width and had two attenuation peaks.
In this way, the shape of rejection band can besynthesized by combining several different
vibrations.

Fig. 5 A method for synthesize the rejection band
shape.

6. Conclusions
By considering the transient characteristics of theultrasonically induced long-period fiber grating, amethod to synthesize the rejection band shape wasdiscussed.
Band width was widened by
combining two vibrations with different frequencies.
To excite the vibration with enough amplitude weneed to utilize the resonance of the transducer. As
a result, large optical rejection ratio is obtained onlyat discrete resonant frequencies over the wholeband. Use of the resonance is the basic limitation
for transient response and sweep speed. A new
configuration should be introduced to achieve fully
continuous sweep and arbitrary rejection band
shape control.

Fig. 6

An example of the rejection band synthesis.

References
1. H. E. Engan, B. Y. Kim, J. N. Blake, and H. J.
Shaw: J. Lightwave Technol. 6 (1988) 428-435.
2. D. Ostling, and H. Engan: Optics Letters 20(1995) 1247–1249.
3. N. Fukuma, K. Nakamura and S. Ueha: Proc.
OFS-17, SPIE 5855 (2005) 852-855.

- 254 -

P2-22

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 255-256
16-18 November, 2005

PIEZOELECTRIC
PHOTOTHERMALAND
PHOTO-REFLECTANCE SPECTRA OF InGaN GROWNBY RADIO FREQUENCY-MOLECULAR BEAM EPITAXY○E. Kawano, Y. Uchibori, T. Shimohara, H. Komaki, R. Katayama(1), K. Onabe(1), A. Fukuyama, and T.
Ikari
Faculty of Engineering, University of Miyazaki, 1-1 Gakuen-Kibanadai-nishi, Miyazaki 889-2192, Japan(1) Graduate School of Frontier Sciences, The University of Tokyo, 5-1-5, Kashiwanoha, Kashiwa-shi, Chiba,277-8562, Japan
Abstr act
Piezoelectric Photothermal Spectroscopy (PPTS) measurements have been carried out on InGaN (In=0~0.32) thin films
grown by radio frequency molecular beam epitaxy. We found that the band energy shifts to the lower energy side (Redshift) by the increase of In composition from 0 to 0.32. For the samples with lower In composition, we could observe
the exciton contribution and the obtained binding energy was estimated to be 30 meV (In=0.01). Since, the usualphoto-reflectance (PR) spectroscopy could not observed the signals for the samples of higher In composition, theusefulness of the present PPTS methodology is concluded.

using a photo-reflectance (PR) method.

1. Introduction
Since band gap energy of nitride semiconductor InGaN
may change by the composition in a wider range from
3.5 (GaN) to 0.7 (InN) eV, it is possible to covering fromthe ultraviolet to the infrared region when this material isused for optoelectronic devices. Accordingly, InGaN isvery promising materials for light-emitting diode,semiconductor laser diode, light-detecting component,
and optical communication devices. However, thegrowthof

InGaN,

especially

for

high

Indiumcomposition, is usually very difficult. Some of thereasons are the lattice mismatch and the different thermal
stabilities of GaN and InN, and the bond lengths and thebinding energies of each atom. These produce defects

2. Exper imental Procedure
InGaN was grown on sapphire (0001) plane bymolecular beam epitaxy with RF plasma (for N2) cell
(RF-MBE) at 575 °C for 90 min. Ga/In flux and RF
plasma output were fixed at 1

2.0 sccm (1.0, 1.2, 1.5, 1.75, 2.0 sccm). Before InGaN
growth, nitridation of sapphire substrate was performedby growing thin AlN and GaN buffer layers to stabilized
the InGaN growth. The sample geometry is shown in
Fig.1. For the PPT measurements, the probing-light froma grating monochromator was mechanically chopped andfocused on the substrate side of the sample at RT.
InGaN
In=0~0.32

growth. Therefore, physical properties of InGaN did notunderstand well yet. Optical absorption spectra of InGaN
thin film was measured by using a Piezoelectric
Photothermal Spectroscopy (PPTS) [1]. The PPTS
method is known to have a high sensitivity for such thin
PPTS method, we also measured the same sample by

10-5 Pa and

400 W, respectively. N2 flow rate was varied from 1.0 tolike phase separation and dislocation during the crystal

film structures. Moreover, to clarify the advantage of the

10-4 Pa/5

300~700(nm)GaN180(nm)AlN~10(nm)Al2O3

0.42(mm)Fig.1 A schematic cross section of the InGaN thin film

The PPT signal was detected by the piezoelectrictransducer (PZT) on the film side of the sample. The- 255 -

signal was amplified by the two-phase lock-in amplifier.

semiconductor. From the PPTS results in Fig. 3, it was

The PR measurements were carried out at room

found that the band gap shifted to the lower energy side

temperature by using a standard experimental setup [2].

by the increase of In composition (Red shift). Although

3. Results and Discussion

the PR measurements could not reveal the band gap for

Figure. 2 shows the result of the PR measurements (In =In>0.13, the PPTS method could determine them. We

0~0.13). The PR spectrum was analyzed by the low-field

think the present PPTS method has advantages for

ER line-shape function, i.e., third derivative functional

measuring the absorption spectra for the crystals withform [3]. The energies of E0 critical point were estimatedcompositional fluctuation.
PPT Signal Intensity (a.u.)

to be 2.73 eV (In=0.13), 3.18 eV (In=0.05) and 3.41 eV(In=0.01), respectively. Band energy shifts to the highenergy side（Red shift) by the increase of In composition.
However, the intensity of PR spectrum decreased with
the increase in In composition, and fitting analysis
became difficult for the sample of In>0.13. We
considered that the fluctuation and phase separation of

InGaN
PPTS
RT

In=0.32
In=0.2
In=0.13
In=0.05

In=0

composition cause the decrease of the signal to noise2

ration when the In composition increased.
InGaN
PR
RT

2.4

2.8

3.2

3.6

Photon Energy (eV)Fig.3 PPTS spectra of InGaN (0<In<0.13)

In=0.13

4. Conclusion
An aim of this work was to investigate the optical⊿R/R

In=0.01

properties of InGaN by using a PPTS method. We foundthat the band energy shifts to the lower energy side（RedIn=0

shift) by the increase of In composition from 0 to 0.32.
Since the intensity of PR spectrum decreased with the

2

2.5

3

increase of In composition, we found that PR technique

3.5

Photon ener gy (eV)

is not good for the sample with compositional fluctuation.

Fig.2 Photo-reflectance spectra of InGaN (0<In<0.13)

For the samples with lower In composition, we could

Figure. 3 shows the PPTS spectra for the samples withalso observe the exciton contribution and the obtained

different In composition (0~0.32). Since the observed

binding energy was estimated to be 30 meV (In=0.01).

spectra indicate the existence of an exciton for low In

5. Referencescompositions, the Quasi-Voigt function was used for[1] A. Fukuyama, A. Memon, K. Sakai, Y. Akashi and T.

decomposing the exciton contribution and the energies of

Ikari, J. App. Phy. 89, (2001) 1751-55.

band gap were estimated.

Consequently, we obtained[2] S. Chichibu, H. Okumura, S. Nakamura, Guy Feuillet,

the binding energy of the exciton as 30 meV (In=0.01),

T. Azuhata, T. Sota, and S. Yoshida：Jpn. J. Appl. Phys.

and was almost the same as the value of 28 meV of36, (1997) 1976-1983.

GaN.

[3]D.E.Aspnes,Surface Science 37, (1973) 418-44

Since the exciton contribution decreases withincreasing the In composition above 0.05, the band gap
2

was calculated from the relations of (PPTS×hν) versus______________________________________________
kawano@pem.miyazaki-u.ac.jphν supposing that the InGaN being a direct gap- 256 -

P2-23

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 257-258
16-18 November, 2005

Calibration for Measurement of Sound Fields
Using Optical Probe◎Takeshi OHBUCHI, Koichi MIZUTANI, Shingo SHIBATA (Univ. Tsukuba)
and Hiroyuki MASUYAMA (Toba Natl. Coll. Mar. Tech.)

1. Introduction
In fields of the medical treatment diagnosis andthe industrial instrumentation etc, the measurementthat uses the ultrasonic wave is widely used from thecharacteristic such as non-contact and non-destructive.
It is necessary to measure the sound fields for thecalibrating of the ultrasonic probe. An optical probe isused for the noncontact measurement of sound fields[1-3]. They use phase deviation of light by the soundwave. The optical system such as interferometers isnecessary for that method. And a sound field also
influences light intensity [4]. In this paper, we propose
the simple system for measuring of sound fields thatuses the amplitude of optical intensity acquired by
Avalanche photodiode (APD) module.
2. Principle
Figure 1 shows the experimental systemmeasuring the sound fields. The experimental system
consists of laser and ultrasonic transducer and APD.
The laser beam passes through the water tankincluding sound fields emitted from transducer. Then,the laser beam is diffracted or is refracted. The lightthat is diffracted or refracted contain information on
the sound fields. Direction emitted laser and ultrasonicfrom are defined y-axis and z-axis, respectively. The
x-direction indicates the transversal direction of thesound. APD put in position on y-axis. Sound wave
defines a sinusoidal plane wave that propagates in thez-direction. And the sound wave emitted to water
gives a refractive index of a medium a change. The
refractive index of water is given byn = n0 + kp ( x, y, z , t ) ,(1)k=(n0 − 1)(n02 + 1.4n0 + 0.4),
(n02 + 0.8n0 + 1) ρ 0 c s2(2)where n0 is refractive index without sound fields in themedium, ρ0 is density of the medium, cs is velocity ofpressure. k is proportionality constant concerned
between amplitude of the sound pressure and changeof the refractive index [5].
The change in the refractive index
according to sound pressure causes refraction
or diffraction. When the width of laser beam is-----------------------------------------------------------ohbuchi@aclab.esys.tsukuba.ac.jp

He-Ne Laser

A.P.D.

Water TankxzyTransducer

Function Generator

Oscilloscope

Fig.1 Experimental setup
very wide compared with the wavelength, the laser isdiffracted or reflected. When frequency of theultrasonic wave is several MHz, the light is diffracted.
The intensity of incident radiation is normalized in the
maximum value. The intensity of m-th order diffracted
light Im is given by
Im=Jm2(ν),(3)
where Jm and ν are the m-th order Bessel function of
the first kind and Raman-Nath parameter. Raman-Nath parameter is given by
2π
(4)
ν =
∫ δ n ( y )dy ,λwhere λ, n and δn(y) are the wavelength of the light,
the refractive index of the medium in themeasurement space and the distribution of the changeof the refractive index of the medium on the directionof propagation of light, respectively.
On the other hand, when the width of laser beam is
not very wide compared with the wavelength of theultrasonic wave, the laser is refracted by refractive
index gradient. Light intensity acquired by the APD
includes information of sound fields. However,relation between light intensity and sound pressure is
not linear. Therefore, it is necessary to calibrate.
3. Measurements and Results
The light source is a He-Ne Laser (05LHP151 /Melles Griot). The piezoelectric transducer (NEPEC- 257 -

References[1] O. L. Lokberg: Applied Optics, 33 (1994) 2574.
[2] T. H. Neighbors, W. G. Mayer and H. J. Ruf: J.
Acoust. Soc. Am. 98 (1995) 1751.
[2] K. Mizutani, T. Ezure, K. Nagai and M. Yoshioka:
Jpn. J. Appl. Phys. 40 (2001) 3617.
[3] T. Obuchi, S. Nakanishi, K. Mizutani and H.
Masuyama: Proc. Symp. Ultrason. 25 (2004) 117.
[4] T. Obuchii, K. Mizutani, S. Nakanish and H.
Masuyama: Proc. ASJ. Spring Meet. (2005) 995(in japanese).
[5] K. L. Zankel, and E. A. Hiedemann: J. Acoust. Soc.
Am. 31 (1959) 44.

APD Output (Vp-p)

4

3

2
1

0
0

60
20
40
80
Input to Transducer (Vp-p)

100

Fig.2 Regression curve of relation between intensityof light that passed sound fields and input
voltage to transducer.
50 Vp-p
25 Vp-pMagnification

4

3
2.5

3
2
2

V50Vp-p/V25Vp-p

APD Output (Vp-p)

5

1.5

1
0
-6

-4

-2

2
0
x (mm)

4

1
6

Fig.3 Experimental result of distribution of intensityof light that passed sound fields Plot by triangle:
Input voltage to transducer is 50 Vp-p, circle:25
Vp-p, broken line: magnification.

80

50 Vp-p
25 Vp-pMagnification

60

2.5

40

2

20

1.5

0
-6
Fig.4

- 258 -

3

V50Vp-p/V25Vp-p

4. Conclusions
The distribution of the projection data of soundfields was measured. The system was able to simplify
by use of amplitude analysis. Thanks to calibration
using characteristic curve, distribution of high soundpressure was not skew. This simple system presented
in this paper proved useful to measure sound fields.

5

Estimation of Input (Vp-p)

21 / TOKIN) is an 8.0 mm aperture in diameter at thefrequency of 1.6 MHz. The transducer is set to themechanical x-stage and is scanned relative to the lightbeam. The optical intensity is measured by the APD
module (C533-03 7L-748 / Hamamatsu Photonics).
The one dimensional distribution of the opticalintensity along the x-axis is acquired by the move ofthe x-stage. The signal is observed with theoscilloscope (54645A / Hewlett Pckard). Fig.1 showsthe experimental setup.
Figure 2 shows characteristic curve for calibration.
Because of relation between input voltage totransducer and the sound pressure emitted bytransducer is linear, the output voltage from APD and
the input voltage to transducer are used for calibrating.
When the characteristic curve is acquired, transducer
is fixed at the point assumed origin. Regression curveof the characteristic curve became normal distribution.
Intensity distribution of laser is Gaussian. It is thought
to be the most likely cause of that regression curve.
Figure 3 shows projection data. The linear
scanning region of x-axial direction is from −6 mm to
6 mm with 0.5 mm step length at a distance of under
10.0 mm from the vibrating surface. The input
voltage to transducer is 25 Vp-p (○) and 50 Vp-p (△).
The broken line is magnification. There is no change
in the normalized sound fields even if the input
voltage changes. Therefore, the projection data is sure
to double, when the input voltage is doubled.
Distribution of high sound pressure is skew.
Figure 4 shows calibrated distribution. Input
voltage to transducer is estimated from Fig.1. The
estimation voltage is proportional to the projection
data when the laser passes the origin. Therefore, theestimated voltage is proportional to the projection data.
Distribution of high sound pressure is not skew.
Because of very small signal, noise thought to because of skew distribution at low sound pressure.

-4

-2

0
2
x (mm)

Calibrated result
magnificationof

4

distribution

1
6
and

P2-24

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 259-260
16-18 November, 2005

A high frequency optical scannerusing flexural vibration of an optical fiber◎Ryoichi Isago, Shunsuke Domae, Daisuke Koyama, Kentaro Nakamura and Sadayuki Ueha
(P&I Lab., Tokyo Inst. of Tech.)
1. Introduction
Mechanical optical scanners are widely used invarious applications such as bar code readers,position sensors, optically ranging devices and laser
projectors. However, it is difficult to increase thescanning angle at high scanning rate withconventional methods. This report presents amethod using a vibrating optical fiber where thelateral displacement and rotation of the fiber end
and the acousto-optic effect due to the stressinduced in the fiber core by the vibration provide alarge deflection angle of the light beam. First, theconfiguration to excite high frequency bending
vibrations on the fiber using a piezoelectric element
is described, and the relationship between theresonant frequency and the fiber length is reviewed
being based on the simple beam theory. Then, theprinciple to deflect the light beam with the vibration
and a ball lens is presented. Second, theacousto-optic effect is observed. Two-dimensionalscanning is demonstrated with the fiber vibration at
42 kHz and the ball lens displacement at 460 Hz.
2. Method and structure
Experimental setup for the proposed opticalscanner is shown in Fig.1. A silica optical fiber
without coating is bonded to the upper surface of amultilayered piezoelectric actuator. An optical fiber
is vibrated in flexural mode of cantilever beam, andoutput light beam is collimated by a ball lens. Theball lens is bonded on the end of a stainless steel
beam that is attached to another multilayered
piezoelectric actuator, and vibrates in the direction
perpendicular to the displacement of the opticalfiber. Two-dimensional scanning is possible by the
conbination of the high frequency vibration of the
optical fiber and the low frequency vibration of theball lens.
Table 1 shows the specifications of the single
mode optical fiber (F-SF, Newport Co.) used in theexperiments. A ball lens of 0.5 mm in diameter anda stainless steel beam of 0.7 mm in diameter areused. Table 2 shows the resonant lengths of theE-mail: isago@sonic.pi.titech.ac.jp

fiber and the stainless beam through the calculation
based on the Bernoulli-Euler hypothesis [1]fn =(β n L ) 2

EI
ρA

(1)
2πL
for the design frequencies 40 kHz and 500 Hz,
respectively. Here, EI is flexural rigidity, ρ isdensity, A is cross-sectional area, and β n L isfrequency constant.
2

Optical fiber

Ball lens

5 mm

Stainlesssteel beam
Multilayered
piezoelectric actuator
10 mm

Fig.1 Experimental setup.
Table 1 Specifications of the optical fiber.
Young’s modulus E
73.1 GPa
Density ρ
2220 kg/m3
Cladding diameter
125 µmCore diameter
7 µm
Refractive index of the cladding
1.453
Refractive index of the core
1.457
Table 2 Resonant lengths in mm.
Mode
1st
2nd
Optical fiber
1.6
4.0
40 kHzStainless steel beam
31.5
78.8
500 Hz

3rd
6.7
131.9

3. Optical scanning characteristics
The frequency characteristics of scanning angle
in the 3rd mode flexural vibration of the opticalfiber are shown in Fig.2 when the ball lensvibration is not activated. The maximum scanning
angle of ±10.5 ° was obtained at 41.97 kHz when
the applied voltage to the multilayered piezoelectric
actuator was 24.3 V. It is considered that excess
adhesive flowed to the fiber might rise theresonance frequency than the designed value.

- 259 -

Displacement of PZT.
● 0.27 µm (Q=362)○ 0.12 µm (Q=209)△ 0.03 µm (Q=107)Angle [deg.]

10
8
6
4
2
0
40

41

42
43
Frequency [kHz]44

Fig.2 Frequency characteristics of the scanning angle
for the vibration of an optical fiber.

4. Observation of the acousto-optic effect
The scanning angle achieved by the fiber
vibration is greater than the value expected by thegeometrical optics with the displacement androtation of the fiber end due to the vibration. It isthought that the acousto-optic effect induced by thevibratory stress in the fiber may contribute the large
scanning angle. The acousto-optic effect is observed
for an optical fiber 16 mm in length vibrating withthe 4th mode at 10 kHz. Vibration displacement andthe locus of output light spot at the fiber end isobserved by stroboscope method using a
microscope and a blinking right emitting diode. The
result is shown in Fig.3. Since the light guided by
the fiber was operated in continuous wave, thelocus of the light was recorded in the figure. The
vibration displacement of the fiber was 8 µm, while
the displacement of the output light spot was 14 µm.
Then, the relative displacement of the output
position to the fiber is 6 µm. This displacement of
the output spot position is thought to be caused byacousto-optic effect induced by the vibration.

5. Two-dimensional scanning
The frequency characteristics of the scanning
angle in the 1st mode flexural vibration of thestainless steel beam are shown in Fig.4. The
maximum scanning angle is ±6.8 ° at 462 Hz when
the applied voltage to the multilayered piezoelectric
actuator is 55.3 V. The ball lens mass lowered theresonance frequency than the designed frequency.
When both the fiber and ball lens vibrations areactivated simultaneously, the collimated light
projected on a screen at the distance of 258 mm isshown in Fig.5. The outer edges are brighter
because of the sinusoidal vibrations.
108Angle [deg.]

12

Displacement of PZT.
● 5.31 µm (Q=24.6)○ 3.12 µm (Q=34.5)△ 0.80 µm (Q=83.6)

6
4
2
0
400

420

440
460
480
Frequency [Hz]500

Fig.4 Frequency characteristics of scanning angle by
the vibration of a stainless steel beam.

±6.8 °(a) Vibration of thestainless steel beam.

±10.5 °(b) Vibration of the
optical fiber.

Locus of output
Neutral position(c) Two-dimensional scanning.
Fig.5 Scanned spot on a screen at 218 mm.
Desplaced position
Displacementof an optical fiber

Displaced position
Optical fiber

8 µm
14 µm

Displacementof output spot
Neutral position

Fig.3 Picture of the vibrating fiber.

6. Conclusion
The method for scanning the laser light in two
dimensions by combining the vibration of an opticalfiber at 42 kHz and a ball lens at 460 Hz was
examined. Two-dimensional scanning of ±10.5 ° ×
6.8 ° was achieved.
References[1] Karl F Graff: WAVE MOTION IN ELASTIC SOLIDS(Dover Pubns, 1991) pp.140-158.

- 260 -

P2-25

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 261-262
16-18 November, 2005

Trapping of particles on radiation surfaces of an ultrasonic actuator
Junhui Hu, Jianbo Yang, Jun Xu (Nanyang Technological University, Singapore)
1. INTRODUCTION
Improved particle separation techniques are increasingly important in biotechnology, water
treatment and many other areas. Ultrasonic separation of particles, which includes trapping andtransportation of particles, is very competitive in this area because it has the advantages such as
being able to handle almost all materials and having a low cost compared to the magnetic andelectrophoretic methods. In this paper, we report a method of trapping small particles on theradiation surfaces of an ultrasonic
Copper Electrodeactuator.
Piezoelectric
ring

20 mm

2. CONSTRUCTION AND PRINCIPLE
Fig. 1 shows the structure and size ofthe actuator for trapping particles. A
Polingdirection 3.0 mm
50 mm
tapered aluminum strip with a width of
4.5 mmx
4.5 mm and length of 30 mm is
30 mm
Tapered metal strip
mechanically driven by a rectangularz
Vibration
sandwich-shaped ultrasonic transducer atyo
one of its four corners. Two piezoelectricCollected particles
rings are sandwiched between two square
Fig. 1 Structure and size of the actuator.
aluminum plates by a bolt structure. The
piezoelectric rings have inner diameter of
6 mm, outer diameter of 12 mm and thickness of 1.2 mm. The
Acoustic
actuator has a resonance frequency of about 105 kHz. It wasboundary layer
observed that particles such as shrimp eggs, mint seeds andgrass seeds may be trapped on the two radiation surfaces near
L1
the tip of the strip at the operation close to the resonance point.
The out-of-plane vibration of the tip of the aluminum stripVh
was clearly observed by microscope system (Olympus BX-51),s1
measuring the position change of the metal surface in the zVndirection due to vibration. Using the method, it wasVwqualitatively found that the vibration amplitude is maximum
L3w
at the tip (x = 0), and decreases as x increases in the vicinity ofothe tip. So the strip operates at a flexural vibration mode.
L2
Using a 1mm needle hydrophone (SN945, Precision
Acoustics, UK), it was observed that the acoustic pressure p Fig.2 Acoustic boundary layer onjust under the vibrating tip in water is much smaller than that the radiation surface and a local
near the radiation surfaces. Hence, we may assume p (0, y, 0) coordinate system.
≈ 0 (1) for −2.25mm≤y≤2.25mm. On the other hand, an
acoustic boundary layer exists on the radiation surfaces [1], asshown in Fig. 2. If Vh, Vw and Vn are L1, L2 and L3 components of the fluid vibration velocity at
the interface between the boundary layer and the metal strip (L3 = 0), respectively, thenVn ≠ 0 (2) Vh ≈ 0 (3) Vw ≈ 0 (4).

Using Eqs. (1)-(4) and grad p = − ρ 0

∂Vf
∂t

(5) ,

where Vf is the fluid vibration velocity, one can deduce that the acoustic pressure at the interface
between the boundary layer and the metal strip is approximately zero. Due to this nodal plane,
small particles are sucked to the radiation surfaces [2, 3]. To verify the existence of the acoustic
radiation force which pushes small particles onto the vibrating strip, an experiment was
conducted, as shown in Fig. 3. In the experiment, the radiation surfaces were perpendicular to- 261 -

gravitation. The trapping of shrimp eggs on the lower radiation surface due to the strip’s vibration
was observed in air.
Trapped
3. EXPERIMENTAL METHODS
particles
AND MATERIALSGravitation
In the following experiments, thelength direction of the metal strip
is parallel to the gravitation; mintseeds were used, which has anaverage density of 0.7 g/cm3,
Direction of
volume per particle of 0.07×10-3
3-5
photographing
cm , mass per particle of 4.9×10
g and diameter of 513 µm in a dry
state.
(a)(b)
4. CHARACTERISTICS AND
Fig.3 Verification of the existence of the acoustic force for
DISCUSSION
trapping. (a) Experimental setup; (b) photo of the trapped
Fig. 4 shows the trapping ofparticles.
mint seeds in water. It wasobserved that the distributionof trapped particles is quite uniform in the trapping area. This provides another experimental
evidence of the existence of the nodal plane of acoustic pressure on the metal strip. Such a
distribution would be impossible if the trapping were due to some isolated nodal points or lines in
the near field or far field. Fig. 5 shows the number of trapped mint seeds versus operating
frequency in water and air for a given driving voltage. It is seen that more mint seeds can betrapped in water than in air and the frequency range in which theactuator can trap particles is wider for the collection in water than in air.
There are two possible reasons for these phenomena. One is that thematching of acoustic impedance between water and aluminum is better
than that between air and aluminum. Another is that the buoyancy
acting on the particles improves the capability of trapping particles.
5. SUMMARY
We have proposed a new method which uses the radiation surfaces ofan ultrasonic actuator to trap small particles. Compared with other Fig. 5 Trapping ofmint seeds in water.
techniques of trapping particles, ourtechnique does not need a specially designed
chamber to generate the necessary soundfield, and can easily transport trapped
particles to a desired location for separation
by simply moving the actuator.
REFERENCE:[1] P. M. Morse and K. U. Ingard,
Theoretical Acoustics. NY: McGraw-Hill,
1968, p. 243 and p. 279-287.
[2] T. Hasegawa and K. Yosioka, J. Acoust.
Soc. Am., vol.46, pp.1139-1143, 1969.
[3] K. Beissner, J. Sound and Vibration, vol.
93, No.4, pp. 537-548, 1984.

Fig. 6 Number of trapped mint seeds versus
driving frequency.

- 262 -

P2-26

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 263-264
16-18 November, 2005

Phase shift of Rayleigh wave beneath slider with preload◎ Yoshito Miyazaki, Takashi Shigematsu, Minoru Kuribayashi Kurosawa (Tokyo Inst. of Tech.)
1. INTRODUCTION
An extremely high performance surface
acoustic wave (SAW) linear motor operating at 10
MHz has been already demonstrated [1,2]. However,
the efﬁciency from the electrical input to mechanicaloutput is still low. To improve the efficiency, anenergy circulation driving method has proposed anddemonstrated [3]. For the energy circulation driving
method, forward scattering beneath a silicon slider
was found to be a problem, due to the shift of thephase relation. Numerical simulation was carried out
to investigate the relationship between the phase shift
and slider parameters using the ﬁnite element method.

E2 = E0 cos ω t

E1 = E0 sin ω tRayleigh wave

Reflector

StatorFig. 1: Schematic view of Rayleigh wave motor statorusing an energy circulation driving method.
Slider
(0.6 mm)

Stator4 mm

1.9 mm{Driving points

2. ENERGY CIRCULATION AND WAVE SCATTERINGThe schematic view of the Rayleigh wave
motor stator using the energy circulation driving
method is shown in Fig 1. [3] The traveling SAW is
excited by putting a sine signal and a cosine signal into
the driving IDTs respectively. The excited travelingwave is received by one unidirectional IDT, andconverted into electric energy. The converted electricenergy is returned to the after unidirectional IDT.
Using this electric energy, the other unidirectional IDT
excites a circulated traveling wave. The circulated
wave is propagated, and received by the initial
unidirectional IDT again. Then the circulated wave
and the initially excited wave are superposed. It isvery important that these two waves are in phase at adesigned value to efﬁciently excite the Rayleigh wave.
On the other hand, many projections were
fabricated on the contact surface by dry etching.
Because of these projections, the sufficient contactpressure is generated and the better contact condition
is obtained. The Rayleigh wave beneath the slider
is scattered by these projections when the slider ispreloaded. As the result, the phase of the wave at
the rear of the slider shifts from the designed value.
This phase shift affects on the superposition of
the circulated wave and the initially excited wave.
Namely, the power of the Rayleigh wave decreases
compared to the complete accord phase condition. As
a result, this phase shift decreases the efﬁciency of the
motor.
y.miyazaki@ae.titech.ac.jp- 263 -

Unidirectional IDT

1 mm

1.8 mmDamping blocks

7.6 mmFig. 2: FEM simulation mesh model of Rayleigh wave
propagation.

3. SIMULATION
To investigate the phase shift of Rayleighwave, simulation was carried out using the finite
element method. The software was ANSYS 7.0.
The FEM stator and slider model is shownin Fig. 2. The model consisted of slider, stator anddamping block. To prevent the reﬂection wave, the left
side of the model had symmetrical boundary conditionand damping elements were attached on the bottom
and the right side of the stator. The dimensions of thedamping element areas were 1.8 x 7.6 mm2 on thebottom and 1.0 x 1.9 mm2 on the right. Four nodes 2-D
plane elements were used in the whole model. The
mesh size of the surface of the stator was 5 µm, thatwas 1/80 of the 400 µm wave length.
The stator was made of 127.8 degrees
Y-rotating X-propagation LiNbO3 [4]. The slider was
made of silicon. Its density was 2.33 x 103 kg/m3 ,Young’s modulus was 1.66 x 1011 Pa, shear modulus
was 0.639 x 1011 Pa. Many projections whose height
were 1 µm were fabricated on the surface of the slider.

Phase shift [degree]

15
10
5
0
-5

Fig. 3: FEM simulation of Reyleigh wave scattered byslider projections.

100
Phase shift [degree]

0

2

4

6 8 10 12 14 16
Gap [nm]Fig. 5: Phase shift aginst gap between the stator andthe slider caused by the silder projections.

120

When the preload was too low, the slider was
pushed by Rayleigh wave and took off the stator. So
soft damping block was mounted on the slider andthe upper surface of the block was ﬁxed. The preload
could be controlled by changing the gap between thestator surface and the slider projections. Phase shift
against gap between the stator and the slider is shownin Fig. 5. Large gaps, namely low preloads, caused theminus phase shit.

80.0
60.0
40.0
20.0
0.00

-10

0

1

4
3
2
Pressure [MPa]56Fig. 4: Phase shift against preload caused by the silder
projections.

The diameter of the projections was 50 µm, the pitchof the projections was 100 µm. Contact elements were
mounted the surface of the slider and the surface of theslider projections. Its contact stiffness was 0.001.
The simulation type was two dimensionalstrain structural-piezoelectric analysis, transient
analysis. The amplitude of 20 nm vertical sinusoidal
displacement at 9.945MHz was given at 5 driving
points.

4. RESULTS OF SIMULATION
The integral time span was 3ns which was
about 1/ 30 of 1 cycle of the wave. The simulation was
carried out for about 1.7 msec, that was the 17 cyclesof the wave motion. On this condition, we could obtainthe traveling Rayleigh wave. The phase velocity of theRayleigh wave of this simulation was 4150 m/sec withno slider.
The deformation model of the simulation
result with slider preload at 1.875 MPa is shown in
Fig. 3. The phase shift against the preload is shown in
Fig. 4. The large preload caused the large phase shit.

5. CONCLUSION
The FEM simulations were carried out toinvestigate the relationship between the phase shift andthe projection parameters of the slider. Phase shift wascaused by the contact of the slider projections. It isnecessary to determinate the relationship between thephase shift and the projection parameters of the slider
quantitatively for the next step.
REFFERENCE[1] Y. Nakamura, M. K. Kurosawa, T. Shigematsu, andK. Asai, “Effects of Ceramic Thin Film Coating on Friction Surfaces for Surface Acoustic Wave Linear Motor,”
Proc. of IEEE Ultrasonics Symp, P2M-3, pp. 1766-1769.
Honolulu, Oct 5-8, 2003.
[2] T. Shigematsu, M. K. Kurosawa and Katsuhiko Asai,“Sub-nanometer Stepping Drive of Surface Acoustic
Wave Motor,” Proc. IEEE-NANO 2003, WE 3, San
Francisco, August 12-14, 2003.
[3] K. Asai, M. K. Kurosawa, “Surface acoustic wave
motro using an energy circulation driving method,” Proc.
of IEEE Ultrasonic Symp. , P2H-5, pp. 525-529 , Atlanta, Oct. 7-10, 2001.
[4] J. Kushibiki, I. Takanaga, M. Arakawa, and T. Sannomiya, “Accurate Measurements of the Acoustical
Physical Constants of LiNbO33 and LiTaO33 Single Crystals,” IEEE Transactions on UFFC, VOL. 46, No. 5, Sep.
1999, pp.1315-1323.

- 264 -

P2-27

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 265-266
16-18 November, 2005

Study of a mini-ultrasonic motor with square metal bar and piezoelectric plate hybrid
Yu Chen, Kai Lu, Tieying Zhou，Tao Liu，Cunyue Lu
Department of physics , Tsinghua University ,Beijing 100084, P R China
A new mini ultrasonic motor with metal bar and piezoelectric plate hybrid as a stator has been studied. Four
piezoelectric plates are bonded to four sides of the metal bar. The metal bar is made of brass and the piezoelectric plates
are made of PZT-4. The polarization direction is vertical to the piezoelectric plate. Compared with the cylindrical type ofthe motor, the square cross-section motor makes polarization more sufficient and signal wire easier to be welded. In thispaper, the modal and harmonic analysis of the stator were performed by using FEM. The resonance frequency of the
motor is about 100 kHz, the no-load revolution is above 6000 r/min and maximum torque is about 60 uNm. This newmini ultrasonic motor is of potential applications.
KEYWORDS: metal bar, ultrasonic motor, piezoelectric plate, bending vibrating
The ultrasonic motor has many excellent
performance features, such as high torque at low speed,
quick response, position with high accuracy, simple
construction, as a result it has the extensively applied
foreground in the fields of industry control, precise
instrument, automobile electric appliances etc. The
bending vibration piezoelectric motor makes use of thevibration and friction force to drive the rotor. Its
outstanding characters are small volume, working withlow energy input and easy to fabricate compared withother types of ultrasonic motors. [1-6]. In our early
work a 1mm cylindrical ultrasonic motor was used in an
endoscope OCT [7,8].
There are various types of bending vibrating ultrasonicmotors, e.g. using piezoelectric slice, piezoelectric
cylinder and tube. Although the types USM of
piezoelectric cylinder and tube have high density of
torque and high speed, they both have a curved surface.
A novel ultrasonic motor with square cross section metalbar and piezoelectric plate hybrid is proposed in thispaper. Compared with the above two, the metal bar one
is easier to fabricate and the thin piezoelectric plates canbe polarized more sufficient; signal wires are easier tobe welded , so that the ultrasonic motor with a squarecross section is easy to be mass-produced.

respectively. Two bending modes excited at two
perpendicular directions, are coupled to form a rotation
bending mode. The surface motion of the stator is a
traveling wave of one wavelength.
Next the rotating stability of the motor is considered.
We examine the moments of inertia in each direction. A
coordinate system is set as in Fig.2(a).

Fig.2 the coordinate system
A rotating coordinate system is x1, y1, as shown in
Fig.2(b) and we have following relationship for two
systems⎧ x1 = x cos α + y sin α⎨⎩ y1 = y cos α − x sin α(1)

The moments of inertia of a bar cross section are givenby [9]

I x = ∫ y 2 dA,

I y = ∫ x 2 dA

A

(2)

A

where dA is the area element. The product of inertia is

I xy = ∫ xydA

(3)

A

The moments of inertial in the x1 –y1 coordinates is2

I x1 = ∫ y12 dA = ∫ ( y cos α − x sin α ) dA
A

(a)
(b)
Fig.1(a)the photo of stator (b)the cross section of stator
The size of the metal bar is 1.6x1.6x6.6mm3. Figure1(b) illustrates the cross section of the stator. Four
piezoelectric plates denoted by 1,2,3,4 in Fig.1(b) arebonded to four sides of the metal bar. The size of eachplate is 1.6x6x0.3mm3. The metal bar is made of brassdenoted by 0 in Fig.1(b) and the piezoelectric plates are
made of PZT-4.
The polarization direction is vertical to thepiezoelectric plate and they are bonded as shown in
Fig.1(b). Four signals which are sinωt、cosωt、-sinωt、-cos ω t, are applied onto four piezoelectric plates= I y sin 2 α + I x cos 2 α − I xy sin 2α(4)
1 − cos 2α
1 + cos 2α+ Ix− I xy sin 2α22
Ix + I y Ix − I y=+cos 2α − I xy sin 2α
2
2

= Iy

Similarly we have

- 265 -

I y1 =

Ix + I y
2

I −I− x y cos 2α + I xy sin 2α2(5)

For a bar with a square cross section we have

Ix = I y =h4, I xy = 0
12(6)

so that4h= Ix = I y
12

I x1 = I y1 =(7)

It has been proved that the moments of inertia of a metalbar are equal in all directions.
On the x-z plane, the curvature, 1/ρx, caused by thebending moments M, can be calculated by the following
equation

1

=−

ρx

M
EI y

(8)

where E is Young’s modulus of the metal. In the x1-y1
coordinate system we have a same formula

1

ρx

=−

1

M
EI y1

(9)

Since the moments of inertia of a metal bar are equal in
all directions, the curvature of a square bar on the
different plane is the same for a constant M. This result
is the same as that of the cylindrical bar. Thus we haveshown that the bending rotation of a square bar is stable.
If we apply the electrical signals on the fourelectrodes the particle motion track on the surface of thestator is an ellipse, which is the same as that of thecylindrical motor [10,11].

(a)(b)
Fig.3 Finite Elements Analysis (a) modal analysis(b) displacement distribution of harmonics analysis
From a Finite Elements Analysis, the result showsthat the resonance frequency of the first bending modefor 2 ends free boundary condition is 114 kHz. The
displacements of the surface particle of the bar along thez axis at the maximum harmonic responding is shown in
fig.3(b) and its maximum displacement of an end point
is about 0.2 um when the applied voltage is 5 V.
Table 1 the measurement result
Frequency kHz
100
95
Effective
V
93.6
90.8
voltagePower
Rotating
speed

90
88.7

W

0.177

0.138

0.118

rpm

6400

3800

2100

The performance of the motor was measured by useof the YOKOGAWA WT1600 power meter. The
parameters obtained are listed in Table 1. The stalling
torque of the motor is about 60uNm.
An ultrasonic motor with a square cross section has
good potential applications in many fields such as
micro-machine, medical machine. The study of using
this motor in an ultrasonic gastroscope is underway.
Acknowledgements
This work was supported by NSFC 50235010.
REFERENCES ：[1]S.Ueha,Y.Tomikawa,M.Korosawa,K.Nakamura.Ultrasonic Motors: Theory and Application.Clarendon
Press,Oxford,1993 p183-184[2] T. Morita,M. Kurosawa and T.Higuchi.A Cylinder
Micro-ultrasonic Motor,Ultrasonics,2000,38(1):33-36[3] Cagatay S., Koc B., Uchino K. A 1.6-mm,metal tube
ultrasonic Motor. UFFC,2003,Vol.50 No.7:782-786[4] Shuxiang Dong etal. Piezoelectric ultrasonicmicromotor with 1.5 mm diameter, UFFC,2003,Vol.50
No.4:361- 367[5]. Zhou Tieying, Zhang Kai, Wang Huan, et al., Studyon piezoelectric cylinder micro ultrasonic motor with
1mm diameter, ACTA ACUSTICA (in Chinese） 2004,
29(3): 258-261.
[6]. Zhou Tieying, Zhang Kai, Jiang kaili, Bending
vibration piezoelectric cylindrical-type micro ultrasonicmotor, Chinese Patent , Application No. 01109049.9（inChinese）2001-01-09.
[7] Zhou Tieying, Liu Xinguo, Chen Yu, etal. ultrasoundelectromotor as the endoscopes driving system,
International Medical Devices （in Chinese）, 2004, 10(1): 10-13.
[8] ZHOU Tieying, ZHANG Kai, CHEN Yu etal, Acylindrical rod ultrasonic motor with 1 mm diameter andits application in endoscopic OCT, Chinese Science
Bulletin Vol. 50, No. 8,April 2005 pp826-830 (inEnglish)[9] Hand Book of Material Mechanics, interpreted by
Fan Xinshan Architecture industry published 1987 (inChinese)[10]. Zhao Xiangdong, Liu Changqing, Zhao
Chunsheng, The movement mechanism research of
cylinder bending travelling wave piezoelectric ultrasonicmotor, Journal of Applied Mechanics （in Chinese）,
2000，17（3）: 120-123[11]. Yang Quangang, Zhang Kai， Zhou Tieying，et al.,Analysis and experiment on the taper angle of contact
surface of cylindrical USM, ACTA ACUSTICA. （inChinese） 2002, 9(5): 413-419.

- 266 -

P2-28Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 267-268
16-18 November, 2005

Analysis of tiny piezoelectric ultrasonic linear motor◎*Hyun-Phill Ko,Chong-Yun Kang, *Sangsig Kim and Seok-Jin Yoon
(Korea Institute of Science and Technology, *Dept. of Electrical Eng. Korea Univ.)

1. Introduction

Shaft

Radial direction

In recent years, a novel tiny piezoelectric ultrasoniclinear motor converting a radial mode vibration to alongitudinal vibration mode driven by impulse
force has been developed for a camera opticalmodule. In this work, the motors are investigatedwith various shaft materials in order to improve thedynamic properties of the motors. The shaft
material has direct influence on efficiency,
reliability and quality of the motors and theirdynamic properties. In this study the frictional
properties of the shaft materials were intensively
considered by changing the materials which havedifferent coefficient of friction, and the optimal
combinations of material for shaft wererecommended[1]. The characteristic of movementwith applied voltage waveforms was also
investigated to suggest optimized movement
condition of the developed tin piezoelectric linear
motor. Finally, the tiny ultrasonic linear motors’
performance with various shaft materials in termsof velocity, maximum force, ideal working
frequency and displacement of the shaft isinvestigated and discussed.
2. Construction and Experiment
The tiny piezoelectric linear motor using the inertia
movement described above is shown in Fig.1.
The motor consists of three parts which are atransducing part, shaft and mobile element. Tetransducing part consists of a piezoelectric ceramic
disk, and a metal disk attached on the piezoelectric
ceramic disk. The transducing part is driven by thesawtooth electrical potential applied to the electrodeon the piezoelectric disk. The piezoelectric ceramic
disk generates the one of the radial modes which isperiodical extension and construction on the radiusdirection of the piezoelectric ceramic disk. The
transducing part is made of round plate, generating
flexural vibration by means of the piezoelectricMovingelement

Piezoceramic

Elastic material

Metal disk

Ceramic

Cap

(a)

(b)

Fig. 1. Configuration of the (a) transducing part and(b) tiny ultrasonic linear motor
ceramic disk and metal disk together. The flexuralvibration which has negative influence on
movement of the transducing part without theconstraint. Constraining the side of piezoelectric
ceramic in this motor, radial mode movement,
which is one of the vibration modes, is converted
into up-and-down movement [2].
In this investigation, four types of materials areused for the shaft, stainless steel coated with DLC(Diamond Like Carbon), uncoated stainless steel,
Pyrex and Graphite As a moving element, a
rectangular stainless steel, which is compressed to asurface of the shaft by rubber tube, is used. The
position of the moving element is detected withlaser interferometer and the position data areconverted into velocity data by using a homedeveloped frequency to voltage converting circuit.
The finite element method (FEM) is used for thevibration analysis of the motor. The finite element
code ATILA is applied to the design of this motor
because it takes into account piezoelectric effectand electromechanical coupling coefficients of the
motor. As we use the FEM, we can assume to the

Author E-mail: ryanko@kist.re.kr- 267 -

Table 1.Mechanical resonant frequency

48

50

Pyrex

Graphite

55

64

Velocity (mm/s)

DLC

15
10
5
0
-5
-10

47

52

10
5
0
-5
-10
-15
-20
-25

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

-30
-0.5

4.0

30

15

Velocity (mm/s)

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

Graphite shaft

25

20

theoretical resonant frequencies of the motors.

0.0

30

Pyrex shaft

25*TFeq.: Theoretical mechanical resonant freq. (ATILA)
EFeq : Experimental mechanical resonant freq.(HP4192)

10
5
0
-5
-10

20
15
10
5
0
-5
-10

-15

-15

-20

-20

-25

-25

-30
-0.5

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

-30
-0.5

4.0

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

Fig. 2. The velocity of the moving element at eachposition of the shaft

Experimental Result and Discussion

20

115

Velocity
Force

110

18
105
16

100
95

14
90
12

85
80

Force (mN-m)To operate motors with various shafts, operating
frequencies of motors with the moving element aremeasured by impedance analyzer as shown in Table
1. The displacement characteristics of the moving
element are measured with a laser interferometer. A
laser beam is irradiated to the mirror attached to themoving element of each motor. The velocity of themoving element is measured at each position of the
shaft, as shown in Fig.2. The length of each shaft ofmotors is 6mm. As indicated in Fig.2, the results
show that the moving element of the motor
performed a stable linear motion at pyrex shaft, as
expected. Also, Fig 2 shows slightly unstable speed
and linear motion characteristics of motors with
DLC coated and Sus shaft. Fig. 3 shows thevelocity and the force of the moving element ofmotors according to various shaft materials. We
operated each motor with its own operating
frequency shown in Fig. 3 and measured the
dynamic characteristic of each motor.
As shown in Fig. 3, when the motor with stainlesssteel shaft is driven at 37 kHz, the motor showed
the velocity and force of 10mm/sec, 90mN-m,
respectively. In the motor with stainless steel coatedwith DLC at 35 kHz, its velocity is 15mm/s and its
force is 75 mN-m. When the motor used the Pyrex
shaft, the force of 110mN-m is reached at 37 kHz.
This motor has a high force and stable velocity
value, because the shaft with Pyrex materials has asleek surface condition. Although the motor withstainless coated with graphite shaft is shows good
dynamic characteristics, after few times operating,motor shows the unstable movements again because
the wear starts and occurs between graphite shaft
and moving element.

Velocity (mm/s)

3.

15

-20
-25

63

20

-15

-30
-0.5

45

SUS shaft25

20

Velocity (mm/s)

T Freq.
(kHz)
E Freq.
(kHz)

Stainlesssteel

DLC shaft25

Velocity (mm/s)

Shaft material
type

30

30

10
75
8

70

SUS

DLC

Pyrex

Graphite

MotorsFig. 3. The velocity and the force of the moving
element of motors
4.

Conclusion

We investigated the velocity and force between theshaft and moving element with a various shaft
materials. When being compared with thepreviously proposed prototype motor, the overall
size of the motor decreases but dynamic properties
and stable operation are improved. The
experimental results confirm that our present motor
has significantly enhanced performance over theprevious motor. It was found that two combinations,
namely Pyrex or graphite for shaft and stainlesssteel for the moving element, exhibits good andstable dynamic behaviors in an ultrasonic linear
motor.
References
1. H.P.Ko et al. Mater. Chem.and Phys., 90 (2005)
322.
2. H.P.Ko, C.Y.Kang ,S.Kim and S.J.Yoon：Sensors
an Actuators:A (accepted SNA4867)- 268 -

P2-29Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 269-270
16-18 November, 2005

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P2-30

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 271-272
16-18 November, 2005

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P2-31

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 273-274
16-18 November, 2005

Tunable Vibration Absorber Incorporating
Piezoceramic Sensoriactuator
Deyu Li, Siu Wing Or* and Helen Lai Wa Chan (The Hong Kong Polytechnic University)
Ping Kong Choy and Peter Chou Kee Liu (ASM Assembly Automation Ltd.)
1. Introduction
Tuned vibration absorbers are pre-tuned passive
vibration absorption devices comprising predeterminedsets of passively damped masses and springs to reduce
the susceptibility to vibrations of a structure at aparticular resonance frequency [1]. Tunable vibration
absorbers (TVAs) often utilize smart materials as theprincipal tuning means to provide the distinct feature of
tunable stiffness via an external tuning signal [2–4]. With
such tuning means, one may easily and precisely adjust
the natural frequencies of such TVAs to optimally matchthe targeted resonance frequencies of the corresponding
under-damped structures so that additional energy
absorption can properly be introduced to the structures
without adding any extra component and affecting thevibration levels of the structures at other frequencies.
While effectiveness and usefulness, the existing TVAs
are incapable of self-sensing structural vibrations for
implementing real-time, closed-loop tuning of theirnatural frequencies; they are limited to an open-loop
mode of operation instead. Although it is possible toestablish a closed-loop operation via an installation ofvibration sensors (e.g., accelerometers) on the vibrating
structures, this method not only increases the engineering
cost, but also leads to difficulties in assuring a reasonably
high degree of absorber-sensor collocation. In view ofthis, we have developed a smart TVA, consisting of apiezoceramic sensoriactuator suspended in a rigid frameby two flexible beams coupled longitudinally to the axialends of the sensoriactuator and the frame, to provide thereal-time, closed-loop tuning capability in a simple,
cost-effective and reliable manner [5]. In this paper, theworking principle and structure of the smart TVA are
described and its tunability and sensibility are reported.

2. Working Principle
The smart TVA is conceptualized using the sensitivityof the bending stiffness and natural frequency of a pair of
flexible beams in response to a force exerted axially tothe beams. Figure 1 shows the conceptual design of thesmart TVA. It basically consists of a lumped, damped
mass (2M), two flexible beams and two rigid bases. Eachof the beams serves as a spring in the transverse direction(i.e., the y-direction) to support the damped mass at one
side. These beams, at the same time, are subject to anaxial actuation force (F) provided by the damped mass.
The beams are assumed to be made of the same material
having Young’s modulus E and density ρ. Besides, they
have the same length L and the same rectangular--------------------------------------------------------------------------------* E-mail: apswor@polyu.edu.hk

Fig. 1. Conceptual design of the smart TVA.

cross-section A of width b and height h. The axial
actuation force on the beams is taken to be compressive(i.e., F is positive). In fact the axial stiffness of the beams(EA/L) is much larger than its bending stiffness (48EI/L3).
The first axial natural frequency of such a beam-masssystem is much higher than its first bending natural
frequency so that the TVA essentially operates in its firstbending mode with the suspended damped mass
vibrating along the transverse direction. The natural
frequency f of the TVA can be calculated using [6]

f =

1
2π

2

π 2  π 
 EI   − F 
L   L 
3ρAL + 8M,(1)

where I = bh3/12 is the inertial moment of the rectangularcross-sectional area.

3. Design Prototype
The fabricated smart TVA is shown in Fig. 3. It has
dimensions 110 mm (length) × 40 mm (wide) × 40 mm(height). The lumped, damped mass illustrated in Fig. 1
was replaced by a piezoceramic sensoriactuator toprovide the attractive functionality of tuning-whilesensing. The sensoriactuator was formed by sandwichinga CETC WTYD1012010 PZT multilayer-stack actuator(20 mm × 10 mm × 10 mm) between two PKI502 PZTpatch sensors (1 mm × 10 mm × 10 mm) and thenclamping the whole actuator-sensor assembly by two
SST304 back plates. The sensoriactuator not only is used
to provide an axial actuation force to the beams (i.e., via
the stack actuator) and to monitor the amount of the axial
actuation force exerted on the beams (i.e., via the two
patch sensors), but also is to serve as a damped mass for
the TVA. Through adjusting the electrical voltage (i.e., a- 273 -

Rigid Frame

5. Conclusion
A novel TVA has been developed and experimentally
evaluated. The TVA consists of a piezoceramic
sensoriactuator suspended in a rigid frame by two
flexible beams coupled mechanically to the axial ends ofthe sensoriactuator and the frame. The measured natural
frequency of the TVA is 152.5 Hz at zero volts. A
significantly reduced natural frequency of 135 Hz,
corresponding to a high tunability of natural frequency of~12%, has been achieved upon the application of a

Flexible Beams

Fig. 2. A smart TVA prototype.

4. Tunability and Sensibility Measurements

100137.5 Hz 145.0 Hz

152.5 Hz

0V
46 V
62 V
92 V

135.0 Hz
-10
-20
-30
-40
100

120

140

160

180

200

Frequrency (Hz)Fig. 3. Measured FRFs of the smart TVA with different tuning
voltages.
30

Output/Input (dB)The smart TVA was installed on a LDS V406 shaker
for evaluating its tunability and sensibility. A SigLab
50-84 dynamic signal analyzer was used to generate a
random analog signal ranging from 0 to 1 kHz to a LDS
PA100E power amplifier so as to excite the shaker toprovide a reference vibration to the smart TVD. An
ENDEVCO ISOTRON 256HX-10 accelerometer was
mounted on the surface of the sensoriactuator of the TVA
to measure the acceleration responses along thetransverse direction. The SigLab was also used to gather
the frequency response functions (FRFs) of both the
input (i.e., the excitation signal) and output (i.e., thepatch sensors and accelerometer signals) signals. A DC
signal generator (Stanford Research DS340) and a power
amplifier (NF 4025) were employed to drive the stackactuator of the sensoriactuator to produce a compressive
axial force for tuning the natural frequency of the TVA.
Figure 3 shows the measured FRFs of the smart TVA
with different tuning voltages. The observed natural
frequency of the TVA at zero tuning voltages is 152.5 Hz.
As the tuning voltage is increased, the stack actuator
expands and exerts a larger compressive axial actuation
force on the beams, thereby reducing the bending
stiffness of the beams and hence the natural frequency of
the TVA. It is clear that the measured natural frequencyof the TVA is 135.0 Hz at an increased tuning voltage of
92 V, giving a large frequency tunability of ~12%.
The FRFs measured from the patch sensors and theaccelerometer at a tuning voltage of 62 V are plotted inFig. 4. The resonance frequency (i.e., 137.5 Hz) detectedby the patch sensors is the same as that captured by theaccelerometer. This good agreement suggests that thepatch sensors can effectively replace the commercial
accelerometer for automatic tuning purposes.

Piezoceramic
Sensoriactuator

Rigid Bases

Acceleration/Input (dB)tuning signal) applied to the stack actuator of thesensoriactuator to change the axial actuation force to thebeams, the transverse bending stiffness of the beams ischanged and the natural frequency of the TVA is tuned.
The variation of the axial actuation force and hence thetransverse bending stiffness and natural frequency areautomatically monitored by the patch sensors. It is
important to note that when the targeted resonancefrequency of the corresponding under-damped structure
is reached, the outputs of the patch sensors are at
maximum, indicating that the TVA has been absorbing
the vibrational energy of the structure. The two flexible
beams were fabricated with SST304 and have the same
dimensions of 32 mm long, 6.5 mm wide and 0.56 mm
thick. The rigid bases mounted on the rigid frame werealso made of SST304.

20
Patch Sensors

10
0137.5 Hz

-10
-20
-30
-40
100

Accelerometer
120

140

160

180

200

Frequency (Hz)Fig. 4. Representative FRFs measured from the patch sensors andthe accelerometer at a tuning voltage of 62 V.

relatively low tuning voltage of 92 V. The FRFs as
measured using the patch sensors agree well with thoseobtained using a commercial accelerometer, indicating agreat promise of deploying the proposed TVA for
real-time, closed-loop vibration absorption.

Acknowledgments
This work was supported by the Innovation andTechnology Fund (UIM/117) of the HKSAR Government
and the Centre for Smart Materials of The Hong Kong
Polytechnic University.

References
1. J.B. Hunt: Dynamic Vibration Absorbers (The Garden City
Press Limited, Great Britain, 1979).
2. A.B. Flatau, M.J. Dapino and F.T. Calkins: J. Intell. Mater. Syst.
Struct. 11 (2000) 923.
3. N. Jalili and D.W. Knowles IV: Smart Mater. Struct. 13 (2004)
998.
4. E. Rustighi, M.J. Brennan and B.R. Mace, Smart Mater. Struct.
14 (2005) 19.
5. S.W. Or, D. Li, H.L.W. Chan, P.K. Choy and P.C.K. Liu: US
Patent 11/051,155 (2005).
6. D. Li, S.W. Or, C.S. Yung, H.L.W. Chan, P.K. Choy and P.C.K.
Liu: Proc. 12th SPIE Annual Int. Symp. on Smart Struct. Mater.
(2005) 5760.

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Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 275-276
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P2-33

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 277-278
16-18 November, 2005

Characteristics of High Frequency Sound Irradiated from
Electromagnetically Driven Plate
Masashi Ohara, Takashi Kubo, Jian Yang, and Mamoru Kuwabara(Dept. of Materials, Physics and Energy Eng. Nagoya Univ.)

Abstract
An ultrasonic oscillator, which electromagneticallygenerates ultrasound by vibration of an iron plate in
the alternating magnetic field, is newly proposed forhigh temperature use. Low Curie point for a normaltransducer made of piezoelectric material is overcome
by use of a ferromagnetic iron plate that has a highmagnetic Curie temperature (1043 K). The joule
heating of a ferromagnetic iron plate is observed by
the infrared camera, and the heating rate of aferromagnetic iron plate is calculated. The rate of theJoule heating is closely related to the Lorentz force
with which the plate is electromagnetically vibrated.
From the viewpoint of Joule heating, the effects ofvarious process parameters on the sound pressure are
discussed. The magnetic field can be concentrated on
the center of the iron plate by setting the water-coolediron pole in the induction coil, and it is expected thatthe iron plate can be vibrated more strongly.

1. Introduction
In recent years, the present authors [1] havedevoted to developing ultrasonic technologies formaterial processing, which is named “Sonoprocessingof materials”. It is considered that various applied
technologies in material processing, such as
convection control of molten metal and inclusion
removal in refining processes, dispersion of reinforced
particles and orientation control of reinforced fibers incomposite material engineering, can be developed byusing the external force of an ultrasonic wave
effectively[2]. The transducers made of piezoelectricmaterials are usually used for generating ultrasonicwaves. But they cannot be directly used in a hightemperature field because its Curie temperature isrelatively low (about 600 K). An alternative way may
be possible by use of a ferromagnetic iron plate which
has a high magnetic Curie temperature of about 1043
K and introducing liquid coolant in order to cool thesteel plate below the Curie temperature. In this study,
an ultrasonic oscillator, which electromagnetically
irradiates ultrasound by vibrating a steel plate in analternating magnetic field generated from an induction
coil, is newly proposed.

In the present paper, the iron plate is vibrated by thestrong electromagnetic field. Transient change intemperature of the electromagnetically driven plate ismeasured by an infrared camera, and the heating rateof the plate is calculated. From the viewpoint of theJoule heating associated with the Lorentz force todrive the plate, the effects of various processparameters on the sound pressure are investigated.
2. Experimental setup
The experimental setup is shown in Fig.1. The
alternate current, which was generated by a highfrequency (440 kHz) and high power (4 kHz) signal
generator, was fed to the water-cooled coil (7 turns, 60
mm in height, 50 mm in diameter). Circular iron platewith different thickness was horizontally placed above
the coil. The iron plate was vibrated by the Lorentz
force and heated with electromagnetically induced
current, which was generated by an alternating
magnetic field from the coil. Distributed thermal levelof the iron plate was observed by an infrared camera.
Temperature change with time was also measured at
specified positions from the center of the iron plate,which were changed at intervals of 30 mm, and theheating rate of the iron plate was calculated. The
measurements described above were performed bychanging the iron plate thickness t and the distance
between the coil and the iron plate d.
Processor
Infrared camera
600mm
t- 277 -

Iron plate
d

3

2

1

Induction coil

I

High frequency
signal generater f=440kHz

Fig.1. Experimental setup.

3. Results and discussion
Typical thermal images of the iron plate during
heating is shown in Fig.2. Stressed temperature range
in Fig.2 is between 15 ℃ and 100 ℃. Operating
conditions and geometries employed are as follows:
the distance between the iron plate and the coil: d=37mm, the thickness of the iron plate: t=0.7 mm, thediameter of the iron plate: 120 mm, the imposedfrequency: f=440 kHz, and the alternate current: I=2 A.
From this figure, it is found that the iron plate washeated from the circular edge. Also transient change intemperature of the iron plate is shown in Fig.3. The
approximated heating on different positions curves
were obtained from plots in this figure. Local heatbalance of the electromagnetic oscillating plate isexpressed as the following equation:CM

dT= QV − hA(T − Ta )
dθ

120

60 mm

t = 0.7 mm
d = 37 mm

Temperature (℃)

100
3

80

2

(1)

1

30 mm

60
40

2

T = -0.0039θ + 1.0462θ + 31.744
2
T = -0.0033θ + 0.9594θ + 26.753
2
T = -0.0021θ + 0.7423θ + 31.987

20
0
0

20

40

60

Time (s)80100Table.1. Heating rates of the iron plate
position

12

3

Q (W/m3)

2.74×106

3.54×106

3.86×106

The specific heat and the density of the iron plate,which were employed in this calculation, were given
as 486.6 J/kg･K and 7870 kg/m3, respectively. Table.1
and Fig.2 indicated that the heating rate and thus thealternating magnetic field were both increased withincreased radial distance. This fact implies theexistence of insufficient profile of the magnetic flux
density under the iron plate to be driven. However, thedifference of the local heating rates between the radial
positions was not so large. Induced current isgenerated by an alternating magnetic field that isparallel to the radial direction of the iron plate. From
the Fleming’s law, Lorentz force that works on theiron plate is proportional to the induced current. So, inorder to increase the magnetic field that is parallel tothe radial direction of the iron plate, it is importantthat the magnetic field is concentrated around thecenter of the iron plate. In this experiment, theair-cored induction coil was used. However, when thewater-cooled iron pole is set in the induction coil inorder to concentrate the magnetic field around thecenter of the iron plate, it is expected that the ironplate could be vibrated more strongly. Detailed designof the new system to generate a more strong magnetic
field and its effect on the electromagnetic vibration of
the iron plate will be also described in the conference
venue.

Fig.3. Transient change in temperature of the iron plate.

4. Conclusion

0s

42 s

83 s

119 s

An ultrasonic oscillator, which electromagneticallygenerates an ultrasound of the frequency of 440 kHz
by vibration of a ferromagnetic plate in the alternating
magnetic field, has been newly designed. The ironplate was vibrated by the strong electromagnetic field.
The Joule heating of the iron plate was observed by
the infrared camera to connect the local heating rate of
the plate with the local Lorentz force that causes
electromagnetic vibration. The heating rate of the ironplate and the Lorentz force that works on the ironplate were not so different between the radial positions
on the iron plate. The iron plate can be vibrated morestrongly when the magnetic flux is concentrated byusing the water-cooled iron pole in the induction coil.

Fig.2. Thermal image of the iron plate.

where C is the specific heat, M is the mass, T is the
local temperature, θ is the time, Q is the local
heating rate, h is heat transfer coefficient, A is thesurface area, V is volume, and Ta is the initial andambient temperature. Since (T-Ta) is zero at θ=0, thelocal heating rate Q can be estimated by substituting
the temperature gradient at θ =0 in Eq. (1) for
corresponding observed one given by Fig.3. The local
heating rates Q thus estimated are shown in Table.1.

References[1] M. Kuwabara et al.: International session, “Recent
Innovation in Material Processes”, CAMP-ISIJ,
16(2004).
[2] M. Kuwabara, et al.: Proc.8th Japan- China Sympo,- 278 -

Science and Technology of Iron and Steel, Chiba,
Japan, 1998, p.175.

P2-34

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 279-280
16-18 November, 2005

Acoustic Cavitation Based Production of Foamed Metallic Material◎Yasuo Saiki, Takashi Kubo, Mamoru Kuwabara and Jian Yang
(Dept. of Materials, Physics and Energy Eng. Nagoya Univ.)

Abstract
An experiment has been carried out to solidify the
Wood’s alloy in an ultrasonic field. The ultrasonicwave employed in this experiment is of output powerof 600 W and of frequency of 60 kHz, respectively.
Microstructures and micro-defects in the solidified
samples are observed by the scanning electron
microscope (SEM). It is found that many
microbubbles and microjets acoustically induced
during solidification of the metal can result in
microvoids in the solidified structure.

10mm
Converter
BoosterWave-guide

Cooling water

Over flow

Cavity due towave-guideAreafor
SEM image

Wood’s alloy
Test tube

Solidified
Wood’s alloy

Water

Fig. 1 Experimental setup for
Fig. 2 Solidified alloy andultrasonic solidification.
area for SEM image.

1. Introduction
Foamed materials have several interesting features.
There are several methods of production of the
foamed materials1). However in the conventional
methods, additives are needed or gas is generated in
the foamed materials such as hydrogen, nitrogen andso on. It is difficult to spread pores uniformly infoamed materials, and mechanical properties of afoamed material whose pores spread non-uniformly
must be inferior to that of original material. Acoustic
cavitation should become more attractive also in thisfield if cavitation malutibubble during solidification ofa metal can be trapped in its solidified structure.
Foamed materials with microstructures andmicrovoids, which are acoustically generated, canhave innovative thermal and mechanical properties.
In this study, an experiment was carried out tosolidify the Wood’s alloy in an ultrasonic field to findmicrovoids in the solidified structure.

3. Result and discussion

2. Experimental setup
The experimental setup is shown in Fig.1. The
molten wood's alloy of 80 grams was charged into thetest tube made of polypropylene, and an ultrasonicwave-guide through which ultrasound with frequencyof 20 kHz and an U.S.P. of 600 W could be imposed,
was set into the melt by the depth of 15 mm. Initialsuperheat of the melt above the melting point ofwood’s alloy (344～349 K) was about 5 K.
During ultrasonic irradiation, cooling water at 303K
started to be flown around the test tube. Even after
irradiation of ultrasound for 90 s, the flow of coolingwater was continued till the whole solidification of themolten Wood's alloy was finished. And the area ofsquare in Fig.2 was observed by SEM.

Fig.3 SEM images of the microscopic structure solidified
without ultrasound.

Fig.3 shows SEM images of the solidification
structure of Wood’s alloy solidified without ultrasonicirradiation. In order to keep the heat transfer
conditions as same as with ultrasound irradiation(Fig.4), the ultrasonic waveguide was only set into themelt and the solidification experiment was carried out.
Fig.3-(b) shows a magnified image of Fig.3-(a). The
eutectic structures were observed in Fig.3-(b). This
eutectic structure is consisted of four kinds of phases.
These phases solidified in order according to therespective melting points2).
(b)

(a)

100μm

10μm

Fig.4 shows typical SEM images of the microscopic
structures solidified with ultrasonic irradiation. The
elapsed time of ultrasound irradiation was 10 s and
U.S.P. was 600 W. As shown in Fig.4-(a), some large
spherical spaces with dark color, whose diameters
were up to 3 mm, were found. It is interesting to notethat several many ting defects, which in closed whiteframes in Fig.4-(a).These large dark colored sphericalspaces should remain to be solid-liquid coexisting
phase until the completion of solidification. The
temperature in the large spherical spaces can be
deduced to be kept homogeneously at the melting
point by ultrasonic violet mixing. After nuclei of the- 279 -

solidification generated at the melting point, the nuclei
gradually grew as small spherical particles of several
tens μm in diameter and the fractional space ofliquid gradually reduce. Finally they gather to become
the large dark colored spherical space with many
spherical particles inside.
Fig.4-(b) shows a magnified image of the whiteframed area (b) in Fig.4-(a), Fig.4-(c) also shows amagnified image of framed area (c) in Fig.4-(a).
Fig.4-(b) shows the solidification structure which
remained as the solid-liquid coexisting phase andsolidified finally. The structure is composed of manysmall particles with a eutectic structure. The
intraparticle pore is occupied by more dark substancein which alloying elements might be concentrated
during growth of the nuclei in the substance at liquidstate. Fig.4-(c) shows the microscopic structure, which
was solidified earlier than the large dark coloredspherical space. The ting defects were observed
everywhere as are shown in Fig.4-(b) and Fig.4-(c).

The mean size of the defect was about 50μm. These
defects were exceedingly small compared with anormal defection of solidification and had a possible
dimension of acoustic cavitation. Fig.5 shows theSEM image of one indentation on aluminum foil
surface attacked in water bath by an acoustic
cavitation. It is interesting to note the shape and thedimension of a defect in Fig.4 is similar to this trace
of indetation.
It is deduced that these defects are generated byacoustic cavitation and microjet. Although vapor
pressure of the solid-liquid coexisting phase is very
low and nearly zero, experimental findings in thisstudy suggest that acoustic cavitation is generated in
the liquid phase of wood’s alloy during solidification.
So it is expected that acoustic cavitation multibubble
can be trapped in the solidifying structure and acoustic
cavitation can be utilized for the production of the
foamed materials.

Waveguide

(a)

c
b

50μm

Fig.5. SEM image of one indentation on aluminum foil
surface attacked by acoustic cavitation.

4. Conclusions

1mm

A laboratory experiments on the ultrasonicsolidification was carried out using the Wood’s alloy.
Frequency and power of the ultrasound employed are
20 kHz and 600 W, respectively. Microstructures andmicro-defects in the solidified samples were
investigated by the SEM imaging and the optical
microscope observation. The solidified structuresindicated that wide area of the melt was uniformly
cooled down until the temperature of the melt was
kept just above its melting point because of intensified
mixing due to acoustic cavitation. Finally a great
many nuclei uniformly evolved and they rapidly grewin the remained melt to result in microstructures inside
the area. Even during the final stage of solidification,
the acoustic cavitation continued in the remaining
liquid accompanied by many microbubbles andmicrojets to cause microvoids in the solidified
structure.

(b)

50μm

(c)

References
1) H. Nkajima: Kinzoku,Vol.74, No.7, 2004, p.646
2) H. Kamioka: Jishin,Vol.2, No.30, 1977, p.463
50μmFig. 4. SEM images of microscopic solidified with ultrasonicirradiation.

- 280 -

P2-35

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 281-282
16-18 November, 2005

Post-Beamforming Second-Order Volterra Filters for Contrast Agent Imaging:
A Frequency-Domain Aspect
Pornchai Phukpattaranont and Kanadit Chetpattananondh (Department of Electrical Eng., Prince of Songkla Univ.)

1 Introduction
Recently, many reports of the improved diagnostic
capabilities exploiting ultrasound contrast agents (UCAs) inclinical applications have been published. This achievement
exploits the nonlinear behavior from interaction between
acoustic energy and UCAs to improve spatial and contrast
resolution. In [1], we have developed post-beamforming
nonlinear filtering algorithm based on the second-order
Volterra filter (SVF) [2] to separate linear and quadraticcomponents from UCA echoes. The proposed imaging
method, referred to as quadratic B-mode (QB) imaging,allows for transmitting pulses optimized for maximum
resolution (as in conventional B-mode) while still providing
images with significant improvement of contrast between
UCA and surrounding tissue. In this paper, we investigate
the characteristics of QB-mode imaging in the frequency
domain.

2 Theory
The QB-mode image is produced using a quadratic filter, in which the relation between input x(n) and output y(n)can be expressed asy(n) =

N
−1 N
−1
X
X

h2 (k1 , k2 )x(n − k1 )x(n − k2 ),(1)

k1 =0 k2 =0

where h2 (·, ·) represents the quadratic kernel and N is thesystem memory. The quadratic kernel for contrast agent
imaging is obtained by forming a system of linear equations
from a segment of beamformed RF data and solving for filter coefficients to minimize the error from the linear plus
quadratic predictor (For complete details of the algorithm,
please see [3]). The QB-mode image is produced by applying the optimal kernel to the beamformed RF data throughout
the B-mode image.
In this paper, we analyze the characteristics of the optimal quadratic kernel, which is capable of enhancing UCA
over tissue, in the frequency domain. The input-output relation of the quadratic filter shown in Equation 1 in the frequency domain is given by
Z Z
1 2π 2πjω
H2 (ejω1 , ejω2 )X(ejω1 )X(ejω2 )dω1 dω2
Y (e ) =2π 0 0(2)
under the assumption
ω1 + ω2 = ω,(3)where Y (ejω ) is the discrete-time Fourier transform (DTFT)
of y(n), X(ejω ) is the DTFT of x(n), and H2 (ejω1 , ejω2 )is the 2D DTFT of h2 (·, ·). Note that the observed
quadratic signal component at ω is the result of the sumof mixing all frequency components ω1 and ω2 such that
ω1 + ω2 = ω weighted by the quadratic frequency responseat ω1 + ω2 = ω.

3 Results and Discussion
We use the RF data acquired in vivo to investigate characteristics of the optimal quadratic kernel in the frequency
domain (Please see [3] for details of the experimental setup
and quadratic kernel derivation). Figure 1(a) shows the Bmode image of the kidney after the injection of 0.01 mL/kg
UCAs acquired using 3-cycle 1.56-MHz pulse (MI = 0.158)
transmissions. Average spectra of tissue and UCA signals
in the left and right boxes of Figure 1(a) are shown in Figure 1(b). We can clearly see that harmonic spectra of UCA
echoes (solid) between 2 MHz and 4 MHz frequency band
are broader than those from tissue echoes (dotted line). Figure 1(c) shows the optimal quadratic kernel used to produce
QB-mode images. The kernel is a square matrix with size 57
by 57. The resulting QB-mode image processed with thequadratic kernel shown in Figure 1(c) is shown in Figure1(d). One can clearly see the improvement of image quality. This significant enhancement of QB-mode image can be
explained using the relation described in Equation 2.
Figure 2 shows a filled contour plot of magnitudeof frequency responses in the (ω1 , ω2 ) plane. The filled
contour plot displays isolines calculated from the magnitude of frequency function in the (ω1 , ω2 ) plane andfills the areas between the isolines using constant colors.
The magnitude frequency responses obtained using RF
data in the tissue (|XT S (ejω1 )XT S (ejω2 )|) and UCA(|XCT (ejω1 )XCT (ejω2 )|) regions are shown in Figure2(a) and (b), respectively. Figure 2(c) shows the ratio ofjω1)XCT (ejω2 )|CT (econtrast to tissue magnitude, i.e. |X.
|XT S (ejω1 )XT S (ejω2 )|Four distinct peaks in the 2D plane of contrast to tissue
magnitude are observed at frequencies corresponding tosecond harmonics due to UCA, i.e. 3 MHz. Figure 2(d)shows the magnitude of 2D frequency response of the
optimal quadratic filter (|H2 (ejω1 , ejω2 )|). The magnitudeof frequency responses of quadratic components fromtissue (|H2 (ejω1 , ejω2 )||XT S (ejω1 )XT S (ejω2 )|) and UCA(|H2 (ejω1 , ejω2 )||XCT (ejω1 )XCT (ejω2 )|) regions in the(ω1 , ω2 ) plane are shown in Figure 2(e) and (f), respectively.
We can clearly see that the quadratic kernel appropriately
amplifies the region where UCA is higher than tissue atfrequency (3,3) and (-3,-3) MHz.
Figure 3(a) shows a more quantitative insight of thequadratic filter in the frequency domain. Line graphs that- 281 -

0

0

−30

−30

−20

−50

−30

−60

−40

2000

50

100
A−line #

−20

−40

−70

1500
−50

−40

0

−40

Magnitude (dB)

1000

UCAs
Tissue20−20

Magnitude (dB)

Amplitude (dB)

−20

40
UCAs
TissueKernel−10

UCAs
Tissue

−10

−10

500
Sample

0

−60

−80
−90

−60

150

−80

0

2

(a)

4
6
Frequency (MHz)

8

10

−100

(b)

−110
−5

0

5

f1 where f1+f2 = 6 MHz

−100

10
(a)

0

2

4

6

8

10
(b)

Frequency (MHz)

0
10

0.02

20

0.01

30

0

40

−0.01

50

−0.02

Figure 3: (a) Solid: |XCT (ejω1 ) XCT (ejω2 ) |, Dotted:|XT S (ejω1 ) XT S (ejω2 ) |, and Thin: |H2 (ejω1 , ejω2 )| wheref1 + f2 = 6 MHz. (b) Solid: |YCT (ejω) |, Dotted: |YT S (ejω) |.

−10−20
Sample

n2

500

1000

−30
−40

1500
−50

10

20

30
n1

40

2000

50

50

(c)

100
A−line #

−60

150

(d)

Figure 1: (a) B-mode image of the kidney. (b) Average spectra of RF data. Solid: |XCT (ejω )|, Dotted: |XT S (ejω )|. (c)The optimal quadratic kernel. (d) QB-mode image.

f2 (MHz)

10

−10

5
0

−20

0
−30

−5

f2 (MHz)

−10

5

−20

−5
0

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−5

(a)
24

10

5

23

5

0

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−5

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−5

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20
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5

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5

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−75

0

−5

−80

−5

0
f1 (MHz)

5

10

−85

5

10

−40
(b)

4 Conclusions(d)

We analyze frequency-domain characteristics of thequadratic kernel that is able to separate quadratic components from UCA signals. QB-mode images produced using
those quadratic components show significant improvement
in both contrast and spatial resolution as shown in Section 3. The understanding of the quadratic kernel in thefrequency domain allows for the improved filter design interm of both kernel size and contrast enhancement capability.

−20
−30

−5
(e)

0
f1

−10

−5

(c)

10

−5

−30

−5

10

−5

f2 (MHz)1000f1 (MHz)5510

corresponding to magnitude frequency where f1 + f2 = 6
MHz in the (ω1 , ω2 ) plane of the UCA, tissue, and kernel
are shown using, thick, dotted, and thin lines, respectively.
One can see that the kernel has a high gain where UCA
is higher than tissue. Figure 3(b) show average spectra ofdata from the QB-mode image in UCA (|YCT (ejω) |) andtissue (|YT S (ejω) |) regions using solid and dotted lines,
respectively. The intersection points between the dashedvertical line and other two horizontally spectral lines arefrom the integration along the diagonal line f1 + f2 = 6 MHz,R 2π R 2π
i.e. 0 0 H2 (ejω1 , ejω2 )XCT (ejω1 )XCT (ejω2 )dω1 dω2
R 2π R 2π
and 0 0 H2 (ejω1 , ejω2 )XT S (ejω1 )XT S (ejω2 )dω1 dω2
where f1 + f2 = 6 MHz.

−60

Acknowledgments−70

The authors thank ESAOTE S.p.A, Genoa, Italy forsupporting the in vivo data used in this paper.

−80

References10(f)Figure 2:
Filled contour plots of magnitude frequency responses in the (ω1 , ω2 ) plane:(a)|XT S (ejω1 )XT S (ejω2 )|,(b) |XCT (ejω1 )XCT (ejω2 )|,|XCT (ejω1 )XCT (ejω2 )|,(d)|H2 (ejω1 , ejω2 )|,(e)|XT S (ejω1 )XT S (ejω2 )|jω1 jω2jω1|H2 (e , e )||XT S (e )XT S (ejω2 )|,and(f)|H2 (ejω1 , ejω2 )||XCT (ejω1 )XCT (ejω2 )|.
In addition,
the diagonal line f1 + f2 = 6 MHz is overlaid in the (ω1 , ω2 )plane.

[1] P. Phukpattaranont and E. S. Ebbini, “Post-beamforming
second-order volterra filter for pulse-echo ultrasonicimaging,” IEEE Trans. Ultrason., Ferroelect., Freq.
Contr., vol. 50, no. 8, pp. 987–1001, Aug. 2003.
[2] G. L. Sicuranza, “Quadratic filters for signal processing,” Proc. of the IEEE, vol. 80, no. 8, pp. 1263–1285,
Aug. 1992.
[3] P. Phukpattaranont and E. S. Ebbini, “Post-beamforming
volterra filters for contrast-assisted ultrasonic imaging:in-vivo results,” in Proc. IEEE Ultrason. Symp., 2003,
vol. 2, pp. 1191–1194.

- 282 -

P2-36

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 283-284
16-18 November, 2005

Development of a novel ultrasonic bonedensitometry using acoustic parametersof cancellous bone for fast and slow waves○Isao MANO*, Kaoru HORII (OYO Electric co., ltd.), Shinro TAKAI (Teikyo Univ.,
Faculty of Medicine), Takuji SUZAKI, Hiroki NAGAOKA (Horiba, Ltd.) and Takahiko
OTANI (Doshisha Univ., Faculty of Engineering)
1. Introduction
An X-ray method is now widely used for noninvasive
measurement of bone mineral density for the assessmentof osteoporosis. An ultrasonic method has been proposed
as an alternative for the diagnosis of osteoporosis.
Ultrasonic measurements of bone status or bone massdensity are generally performed using the following
ultrasonic parameters, the slope of frequency dependent
attenuation (or broadband ultrasonic attenuation, BUA[dB/MHz]) and the speed of sound (SOS [m/s]). The
ultrasonic method has several advantages over X-ray bonedensitometry as it is radiation free, inexpensive andportable and has the potential to evaluate the elastic ormechanical properties of bone which the X-ray method
does not. However, a problem is inherent in the ultrasonicmethod regarding the reproducibility and the uncertainty
in measured ultrasonic parameters. A novel ultrasonicbone densitometry have been developed to overcome theploblem inherent in the ultrasonic method, and to obtain
both the bone density in the real unit [mg/cm3] or [kg/m3](not in the form of ultrasonic parameters) and the elasticity
with a comparable space resolution to the pQCT system.
2. Measurement system and method
The authors have previously reported a peculiar
propagation phenomena of ultrasonic wave in cancellous
bone (two distinct longitudinal waves, fast and slow
waves, propagating through cancellous bone),1 〜3) andtheoretically formulated the propagated ultrasonic signals
in measurement site for the fast and slow waves.4) This
formulation is applied to the developed system to obtainreal quantitative bone density and elasticity of cancellous
bone in vivo clinical measurements. The peripheral
quantitative computed tomography (pQCT, a micro focus
X-ray system) is characterized, above all X-ray methodsof bone densitometry, by measurements of a 2
dimensional local bone mass density and its imaging at
4% distal radius site. The same measurement site as pQCTis selected to compare the characteristics between the
newly developed ultrasonic system and the pQCT system.
The site to be measured, or a left distal radius (a left-handwrist) is set between a pair of broadband-focusedtransducers (transmitter and receiver) through water-filled
bags. The transmitter is driven by a single sinusoidal pulse
voltage at a frequency of 1MHz. The beam width at thefocus is approximately 3mm for a single sinusoidal pulse.
Both transducers are coaxially and confocally aligned andmove simultaneously for scanning. The ultrasound beamscans mechanically in a raster pattern through themeasurement site. The ultrasonic measurement consists of
two scanning processes, the first and the second scannings.
In the first scannings, transmitted ultrasonic signals aretaken and recorded every 2mm over a scanning area inboth X-Y directions 28×28mm. Overall amplitude of thetransmitted signals including both the fast and slow waves
are analyzed to obtain a local attenuation distribution of
the measurement site. The psuedo-bone density image
obtained by the first scanning is used for confirming thebone geometry of measurement site and for determining
the second scanning position to measure the bone density
and the bone elasticity. The second scanning position is
automatically selected by a specially developed algorithm
at almost the same measurement site of pQCT in the distal
radius and at a good likelihood site of the fast wave incancellous bone. The second scanning area is 4×4mm.
During the second scanning, transmitted signals arerecorded every 1mm and analyzed in transmission mode
to obtain the amplitudes and the propagation times for theboth fast and slow waves. Echo signals are analyzed inecho mode of the second scanning to obtain the
thicknesses of soft tissue, cortical bone and cancellous of
the measurement site.
3. Evaluation of bone density and elasticity
The ultrasonic wave is separated into two longitudinal
waves, the fast and slow waves, during the propagation incancellous bone at distal radius. Assuming that thereceived signal voltage of the receiver is E2 for the fast
wave and E2’ for the slow wave, and that the received
signal voltage without the subject (the propagation
through water) is E0. The ultrasonic parameters for thecancellous bone density are expressed for the fast wave as
T34T45 = B0 / ATe −α 4 x4 ( E 2 / E0 )(1)
and for the slow wave as* E-mail address: imano@oyoe.jp- 283 -[]

e −α 4 ' x 4 T34 'T45 ' = [B0 / AT ]( E2 ' / E0 )Slow wave level E2'/E0[dB/mm]

0.0
-1.0
-2.0
-3.0
0
100
200
300
400
Bone mass density [mg/cm3] measured by pQCTSlow wave level E2'/E0 normalized byradius thickness and bone mass density
measured by pQCT.
400

slow wave level E2'/E0

Bone mass density[mg/cm3] deduced from the

Fig.1

300
200
100
0
0
100
200
300
400
Bone mass density [mg/cm3] measured by pQCTBone density deduced from the slow wave
level and bone mass density measured by
pQCT.
3

Propagation speed of thefast wave [mm/µs] [km/s]

Fig.2

2
1
0
0
100
200
300
400
Bone mass density [mg/cm3] measured by pQCTElasticity [MPa] estimatedby fast wave

Fig.3

Propagation speed of the fast wave incancellous bone of radius and bone massdensity measured by pQCT.
8000

4. Measured results
Some examples of measured results obtained by thedeveloped system LD100 are shown from Fig.1 to Fig.4.
The slow wave level normalized by cancellous bone
thickness and the bone density deduced from eq.(2) are
shown in Fig.1 and Fig.2. The propagation speed of thefast wave and the elasticity for the fast wave are shown in
Fig.3 and Fig.4. All measured values are shown as afunction of the bone mass density measured by pQCT
system (Stratec XCT-960).
5. Remarks and conclusions
The bone density and the elasticity of in vivo cancellous
bone are evaluated by the developed ultrasonic bonedensitometry, prototype LD100.
1) Measured values of the developed system LD100 are
given with the real unit of density [mg/cm3]or [kg/m3] andof elasticity [MPa].
2) Measured densities obtained by the developed
ultrasonic system agree well with that obtained by pQCT
system and the correlation coefficient between both
systems attains 0.9.
3) A possibility of 2 dimensional imaging of the bonedensity, comparable to the pQCT system is confirmed bymeasured results.

6000

Acknowledgments

4000

This study is supported by the Japan Science andTechnology Agency as a development project of newtechnologies.

2000
0
0
100
200
300
400
Bone mass density [mg/cm3] measured by pQCT

Fig.4

(2)
where A is the total attenuation except cancellous bone, T
is the product of the transmission coefficients of all
medium boundaries except cancellous bone, B0 is theattenuation in water (without the measurement site), T34
T45 and T34’ T45’ are the products of transmission
coefficients of both boundaries of cancellous bone for the
fast and slow waves.
The dependence of the parameters for the fast and slow
waves, eqs.(1) and (2) on the bone density should be
experimentally established a priori. Then the bone densityof cancellous bone is evaluated using eqs.(1) and (2). The
propagation speed of the fast wave in cancellous bone isgiven byc4 = x4 / t4(3)x 4 and t4 are obtained by measured signals during thesecond scanning in both transmission and echo modes.
The elasticity of cancellous bone is evaluated using c4.

Elasticity corresponding to the fast wave
and bone mass density measured bypQCT.

References
1. A. Hosokawa and T. Otani : J. Acoust. Soc. Am. 101 (1997) 558.
2. A. Hosokawa, T. Otani, T. Suzaki, Y. Kubo and S. Takai,: Jpn. J. Appl.
Phys. 36, 5B (1997) 3233.
3. A. Hosokawa and T. Otani : J. Acoust. Soc. Am. 103 (1998) 2718.
4. T.Otani,: Jpn. J. Appl. Phys. 44, 6B (2005) 4578.

- 284 -

P2-37

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 285-286
16-18 November, 2005

Ultrasonic Beam Steering for Accurate Measurementof Intima-Media Thickness at Carotid Sinus○ Takashi Mashiyama, Hideyuki Hasegawa and Hiroshi Kanai(Dept. of Electronic Eng., Tohoku University)
1. Introduction
Recently, the number of patients who died due tocardiovascular diseases are in the second place following the malignant neoplasm. Since atherosclerosis is one of main causes of these diseases, it is necessary to diagnose atherosclerosis in an early stage.
The carotid sinus is a site which is easily affected byatherosclerosis [1]. However, it is impossible to diagnose atherosclerosis accurately using the conventional ultrasonography because the carotid sinus is
not flat along the axis of the vessel, and ultrasonicbeams of linear scanning are orthogonal to the arterial wall in a limited region. The reflection intensity
in a non-orthogonal region is very weak and the arterial wall in such a region is hardly recognized. In
this paper, the rough position of the arterial wall isdetermined from the B-mode image obtained by conventional linear scanning, then ultrasonic beams aretransmitted again so that all beams are orthogonal tothe arterial wall.

∆x
x

probe
skin

xi−1 xi xi+1
bi
θi
yi−1
yi
yi+1

y

ai

arterial wall
(posterior)

Figure 1: Determination of optimum beam position
and angle (N = 1).

(a)steeredbeam No:

29 ... 17 16 15 14 13 12 11 10

9

8

7

6

5

4

3

2

1

93.8
90.0
86.2
82.5
78.9
75.2
71.7
68.4
65.2
114.8 ... 91.9
88.1
84.3
80.6
77.0
73.4
70.0
66.7

θ [deg]:

θi

x θ
y

0.4 mmdesigned beam(θ i , bi )...

θi = tan−1 ai [rad],yiyi= xi −[mm],
bi = xi −tan θiai(1)(2)

where θi is the angle of beam which is perpendicular
to the position, (xi , yi ), of the arterial wall. From thedetermined beam angle, θi , and the position, (xi , yi ),of the arterial wall, the transmitting position, bi , ofthe ultrasonic beam, which perpendicularly passes
through the position, (xi , yi ), is determined. By determining θi and bi for all positions, {(xi , yi )}, alle-mail: masiyama@us.ecei.tohoku.ac.jp(x i , yi )(b)arterial walltransmit position #: 1
distance [mm]: 0( 0.2)+

2. Principle
Figure 1 shows schematic diagram of beam steering. The ultrasonic beam was scanned at M positions along the x-axis by conventional linear scanning. The depth of the arterial wall, yi (i = 1, 2, ..., M ),at each beam position, xi (i = 1, 2, ..., M ), is detected from the B-mode image obtained by conventional linear scanning. The interval of ultrasonic beams,∆x = xi+1 − xi , is 0.2 mm. With respect to a position, (xi , yi ), of i-th ultrasonic beam, a regionalslope, ai , of the arterial wall to the surface of the ultrasonic probe is estimated by the least-squares method
using the positions of the arterial wall, {(xi , yi )}, atneighboring ±N ultrasonic beams, as follows:... j−2 j−1 j j+1 j+2 ... M... j−3 j−2 j−1 j j+1 ... M−1
x

θi’

bi

y
designed beam(x i , yi )Figure 2: (a) Determination of the discrete beam angle (red line: designed beam). (b) Determination of
the discrete transmit position.
beams are designed so as to be orthogonal to the arterial wall.
In the actual diagnostic equipment, the beam angle, θi , has a discrete value. Therefore, a discrete
beam angle, θi , that is the nearest to the angle, θi , isselected from pre-assigned angles of 29 beams (Fig.
2(a)). The transmit position, bi , is calculated again
using the discrete beam angle, θi . The transmit position, bi , is also a discrete value which depends onthe element pitch of the ultrasonic probe. As shownin Fig. 2(b), the discrete transmit position, bi , whichis the nearest to bi , is selected as well as the angle,
θi . In this way, a beam transmitted to the position,
where is the nearest to (xi , yi ), is selected.

- 285 -

(a)xultrasounddiagnostic
equipment[1] C. K. Zarins, D. P. Giddens, B. K. Bharadvaj, V. S.
Sottiurai, R. F. Mabon and S. Glagov: Circ. Res. 53(1983) 502.

silicone rubber(fixed)XYZ stage

water bath(b)
3 mm/s

probe
x
y
...

O ...

silicone rubber

13 mm

Figure 3: (a) Experimental system. (b) Schematic
diagram of measurement.

0

distance [mm]
5
10

0
0

5

5

depth [mm]0distance [mm]
5
10

10

10

4. Conclusion
In this paper, a method for designing beam steering which makes ultrasonic beams orthogonal to thearterial wall at the carotid sinus was proposed. In basic experiments, it was shown that it was possible toimage a nonflat object in the entire scanned area using the proposed beam steering.
References

probe

y

depth [mm]3. Basic experiments using silicone rubber
As shown in Fig. 3(a), a silicone rubber tube which
was open up was fixed in a water bath and it was measured by a linear ultrasonic probe at 10 MHz of thediagnostic equipment (Aloka SSD-6500). The sizeof the silicone rubber was configured to be 13 mm
because the size of the carotid sinus in the longitudinal direction is 10-13 mm. Using the beam scanning
method [2], all ultrasonic beams are transmitted so as
to pass the point, O, and 29 beams are transmitted atvarious angles θi (65-115 deg). The ultrasonic probe
was moved in the x-axis direction at 3 mm/s for 4 s,
and the point, O, was moved as shown in Fig. 3(b).
Ultrasonic RF echos were obtained 60 times by moving the point, O, with a pitch of 0.2 mm along x-axis.
A B-mode image of conventional linear scanning(Fig. 4(a)) was composed of 60 ultrasonic beams. 60
positions of the silicone rubber surface {(xi , yi )} (i =1, 2, ..., 60) were assigned on the B-mode image (redline in Fig. 4(a)). Based on the principle in Sec. 2,
beams (angle θi , transmit position bi ) which are orthogonal to the surface of the silicone rubber were
calculated using {(xi , yi )}. The yellow line in Fig. 4(a)shows a designed optimum beam (i = 9). Designed
beam angle θ9 is 67.8 degrees and transmit position
b9 is 7.77 mm. Then, a beam with a discrete beam
angle of θ9 = 68.4 degrees and a discrete transmit
position of b9 = 7.6 mm was selected.
Figure 4(b) shows a B-mode image composed byultrasonic RF echos obtained by the designed beams.
The ultrasonic beam inclines from a vertical direction
by 22.2 degrees at the surface position (x9 , y9 ). At
this position (x9 , y9 ), the silicone rubber cannot be
imaged by the conventional linear scanning as shownin Fig. 4(a). On the other hand, a strong reflectedecho was obtained from such a position (x9 , y9 ) using the proposed method, and the silicone rubber was
imaged in the entire scanned area as shown in Fig. 4(b).

15

15

20

20
(a)

(b)

Figure 4: (a) B-mode image obtained by conventional linear scanning (red line: assigned silicone
rubber surface, yellow line: designed beam (i = 9)).
(b) B-mode image obtained by the proposed steered
beams.

[2] N. Nakagawa, H. Hasegawa and H. Kanai: Jpn. J.
Appl. Phys. 43 (2004) 3220.

- 286 -

P2-38Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 287-288
16-18 November, 2005

Accurate Ultrasonic Measurement of Surface Profile
by Detecting Phase Shift During Scanning○ Chihiro Arihara, Hideyuki Hasegawa, and Hiroshi Kanai(Dept. of Electronic Eng., Tohoku University)

1. Introduction
Recently, increasing circulatory disease such as myocardial infarction and cerebral infarction becomes aserious problem. It is important to diagnose atherosclerosis, which is a main cause of diseases in an earlystage.
Luminal surface of clinically normal artery is covered with a piece of endothelial cells and will besmooth. In an early stage of atherosclerosis, luminal
surface is damaged [1], edema develops under endothelium, endothelial cell is separated, and then luminal surface becomes rough. To detect atherosclerosis in an early stage, roughness of luminal surface
needs to be detected with an accuracy of micrometer that cannot be achieved by traditional ultrasonicdiagnostic equipments.
Interval of the ultrasonic beams is about 100 µm,
and B-mode image is created using the amplitude of
the received wave. Therefore, resolution of B-mode
image of ultrasonic diagnostic equipment is 100 µm
in a lateral direction and is 150 µm, which corresponds to the wavelength at 10 MHz, in the depth
direction. In this paper, accurate detection of roughness in the depth direction was achieved using thephase of the received wave. Furthermore, the lateral
resolution, which is deteriorated by a finite beam diameter, is improved by an inverse filter.

2. Principle
2.1 Method
A linear-type ultrasonic probe of ultrasonic diagnostic equipment (Aloka SSD-6500) used in this experiment transmits and receives ultrasonic pulse byscanning the ultrasonic beam along the surface of an
object (center frequency f0 =10 MHz, scanning pitch
L=0.1 mm, and number of beams M =70 (beam position: l1 , l2 , · · · , lm , · · · , lM )). Ultrasonic RF signals
are sampled with 40 MHz. B-mode images consist of
sampled points, which are obtained every 19.25 µm
in the depth direction (sound speed c0 =1,540 m/s)
and 0.1 mm in the lateral direction (a set of two dimensional discrete data: frame).
Figure 1 shows the experimental system and sizeof an object (extension board: 0.5 mm pitch). The
object was stuck to rubber board with thumbtacks.
Water was boiled to expel any remaining air, and thewall of the tank was covered with acoustic absorbing
material.
The object was measured with a frame rate FR=188
frames/s while the probe moved automatically at theultrasonic probex

ultrasonicdiagnostic
equipment

z

water

XYZ stagex av =220 µ m
x =500 µ m
controller0zav =32 µ m
extension board

Figure 1: Experimental system for measurement of
surface of object by ultrasound.
constant speed vstg =50 µm/s for 2 seconds. The distance moved L=100 µm, which corresponds to thescanning interval of ultrasonic beam, in x-direction
with an XYZ stage. Let us define the measurement
location of beam l1 of the first frame by x(1, 1)=0mm, the measurement location of beam lm of nthframe is obtained as follows:(n − 1) · vstg
(1)
FR.
In this paper, the phase difference ∆φ(x) of the received wave at each measurement location was calculated as interval of frame ∆N =19. In this case,the probe moved ∆x=5 µm during ∆N . Therefore,
in the x direction, interval of measured location was
smaller than scanning interval of L=100 µm.
The demodulated signal of received RF echo wasused to measure surface profile with an accuracy of
micrometer in the depth direction. When the phasedifference of received wave between locations, x andx + ∆x, is ∆φ(x), the surface profile, z(x), is obtained as follows:x(n, m) = (m − 1) · L +c0 ∆φ(x)4πf0 , x∆z(x )dx.
z(x) =∆z(x) =

0

(2)
(3)

At this time, the measured depth, d(x), corresponds
to the surface of the object (about 11 mm), and focal
point was about 14 mm from the surface of the probe.
It is difficult that the surface of the ultrasonic probemoves in a direction perfectly parallel to the surfaceof the object. Therefore, the angle between the surface of the ultrasonic probe and the surface of theobject needs to be compensated. The compensatedsurface profile, g(x), is obtained by subtracting themeasure line, which is calculated by the least-square
method from z(x).

- 287 -

2.2 Improvement of deteriorated lateral resolution using inverse filter
Beam diameter deteriorates the lateral resolution.
The actual surface profile, f (x), of the object is expressed by the measured surface profile, g(x), thebeam diameter (point spread function), h(x), noise
n(x), and their Fourier spectra, F (u), G(u), H(u),and N (u), as follows:g(x) = f (x) ∗ h(x) + n(x),(4)(5)

where u is spatial frequency.
In this paper, the measured signal is assumed to
be noise free. The estimate of the surface profile,f(x), is calculated by the inverse Fourier transformof F(u), after applying the inverse filter M (u), named
Wiener filter to G(u) as follows [2]:F(u) = G(u) · M (u),(6)

3. Experimental Results

In this paper, the surface profile of an object with asurface roughness in micrometer order was detectedusing the phase of the RF signal. By applying an inverse filter as point spread function to the measured
surface profile, the estimated surface profile showshigher accuracy than the measured surface profile without the filter.
References[1] R. Ross: N. Engl. J. Med., 340 (2004) 115.

1000

2000

6000

2
4
6
8
10

10
0
−10
−20

0

(c)

3000

distance x [ µ m]

8000

| F(u) |>| F(u) || G(u) || H(u) |spectrum power

12000

4000
0
0

| H(u)|
1500
1000
500

2.5

5

7.5

10

12.5

spatial frequency u [1/mm](d)depth z [ µ m]

4. Conclusion

4000

20

(7)

Figure 2(a) shows a B-mode image obtained of theobject by ultrasonic diagnostic equipment. Figure2(b) shows the measured surface of the object g(x).
Roughness, which could not be found in Fig. 2(a),could be found easily in Fig. 2(b). The step width ofg(x) was equivalent to the actual pitch of the surfaceof the object (500 µm).
Figure 2(c) shows power spectra, |F (u)|, |G(u)|,|H(u)|, and |F(u)|, of the point spread function, h(x),and surface profiles, f (x), g(x), and f(x), respectively. In this paper, H(u) is obtained by analyzing
the profile of acoustic pressure within the width (500µm) where the acoustic pressure is larger than 10%of the maximum. Figure 2(d) shows the measured
surface profile, g(x), the estimated surface profile,f(x), using the inverse filter, and the actual surface
profile f (x). The shape of f(x) was more similar tothat of f (x) in comparison with g(x).

2000

f(x)
g(x)
f(x)

20

>

|H(u)|2
1
H(u) |H(u)|2 + Pnn (u)/Pf f (u) .

distance x [ µ m](b)depth z [ µ m]

M (u) =

00depth z [mm]G(u) = F (u) · H(u) + N (u),

(a)

10
0
−10
−20
0

500

1000

distance x [ µ m]

1500

Figure 2: (a) B-mode image. (b) Surface profile,g(x). (c) Spectra of point spread function and surface
profiles , f (x), g(x), and f(x). (d) Surface profiles,
f (x), g(x), and f(x).

[2] H. C. Andrews, B. R. Hunt: Digital Image Restoration, (1977) 126.

- 288 -

P2-39Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 289-290
16-18 November, 2005

In Vitro Measurement of Ultrasonic Scattering
Dependence on Myocardial Direction○ Hiroaki Kamada, Hideyuki Hasegawa and Hiroshi Kanai(Dept. of Electronic Eng., Tohoku University)
1. Introduction
The direction of myocardial fiber of normal humanheart wall changes gradually from the epicardium tothe endocardium [1]. However, in hypertrophic heartdisease, the myocardial direction becomes disarrayed.
It is supposed that ultrasonic backscatter from heart
wall reflects such a change of myocardial direction.
Therefore, basic researches have been conducted forquantitative estimation of myocardial tissue [2]. In
order to quantify the correspondence between acoustic properties of a object and ultrasonic backscattering, Knipp et al. made use of a reference phantom ofknown acoustic properties for measuring ultrasonicbackscattering with eliminating influences of the measurement system [3].
In this paper, ultrasonic echoes from cloth fibers
mimicking myocardium and from a porcine ventricular wall were measured with changing the angle of insonification relative to the fiber direction. From thesemeasurements, ultrasonic scattering dependence on
the fiber direction was investigated.

transducers was 30 mm and the transducers were inclined from the vertical direction by 30 degrees. A
burst sine wave of seven cycle was applied to transmit transducer. The reflected signal from the object
was received by receiving transducer. The digitizeddata were acquired after averaging the reflected signals 128 times.

2. Measurement Methods
In this paper, an ultrasonic echo from an object
was measured with changing the angle of insonification relative to the fiber direction by rotating theobject. As shown in Fig. 1, the reflected signal wasmeasured at every 10 degrees by rotating the object.

Figure 2: Measurement System.

transducer
o

o

30 30
0

.
. .
. .
.

gauze
Figure 1: Schematic diagram of the measurement.
The measurement system is shown in Fig. 2. Bothtransducers for transmission and reception are single
element concave transducers. Focal distance of theE-mail: kamada@us.ecei.tohoku.ac.jp

Waveform
Generator
DigitalOscilloscopePower AMP

PC

30o30ogauze
XYZ StageXYZ Stage

3. Basic Experiment Using Cloth Fibers
The object made of cloth fibers consists of 16 layers of gauze. Unnecessary fibers were removed fromgauzes so that fibers becomes homogeneously lined
along one direction. The diameter of a human myocardial fiber is smaller than the wavelength of ultrasound used in ultrasonic diagnostic equipment. Therefore, ultrasonic echoes from a human myocardium
becomes scattered waves. To make echoes from thecloth fibers scattered waves as in the case of humanmyocardium, the diameter of the used fiber was about
150 µm, which is smaller than the wavelength of ultrasound (214 µm) used in this basic experiment. The
thickness of the whole gauze is about 9 mm.
At first, the ultrasonic echoes from the gauze were
measured with setting the focal position at the surface of the gauze. Secondly, the focal position was
set inside the gauze by moving both transducers to
depth direction by 4.5 mm. Figure 3 shows the maximum amplitude of the reflected signal from the surface and the inside of the gauze at each insonification
angle relative to fibers. In both cases, the amplitudeof the reflected signal changed due to the direction- 289 -

of fibers. In the case of the parallel insonification
relative to the direction of fibers (θ = 0 degrees and180 degrees), the amplitude of the reflected signal
became the maximum.

40
amplitude (mV)from surface

from inside

30
20
10
00

90

180
270
angle θ (deg)360

imen as well as in the case of the gauze. Figure 4
shows the maximum amplitude of the reflected signal from the specimen. The reflected signal seemed
to change randomly with respect to the insonification
angle. However, in the case of θ = 20 degrees and 350
degrees, the reflected signal was significantly large
in comparison with those at other insonification angles. It is supposed that the directions of myocardial fibers were homogeneously lined along one direction at each depth because the porcine heart measured in this experiment was excised from the normal
pig (considered not to have any heart disease). Therefore, the reflected signal from the porcine heart wallshowed the angular dependence as well as the basic
experiments using cloth fibers.

8

The insonification angle dependence of ultrasonicecho is significant in the case where an object is larger
than the wavelength of ultrasound. However, in thecase of scattering, the angle dependence is less significant. However, when scatterers which are smaller
than wavelength of ultrasound are located at a constant pitch, ultrasonic echo changes due to the angle of insonification relative to the fiber direction.
Though the gauze used in this measurement was composed of fibers which are smaller than the wavelengthof ultrasound, the fibers were located at a constant
pitch and were lined only along one direction. Therefore, the reflected signals from the gauze showed thechange due to the direction of the fiber.
4. In Vitro Experiment Using Excised Porcine
Heart
The ultrasonic echoes from an excised porcine
heart wall were measured. In the porcine heart wall,the direction of myocardium changes gradually fromthe epicardium to the endocardium as well as thehuman heart wall [1]. Therefore, the angle dependence of the echo reflected from the porcine heart
wall was investigated. After slaughter, the porcine
heart was measured within a half day. A part of theleft ventricular wall was excised from the porcine
heart and was put the epicardium side up in the water tank which was filled with the normal saline solution. The focal position was set at the surface ofthe specimen (epicardium), the reflected signal wasmeasured at every 10 degrees by rotating the spec-

amplitude (mV)

Figure 3: Changes in the maximum amplitude of thereflected signal due to the direction of the fiber. The
solid and dotted lines show the maximum amplitudesof the reflected signals from the surface and the inside of the gauze, respectively.

6
4
2
0
0

90

180
270
angle (deg)360

Figure 4: Change in the maximum amplitude of thereflected signal from the porcine ventricular wall.
5. Conclusion
In this study, in basic experiments using cloth fibers,
it showed that the amplitude of the reflected signal
changed due to the angle of insonification relative tothe fiber direction. As well as the basic experiments,
in the case of the porcine heart wall, the reflected signal showed the angle dependence. Since the directionof myocardial fibers becomes disarrayed in patientswith hypertrophic heart disease, it is possible to estimate the condition of myocardial fibers using thefeature of the change in ultrasonic scattering due tothe directions of fibers.
References- 290 -

[1] F. Rushmer: WB Saunders (1976) 76.
[2] L. Baldwin, R. Marutyan, M. Yang, D. Wallace, R.
Holland, and G. Miller: Ultrasound in Med. & Biol.
31 (2005) 477.
[3] S. Knipp, A. Zagzebski, A. Wilson, F. Dong and L.
Madsen: Ultrason Imaging 19 (1997) 221.

P2-40

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 291-292
16-18 November, 2005

Measurement of Spatial Distribution of Strain Generatedby Dual Acoustic Radiation Forces○ Mikito Takahashi, Hideyuki Hasegawa, and Hiroshi Kanai(Dept. of Electronic Eng., Tohoku University)

2. Principle
When the ultrasound propagates in the medium, aconstant force is generated in the direction of propagation. This force is called as the acoustic radiation
force. The acoustic radiation pressure is defined as
the acoustic radiation force per unit area.
When two ultrasonic beams of the amplitude ofsound pressure, p0 , at slightly different frequencies,
f and f + ∆f , are crossed each other [1, 3, 4], an
acoustic radiation pressure which fluctuates at the frequency difference, ∆f , is generated. The sound pressure, psum, on the surface of the object is expressed
as follows:on the interface is given by
PR (t) = (1 + R2 )synchronization
signal generator

f1+∆ fpsum =p0 cos(2πf t + θ) + p0 cos 2π(f + ∆f )t, (1)p2= 0 2 [1 + cos(2π∆f t − θ)ρ1 c1
+ cos 2π{(2f + ∆f )t + θ}1
+ cos 2(2πf t + θ)2
1
+ cos 4π(f + ∆f )t],2(2)where ρ1 and c1 are the density and sound speed ofthe medium, respectively. From the second term of
the right-hand side of eq. (2), it is found that the energy density, e(t), has a component at the frequency
difference ∆f . With respect to the low-frequency component, the acoustic radiation pressure, PR (t), acting

signal generator

f1

f2+∆ f f2

adder

adder

mixer

divider

dividermixer

power AMP
ultrasonictransducer
1
φ1
water

ultrasonic
probe

LPF

rectangular
A/D converter
pulsesfunction
generator

electronic controlled switch

where f and θ are the frequency of the transmitted
pulse and the phase difference between sinusoidal
waves at f and f +∆f , respectively. The energy density, e(t), is expressed as follows:
1e(t) ={psum (t)}2
ρ1 c21
1={p0 cos(2πf t + θ)
ρ1 c21+p0 cos 2π(f + ∆f )t}2p20
[1 + cos(2π∆f t − θ)], (3)ρ1 c21

where R is the pressure reflection coefficient on thesurface of the object. When multiple acoustic radiation pressures, PR (t), are employed, the phase of anradiation pressure relative to the other radiation pressures can be controlled by changing θ in eq. (3).
3. Experiments
The measurement system is illustrated in Fig. 1[4]. Insonification angles, φ1 and φ2 , of two ultrasound beams for actuation were set at 25 and 155 degrees, respectively. The object made of gel (45 mm×45
mm×17 mm, containing carbon powder to obtain areflection from the inside of gel) was used.

17mm

1. Introduction
In recent years, some remote actuation methods
based on acoustic radiation forces have been reported.
Fatemi and Greenleaf proposed an imaging modalitythat uses the acoustic response of the object to a localized dynamic radiation force [1].
In this paper, the spatial distribution of strain generated by dual acoustic radiation forces was measured by the ultrasonic phased tracking method [2].

power AMP

real

TRIG

imag

diagnostic
ultrasoundsystem
2

GPIB

φ2

workstation

focal points
gel
block
water tank

Figure 1: Measurement system.
The spatial distribution of strain generated by phasecontrolled dual acoustic radiation forces (θ = 0) is
illustrated in Fig. 2. Two ultrasonic transducers for
actuation (center frequency: 1 MHz, focal point: 52
mm from the surface of the transducer) were driven
by a sum of two continuous waves at two slightly
different frequencies of 1 MHz and 1 M+∆f Hz.
Phase-controlled dual acoustic radiation forces (θ =0) were applied by setting the focal points at two
different positions in the object. These two applied
radiation forces cyclically compress the region between two focal points at the frequency difference,∆f , in the horizontal direction. Therefore, this region is cyclically thickened in the vertical direction.

- 291 -

The displacement distribution is also measured withultrasound combined with the phased tracking method[2].
P R1 (t)
0

t1

time

t1

time

P R2 (t)0acoustic radiation acoustic radiation
force (P R2 (t) )force (P R1 (t) )strain(period A in Fig. 4), points at depths of 0 mm and
4.0 mm move upward and downward, respectively.
This phenomenon shows that the thickness along theultrasonic beam at position e increases by applicationof radiation pressures.
Figure 5 shows the spatial distribution of the maximum displacements. Lengths and directions up-arrows
and down-arrows show the magnitudes and directionsof the measured displacements.

objectobjectt= 0

t= t 1

1.0 µ mfocal points

Figure 2: Generation of strain.
Positions where displacements were measured areillustrated in Fig. 3. The ultrasonic probe for measurement was moved with a pitch of 0.3 mm in the
direction of the x axis and the displacement distribution was measured at nine positions (a, b, c, ・・・, i).
A position e is the center between two focal points.
The frequency difference, ∆f , was set at 10 Hz.
xultrasonic probe
for measurementy∆ d=0.3 mm
a

b

c

d

e

f

g

h

i

17 mm

2.4 mmsilicone gel
45 mm

Figure 3: Illustration of positions where displacements were measured.
A+2.0 µ m
0 µm-2.0 µ m

b

c

d

e

f

g

h i

0

0

0

0

0

(1) 0 mm0(2) 2.7 mm(3) 4.0 mm

0

0 0

0

0

0

0

0

0

(4) 4.7 mm

0.3 mmFigure 5: Spatial distribution of the maximum displacements.
As well as Fig. 4, the displacement distribution
shows the thickness in the region between two focal
points increases by radiation forces. In Fig. 5, it isfound that displacements measured at a depth of 4.0
mm was larger than those measured at 0 mm. Ultrasonic beams for actuation were insonified along theoblique directions, and the applied radiation forces
have components acting downward.
These results show that the change in thickness(strain) inside the object was successfully generatedusing phase-controlled dual acoustic radiation forces(θ = 0).

acoustic radiation
force (P R2 (t) )acoustic radiation
force (PR1 (t) )a

4. Conclusion
In this paper, it was confirmed that the strain could
be generated by applying phase-controlled dual acoustic radiation forces. Furthermore, it was shown thatthe generated strain was measurable by the ultrasonicphased tracking method.

(a)+2.0 µ m

References

0 µm
(b)

-2.0 µ m[1] M. Fatemi and J. F. Greenleaf: IEEE Trans. on Med.
Imag. 21 (2002) 1.

+2.0 µ m
0 µm(c)-2.0 µ m[2] H. Kanai, M. Sato, Y. Koiwa, N. Chubachi: IEEE
Trans. on Ultrason. Ferroelectr. Freq. Control. 43(1996) 791.

+2.0 µ m
0 µm(d)-2.0 µ m
0s

0.5 s

time

Figure 4: Measured displacements at depths of 0 mm(a), 2.7 mm (b), 4.0 mm (c), and 4.7 mm (d).
Figure 4 shows the measured displacement at each
depth at position e. From the results shown in Fig.
4, it is found that when radiation pressure increases- 292 -

[3] K. Michishita, H. Hasegawa and H. Kanai: Jpn. J.
Appl. Phys. 42 (2003) 4608.
[4] Y. Nishio, N. Hashimoto, H. Hasegawa and H.
Kanai: IEEE Intnl. Ultrasonic Symp. (2004) 1773.

P2-41

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 293-294
16-18 November, 2005

Tissue Classification of Arterial Wall Based on Correlation
Between Regional Elasticity Distributions andElasticity Histograms of Tissues○ Jun Inagaki, Hideyuki Hasegawa, Hiroshi Kanai (Dept of Electronic Eng., Tohoku University)
Masataka Ichiki (Sendai Hospital of East Railway Company)
Fumiaki Tezuka (Sendai Medical Center)tissue) were assigned on the elasticity images. From
the regions assigned for respective tissues, the elasticity histograms were obtained. Figure 1(c) shows
the respective elasticity histograms of tissues.
elasticty image
ROI size
0.9x0.9 mm

inagaki@us.ecei.tohoku.ac.jp

3
2

lumen

1.5
adventitia

3 mm

1

regional elasticity histogram
8

(b−1)

frequency [%]frequency [%]

0.5
6
4
2
0

0

1

2

3

(b−2)

30

0
20
10
0

4

elastic modulus [MPa]correlation0

1

2

elastic modulus [MPa]4fibrous tissue3_ 0.63 MPa
1.0 +(N=5213)

2
1
0

0

1

2

3

elastic modulus [MPa]4frequency [%]frequency [%]4

calcified tissue3_ 1.0 MPa (N=218)
2.0 +

2
1
0

0

1

2

3

(c)

60
lipid
50
_ 40 kPa
81+
40(N=425)
30
20
10
00
1
2
elastic modulus [MPa]40frequency [%]frequency [%]elasiticity histogram for each tissue

2.1 Construction of Elasticity Histograms of Respective Tissues

In the experimental setup for in vitro experiments
using excised femoral arteries of patients with arteriosclerosis obliterans, the change in internal pressure was generated using an artificial heart, and theinternal pressure was measured by a pressure detector placed in lumen of the artery. The artery fixed in
the water tank was measured in the long-axis planewith a linear-type probe (center frequency: 7.5 MHz).
From the measured change in thickness and pressure,the elasticity of arterial wall was estimated at intervals of 75 µm in depth direction and 300 µm in axial
direction.
After ultrasonic elasticity measurement, the pathological images were made to identify tissue component of the arterial wall in the scanned plane. The
plane scanned by ultrasound was identified by imaging a needle fixed on the external surface of the posterior wall. By referring the pathological images, theregions which corresponding to one of the respective
tissues (lipid, blood clot, fibrous tissue and calcified

4 MPa

(a)

elastic modulus

1. Introduction
Cerebral infarction and myocardial infarction are
mainly caused as terminal symptoms of atherosclerosis by occlusion of the artery due to blood clots
which are generated by the rupture of atherosclerotic
plaque [1]. Direct characterization of the composition and vulnerability of atherosclerotic plaque may
offer insight into the mechanism of plaque regression and progression [2, 3]. The change in elasticity is caused greatly by the change in composition of
the arterial wall due to development of atherosclerosis [5]. Therefore, the measurement of elasticity has
potential for tissue characterization of plaque.
The elasticity histograms of respective tissues (lipid,
blood clot, fibrous tissue, and calcified tissue) in artery
were constructed to classify arterial tissue based onelasticity [6]. In this paper, tissue classification basedon correlation between the elasticity distribution ofeach small region in an elasticity image and elasticity histograms of respective tissues is investigated to
improve tissue classification based on elasticity.
2. Method

4

elastic modulus [MPa]56blood clot30

_ 57 kPa
96 +(N=612)

20
10
0

0

1

2

elastic modulus [MPa]Figure 1: (a) Elasticity image of arterial wall. (b)Regional elasticity distributions in elasticity image.
(c) Elasticity histograms of lipid, blood clot, fibrous
tissue, and calcified tissue.
2.2 Tissue Classification Based on Correlation Between
Elasticity Histograms

The region of interest (ROI) was assigned on theelasticity image (regions surrounded by green rectangles in Fig. 1(a)) which was obtained by ultrasonicmeasurement, and the elasticity distributions within
the ROI were obtained (Fig. 1(b)). The size of ROI
was 900 µm in the depth direction and 900 µm in thearterial axial direction. Therefore, 36 elasticity values were included in each ROI. Although the number of elasticity values within a ROI whose center located at the edge of the elasticity image becomes lessthan 36, the elasticity distribution within each ROI
was also obtained.
The ROI was scanned in the entire elasticity imageat the intervals of 75 µm in the depth direction and- 293 -

Table 1: Correlation between regional elasticity distribution and elasticity histogram of each tissue.
b-1
b-2

rl
−0.07
0.80

rb
−0.06
0.84

(a)

elasticity image

rf
0.80
−0.04

rc
0.35
−0.15
elastic
modulus

3 mm

4
3
2

(b) pathological image
3 mm

1.5

blood clot
lipid

1

0.5

fibrous tissue
0
MPa

(c−1)
lipid

(c−2)
blood clotcorrelation
coefficient

300 µm in the axial direction, and the correlation coefficients between the regional elasticity distribution
at each position and elasticity histograms of respective tissues were calculated. The correlation images
were obtained by color coding corresponding to thecorrelation coefficient of each ROI.
3. Results
Table 1 shows the correlation coefficients (rl : lipid,
rb : blood clot, rf : fibrous tissue, rc : calcified tissue)
between regional elasticity distributions (Figs. 1(b1) and 1(b-2)) and elasticity histogram of each tissue. Regional elasticity distributions of Figs. 1(b1) and 1(b-2) showed the highest correlation with fibrous tissue and blood clot, respectively.
Figures 2(a) and 2(b) show the elasticity image
and the corresponding pathological image, respectively. Lipid, blood clot, and fibrous tissue were identified by referring to the pathological image. Figure2(c) shows the correlation image for each tissue. The
color bar indicates the correlation coefficient for eachtissue.
In Figs. 2(c-1), 2(c-2) and 2(c-3), the region oflipid, blood clot and fibrous tissue was superimposedon the correlation images by yellow, red, and green
broken line, respectively. As shown in Figs. 2(c1), 2(c-2), and 2(c-3), elasticity histograms for lipid,
blood clot, and fibrous tissue give high correlation
coefficient in the regions which were identified to berespective tissues. Althogh it is difficult to classify
a region to be lipid or blood clot, it is possible todifferentiate lipid and blood clot from fibrous tissue
and calcified tissue as shown in Fig. 2(c).
4. Conclusion
In this paper, the tissue classification based on thecorrelation between the measured elasticity distribution and the pre-determined elasticity distribution ofeach tissue components was investigated. Fibrous
tissue was found to be differentiated from lipids or
blood clots. However, differentiation of lipids andblood clots seemed to be still difficult.

(c−3)
fibrous tissue

References(c−4)[1] P. R. Moreno, E. Falk, I. F. Palacios, et al: Circulation 90 (1994) 775.
[2] H. M. Loree, R. D. Kamm, R. G. Stringfellow, et al:
Circ. Res. 71 (1992) 850.
[3] E. Falk, K. Prediman, P. K. Shah, at al: Circulation
92 (1995) 657.
[4] R. T. Lee, A. J. Grodzinsky, E. H. Frank, et al: Circulation 83 (1991) 1764.
[5] J. Inagaki, H. Hasegawa, H. Kanai, et al: Jpn. J.
Appl. Phys. 44 (2005) 4593.
[6] H. Kanai, H. Hasegawa, M. Ichiki, et al: Circulation
107 (2003) 3018.

calcified tissue

1

0

− 0.5

3 mm

Figure 2: (a) Elasticity image of arterial wall. (b)Pathological image of corresponding section. (c)Correlation between regional elasticity distribution
and elasticity histogram of each tissue.

- 294 -

P2-42

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 295-296
16-18 November, 2005

Measurement of the amount of embolic usingtransesophageal echocardiography◎Tadashi YAMAGUCHI, Kazuki HIRAI, Masami AOKI, Jin MIYAGI,
Masahiko SUZUKI, Hideshige MORIYA and Hiroyuki HACHIYA (Chiba Univ.)
1. INTRODUCTION
Deep venous Thrombosisi (DVT) and Pulmonary
Thromboembolism (PTE) are serious complications
after Total Knee Replacement (TKR). The tighttourniquet usually used in TKR placed around thepatient’s leg caused blood stasis and encourage theformation of a thrombus. This thrombus may travel
through the blood stream, embolizes the pulmonary
arte and causes DVT. Therefore, understanding of theflowing thrombus volume is important to curing andprevention DVT.
In this study, we propose the measurement methodof the amount of embolus that are carried from the leg
to lung by the bloodstream.

(a) Original image

2. DATA ACQUISITION
We acquired transeophageal echocardiography(TEE) data of right atrium of heart by ultrasonicdiagnosis equipment (Hewlett Packard SONOS5500)
in the period throughout the operation end out of TKR.
The center frequency of ultrasound is 4 MHz. TEE
datasets are saved as S-VHS movies (frame rate is
30fps). Each movie are converted to the sequence of
TIFF images for data analysis and image processing.
3. EXTRACTION OF EMBOLUS
We tried to extract embolus information from TEE
image by Fiber Structure Extraction Technique(FSET)[1]. In heart chamber that contains no embolus,
the background echoes from each randomly
distributed scatter interfere mutually, and it is wellknown that the probability density function (PDF) of
echo envelope amplitude can be mostly approximatedby the Rayleigh distribution. On the other hand, theecho envelope amplitude from embolus is larger thanthat from surrounding area. Hence, the PDF will notobey with Rayleigh distribution. By constraining theRayleigh elements and accent non-Rayleigh elements,
we can extract the embolus information from TEE
images.

(b) Image of afer FSET. Non Rayleigh elements are
extracted as high echo.
Fig.1 Result of FSET for TEE image
Figure1(b) is the result of extracting the embolus
information from Fig.1(a) by FSET. The punctuatestructures seems like embolus are extracted in theright atrium of heart. FSET leverages the statisticalproperty of echo signals, so it can extract 䏈䏏䏈䏐䏈䏑䏗䏖that can't be extracted by simp䏏䏜 threshold䎃䏓䏕䏒䏆䏈䏖䏖䏌䏑䏊.
4. VOLMETRIC RATIO OF THE EMBOLUS
Figure 2 shows the time change of extracted images.
The time tourniquet became loose after KTC
operation is 0 second in Fig.2. It is shown amounts ofembolus are different in each result. Then, we
appropriated about time rate of change of the embolus.
To measure the amount of the embolus, it is necessary
to extract the endocardium from TEE images. We used
Brain Extract Tool (BET) method[2][3] , it is one of theactive contour model to endocardium extraction. Fig.3
shows the result of endocardium extraction by BET
method. The endocardium of right atrium of heart
were extracted clearly regardless of the existence ofembolus.

- 295 -

heart beat.
5. CONCLUSION
We proposed the method to measuring the amount ofembolus, and demonstrated that our method can detect
the variation of embolus flow in the right atrium ofheart.
The processing speed of each operation isimportant to achieve the real-time measurement ofembolus flow. In the actual condition, FSET
operates approximately in real time, but BET
method doesn’t. So, we examine about extraction
method of an endocardium by two kinds of theprocessing methods of BET method and the move
distance pursuing method in the next step.
REFERENCES
1. T. Yamaguchi et. al.: JJAP, Vol.44, No.6B (2005)
4615.
2. S. M. Smith: Human Brain Mapping 17 (2002) 143
3. T. Yamaguchi et. al.: IEICE Rep. US2004-103(2005) 7(a) 20 second(b) 60 second
Fig.2 Time change of the result of FSET.

䎓䎑䎔䎙
䎓䎑䎔䎗
䎓䎑䎔䎕

(a) Case A: there are no embolus in heart chamber.

䎵䏄䏗䏈 䎓䎑䎔
䎓䎑䎓䎛
䎓䎑䎓䎙
䎓䎑䎓䎗
䎓䎑䎓䎕
䎓
䎓

䎔䎓

䎕䎓

䎖䎓
䎗䎓
䎷䏌䏐䏈䎃䎋䏖䎌

䎘䎓

䎙䎓

䎚䎓

(a) long period case (100s)
䎓䎑䎔䎙
䎓䎑䎔䎗

(b) Case B: many embolus are shown in heartchamber.

䎓䎑䎔䎕䎵䏄䏗䏈 䎓䎑䎔

Fig.3 Result of endocardium extraction

䎓䎑䎓䎛
䎓䎑䎓䎙

Fig.4 is the time series graph of volumetric ratio of
the embolus. In fig. 4(a), the rate of embolus starts
increase after about 10 seconds from blood-flow
resumption. As for after 20 seconds, there are mostembolus, and it decreases gradually after that. In
addition, the rate of embolus is changing finely atintervals of short time through the whole observation
period. This minute change synchronizes with the- 296 -

䎓䎑䎓䎗
䎓䎑䎓䎕
䎓

䎚

䎛

䎜

䎔䎓

䎔䎔
䎔䎕
䎷䏌䏐䏈䎃䎋䏖䎌

䎔䎖

䎔䎗

䎔䎘

䎔䎙

(b) short period case (10s)Fig.4 Measurement result of rate of the embolus
in the right atrium of heartP2-43

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 297-298
16-18 November, 2005

Dynamic and Precise Visualization of Contrast Agent in Blood Vessels
with Motion Correction

Hideki Yoshikawa, Takashi Azuma, Kazuaki Sasaki, Ken-ichi Kawabata and Shin-ichiro Umemura
Central Research Laboratory, Hitachi, Ltd. Tokyo, Japan
1. INTRODUCTION

artifact is demonstrated. If the minute vessels in each

Vascularity gives very important information onframe are moved by the tissue motion, the signals

properties of tumors such as borders or stages.

from the agent cannot be accurately connected to

Especially for detecting borders between normal andform a line structure. This motion artifact may lead totumor tissues, not only large vessels but also minutemisdiagnosis

ones are to be visualized because large portions of

inappropriate treatment. To average signals in a

vessels into tumors are minute at the end. One canspecific region, both measurement and correction

obtain precise information on tumor tissues withmethods of the tissue motion, which are caused bydetailed tumor vasucularity which may lead tonormal physiological nature, such as the breathing or

accurate diagnoses. Also, such information would be

the cardiac motion, are needed.

andpotential

prescriptionofuseful in monitoring the effect of tumor therapy.
Contrast agent

As a relatively low invasive method to visualize
tumor

vascularity,

ultrasoundechography

Averaged image

iscommonly uses with contrast agents made ofWithout tissue motion

stabilized microbubbles. Conventional ultrasonic

Averaged image

imaging, however, cannot clearly visualize minutevessels mentioned above, because too small numbersof contrast media can pass the vessels at a time.

With tissue motion

Moriyasu et. al, reported on a method toFig. 1 Influence of the tissue motion for averaging frames.

visualize minute vessels by averaging multiple
frames[1]. Although each contrast agent in the minutevessels is imaged as an individual dot signal in theconventional contrast image, their method makes
virtual flow lines of contrast agents by overlapping
the dots signals. Using this method, the minutevessels can be imaged as a line structure (upper
images in Fig. 1).
Such approach enables an operator to clearly
identify the minute vessels and observing dynamic
inflow to the vessels of the contrast agents. However,this approach cannot simply be applied to objects

We propose a method to produce a clear imageof the vascularity of both large and minute vessels byusing the time-averaging method with multiple
frames which can be applied to the moving objects.
The purpose of this study is 1) to investigate thefeasibility of our tissue-motion measurement andcorrection method[2], and 2) to apply the correction
method to the time-averaging technique in in vivo
situation including visualizing the vascularity of theminute vessels using the time-averaging method withmultiple contrast images.
2. METHODS

with large motion, which may lead to the motion
artifacts. In the lower image of Fig. 1, the motion---------------------------------------------------------------hi-yoshi@crl.hitachi.co.jp

We performed in vivo experiments with rabbit
kidney filled with the contrast agent Definity®.
Contrast image frames were obtained from anultrasonic scanner EUB-8500 (Hitachi Medical Corp.,- 297 -

Chiba, Japan) and linear array probes which were(b).

fixed by hand. The Base-frame was set from the

Figure 3 (a) is the time-averaged image obtained

obtained frames, and other frames were then added towith motion correction of mesenterium vessels using

the Base-frame by using motion correction. Then,

150 frames and Fig. 3 (b) is the time-averaged image

subregions were set on the Base-frame to measureobtained without motion correction. The vascularity[2]

tissue motions . The subregion size was 3×3 mm

in Fig. 3 (a) can be imaged more precise and clearer

laterally and longitudinally, which was larger than

than in Fig. 3 (b). We also found that the

speckle size in this experimental condition. Tissue

tissue-motion measurement and correction method

motion was measured between each subregion on thewere performed accurately by comparing both images

Base-frame and obtained other frames using the leastin Fig. 3.

square method. The region that best matches the set(a)subregion was extracted and then added to theBase-frame. This extraction and addition of matchingregion was performed in all subregions and in all
obtained frames.

5mm

3. RESULTS AND DISCUSSION(b)

Figure 2 shows the time-averaging results of a
rabbit kidney using contrast images with motion
correction. Figure 2 (a) is the averaged image
obtained using 100 frames, and Fig. 2 (b) is a typical

5mm

conventional contrast image used to obtain Fig. 2 (a).

Fig. 3 Comparison of time-averaged images obtained with(a) and without (b) tissue motion correction.

(a)

The results of the precise imaging of the minutevascularity shown in Fig. 2 (a) and in Fig. 3 (a) werecaused by the slight tissue motion in the slice
direction. Tissue normally moves three-dimensionally

5mm

and an ultrasound beam scanning an object has a slice
direction range so that signals from various angles in(b)

the slice direction are included in obtained multiple
frames. Therefore, dynamic signals that were lessthan the slice range, like speckle signals, werereduced and static signals, like vessels, were clearly

5mm

imaged as a continuous line protectively and the highFig. 2 Comparison between time-averaged image with motion

contrast imaging of vascularity could be performed.
REFERENCES

correction (a) and single frame (b) of the contrast mode.

When

comparing

both

images,minute[1] F. Moriyasu et. al, Jpn. J. Med. Ultrason. 31

vascularity was visualized more clearly in Fig. 2 (a)(2004) 241.

than in Fig. 2 (b). Especially, the peripheral vessels in[2] H. Yoshikawa et. al, J. Appl. Phys. 44 (2005)

the renal cortex were imaged as a precise line

4561.

structure in Fig. 2 (a), which cannot be seen in Fig. 2- 298 -

P2-44

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 299-300
16-18 November, 2005

Behavior of Marine Animals Using Underwater Acoustic Camera○Kohji IIDA1), Yong Tang1), Tohru MUKAI1), Rika Takahashi1), and Masanori SATO2)
1)Graduate School of Fisheries Sciences, Hokkaido University, 2)Honda Electronics Company, Ltd.
1. INTRODUCTION

95dB, is displayed on a CRT monitor as a 256 levels

A high resolution underwater acoustic camera enclosed in a

monochrome image, and is outputted the video signal as NTSC

pressure-resistant case was developed to observe underwater

format. An optical camera consists of a CCD color video camera

animals. It enables to measure size, shape, and movement ofand an infrared LED projector in a cylindrical waterproof case.

living marine animals. Acoustic camera is originally used for

An optical camera was set beside the acoustic camera directed

medical diagnosis. A convex type transducer array is driven by

to the forward.

3.5 MHz ultrasonic signals and the echo from an object isBoth acoustic and optical camera were enclosed in adisplayed as an acoustic tomogram. It covers a range up tocylindrical pressure-resist case (100m in maximum) which was

240cm in waters. The acoustical B mode image and an opticalmade of aluminum alloy with 480mm of diameter and 450mm

image by an underwater camera were transferred throughof length. It was placed on the 1m height of aluminum frame

underwater cable.

stand and set at the sea bottom. The length of 300m underwaterSpecial features of the underwater acoustic camera are, 1) To

cables was used to supply electric power to the underwaterobserve marine animals in dark or opaque water, 2) To observe

system and to transmit the video signals to facilities on shore.

the tissue and shape of internal organ, 3) To observe the shapeand movement of marine animals in natural conditions.

Table 1. Specification of underwater observation system.

The underwater acoustic camera was set on a sea bottom toComponent

Item

Specification

observe behavior of marine animals. Several species of marine

Frequency

3.5MHz / 5.0MHz

Resolution

Direction 2mm / Distance 3mmanimals including zooplankton were observed during theRange

0.24〜1m

Probe typeLinear / Convex

Display mode

B-mode, M-mode

Video format

NTSC

Power supply

AC100V

Camera type

CCD camera

Video format

NTSC

Lighting

Infrared LED Projector

Power supply

DC12V

Acoustic
Cameraperiods.

2. METHODS
Optical
CameraUnderwater acoustic camera: To observe marine organisms innatural condition, an underwater acoustic animal observation

Casingsystem was developed. The system consists of acoustic camera
and underwater optical camera in the pressure-resistant case as

Monitorshown in Fig. 1.

Size

500W×450D×1260H (480φ)

Cable

300ｍCoaxial cable

Pressure-resistant

Depth 50m

Display

CRT Display

Video recorder

DV, VHS, Computer

Observation of marine animals in natural condition at sea:
Field experiments were conducted in July 2004 and June 2005
at the east coast of Oshima Peninsula, northern Japan where is
Usujiri Fisheries Experiment Station, Hokkaido University. An
acoustical/optical underwater animal observation system wasinstalled on the sea bottom at the depth of 12m. Both acousticaland optical images of marine organisms were continuously
recorded to VCR for 5 days in natural condition. And the tapes
Fig.1. Underwater acoustic camera.

were replayed simultaneously to compare two images foranalyzing species, size, and swimming behavior of animals.

An acoustic camera uses convex type ultrasonic probes whichworks at 3.5 MHz and forms fan shaped acoustical beam using

3. RESULTS

192 elemental beams with the coverage angle of 60 degrees andSeveral marine organisms in the video tapes were recorded.

the range of 240 cm. Acoustical image has a dynamic range ofGenerally, marine animals appeared more in nighttime than in- 299 -

daytime. In the daytime fish were appeared in optical camera
50

shown in Fig.2. In the night time many animals including
zooplankton were observed in acoustic camera. However, the

250

40

Fish
Plankton

30

200
150

20

100

22

20

18

16

14

12

Main animals which appeared in daytime were sand lance, sea

8
10

0

6

including zooplanktons were recorded by acoustic camera.

4

50

0

short range. However a large number of marine organisms

2

10

0

optical camera with an infrared light projector was restricted

300
Plankton (Number/min.)Fish (Number/5days)

and some of them were also displayed in acoustic camera as

Time (Hour)smelt, jelly fish, Arabesque greenling, etc. Especially a coupleof sand lance was showed spawning behavior and looked likeFig.3. Variation of marine animals in acoustic camera.

holding territory to other male individuals.
Many different kinds of animals were observed in nighttime by

4. DISCUSSION

both acoustical and optical cameras. Swarms of small organisms

Another merit of an acoustic camera is to observe marine

such as zooplankton were intermittently appeared. Fish,organisms under dark and muddy waters. Especially as anzooplankton, and unknown animals with long tail were also

acoustic camera provides B-mode image which can beobserved. Since an infrared light was attenuated dramatically inmeasured certain distance from camera to the object, the size

water, it was difficult to confirm that the animals of opticaland the shape of animals can be measured, whereas an opticalimages were the same as those in acoustical camera. Only few

camera only provides an apparent shape depending on thefish who were passed through in short distance from the camera

distance to an object. Shapes, sizes and swimming speeds of fish

was identified.

and zooplankton were observed by acoustical image in the field
experiment. These functions are applicable to monitor behaviors
of marine animals in the underwater facilities.

Fig.2. Fishes appeared in an acoustic camera.
Figure 3 shows a diurnal variation of organisms appeared inacousticalimage.

Appearancesof

zooplankton

Fig. 4. Unknown animals passed through in short range.

wereconcentrated between sunset and sunrise. The other fishlike

Acknowledgments

images were increased after sunset and before sunrise during theThis research was supported by the Ministry of Education,

field experiment.

Science, Sports and Culture, Grant-in-Aid for Exploratory

Even though it was difficult to measure the sizes of animals by

Research, 15658057, 2003-2005.

optical images due to distortion, acoustical images could
provide correct shape and size of the animals in a display.

REFERENCES

Swimming speeds of animals were also measured. Internal

1. Kohji Iida, Tohru Mukai, Donhyug Kang, Masanori Sato, “MorphologicalObservation of Marine Organism by Underwater Ultrasonography”,
Proceeding of OCEANS‘04, 357-36 (2004)
2. P. N. T. Wells, "State-of-art of ultrasound imaging in medicine andbiology", Acoustical Imaging, Vol. 22, 137-144(1995)
3. R. K. Hansen and P. A. Andersen, "A 3D underwater acoustic camera- Properties and applications," Acoustical Imaging, 22, 607-611(1995)

structures including a vertebra and a swimbladder of swimmingfish were recorded as shown in Fig. 4. This animal was thought
to be a bladdered fish having a large gill and the body length
was estimated to 18 cm.

- 300 -

P2-45

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 301-302
16-18 November, 2005

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Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 303-304
16-18 November, 2005

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Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 305-306
16-18 November, 2005

An Approach for Tonal Signal Automatic Recognitionof Ship Radiated Noise◎ Kyu-Chil Park, Phil-Ho Lee* and Jong-Rak Yoon
(Pukyong National University, Korea, * Agency for Defense Development, Korea)
1. Introduction

Line spectrum

The ship-radiated noise appear the various
characteristic signals due to the mechanic system inthe ship, the propeller and the interaction between
ship body and sea water. It is important
informations for the ship's design, the maintenance,
the operating condition dignosis and machinery
monitering. Generally, it is classified the speed
dependent component and the speed independent
component.[1-3] Therefore, it is required that someprocedures to classify the signal origin from theship-radiated noise.
This paper presents a technique to detect the tonal
signal and classify the tonal signals from ship
radiated noise, using peak extraction and Q-factor.
[4]

LOFAR gram

Fig. 1

An example of line spectrum and LOFAR
gram from ship-radiated noise.

2. Detection of tonal signals
Fig. 1 shows an example of a line spectrum and a
LOFAR-gram from ship-radiated noise. It is
required to classify what tonal signal is important
and what origin of tonal signal is. As shown Fig. 1,
the tonal signals have some characteristics – peaks
and higher amplitudes. And the speed dependent
components have some frequency variation widthin some time zone. In this paper, we approach thestudy as following:
1) Peak extraction from line spectrum
2) Selection of tonal signals exceed thresholdamong results of step 1)3) Classification of tonal signal by checking thefrequency variation width (by Q-factor)
In step 1), the peaks are extracted by differential
coefficients ( ∆x1 − ∆x2 ≥ 0 ) from adjacent two
points shown as Fig. 2.
-----------------------------------------------------------E-mail : kcpark@pknu.ac.kr

Fig. 2

Peak extraction of tonal signal bydifferential coefficients

For step 2), the threshold value is given the half ofmaximum spectrum amplitude.
For step 3), we apply the Q-factor to classify two
components - the speed dependent component(SDC) and the speed independent component(SIDC). Fig. 3 shows an example of the bandwidth
calculation from a test signal with Gaussian
distribution of 20Hz divergence. At the top, fromleft, the frequency variations are 0, 1, 2, …, 7Hz.
The rest show the amplitudes variations and the
bandwidth variations. From Fig. 3, we obtain thatthe frequency variation in the SDC is larger than the
SIDC. The frequency variation width of the tonal
signals are calculated in 22 frames. We classify the
tonal signals by counting the number of the peak
extracted signal in each frequency domain.

- 305 -

frequency components below 50Hz.

(a) Real ship signal’s spectrogram

Fig. 3

An example of the bandwidth calculation(b) Exceeding threshold and peak extraction signal(a) Test ship signal’s spectrogram(c) Classified signal by Q-factor
Fig. 54.
(b) Exceeding threshold and peak extraction signal

Conclusions

This paper presents techniques to detect andclassify the tonal signals from ship-radiated noise,using peak extraction and Q-factor. We obtained thereasonable results of the classification of tonal
signals using frequency variation width infrequency domain.

(c) Classified signal by Q-factor
Fig. 4

Classification result for real ship signal
applied by Q-factor

Acknowledgment
This work was supported by STX Engine Co.,Ltd.
and the Brain Busan 21 Project.

Simulation result for test ship signal
applied by Q-factor

3. Numerical experiments and results
For the numerical simulation, we made test signals
with 1 SIDC and 9 SDCs and 20% Gaussian noisesuch as Fig. 4. We exactly classified 1 SIDC
marked 'o' and 9 SDCs marked 'x'.
At next, we apply this approach to real achieved
noise data from ship radiated as shown Fig. 5. The
horizontal two lines represent 22 frames forclassification around CPA(Closest Point ofApproach). We successfully classified 3 SIDCs
marked 'o' and 10 SDCs marked 'x' except very low

References
1. Chu-Kuei Tu, Huang-Chia Huang: UnderwaterTechnology (1998) 253.
2. Phil-Ho Lee, Jong-Rak Yoon, Kyu-Chil Park,
Ki-Hun Lim: Acoustical Society of Korea, 23,(2004) 175.
3. Thomas L. Hemminger, Yoh-Han Pao: IEEE 5(1994) 712.
4. C.K.Alexander and M.N.O.Sadiku: Fundamentalof Electric Circuits 2nd Edition (McGRAW HILL,
2003) p.632.

- 306 -

P2-48Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 307-308
16-18 November, 2005

- 307 -

- 308 -P2-49Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 308a-308b
16-18 November, 2005

Dynamic Behavior of Multibubble Cavitation Dur ing Ultr asonic Degassing○Takashi Kubo, Noriyasu Kawakita, Mamoru Kuwabara and Jian Yang
(Dept. of Material Processing Eng. Nagoya Univ.)
Abstr act
A water model experiment has been carried out toclarify the effects of atmospheric pressure, the powerof ultrasound imposed and concentration of suspended
fine particles on the rate of degassing with acoustic
cavitation. It is clarified that operating factors
described above accelerate the rate of ultrasonicdegassing. Furthermore dynamic behaviors of acoustic
cavitation during degassing have been visualized by ahigh-speed digital video camera, and the observed
behaviors were well correlated to the rates ofultrasonic degassing for corresponding operating
conditions.

depth. Distilled water, which was prepared for thesolution, was pre-degassed by stirring it for about 1.5hours under the vacuum atmosphere. Atmosphericpressure above the solution was adjusted at a specified
level with a vacuum pump. The pH meter to indicate
variation of CO2 concentration in the NaOH solution
intermittently detected its transitional change. The
influences of the atmospheric pressure, frequency andthe concentration of dispersed fine particles on thedegassing speed were investigated. The fine particles
used in this experiment were polystyrene balls of
241μm in diameter and 1050kg/m3 in density.
Furthermore, the behavior of the cavitation during
irradiation of ultrasound was photographed with a

1. Introduction
Sonoprocessing of Materials in which ultrasound isapplied to material processing operations can create
many innovative technologies [1,2]. Acousticallyinduced cavitation multibubble is especially expected
to cause high efficient degassing, because of increasein gas-liquid interface generated as acoustic cavitation
multibubble, and because of intensive mixing of liquidby microjets generated when the bubbles collapse [3].
For high efficient ultrasonic degassing, it is important
to investigate the operating factors described below. In
this study, a water model experiment has been carriedout to clarify the effects of atmospheric pressure,frequency, the power of ultrasound imposed andthe amount of suspended fine particles on the rate ofultrasonic degassing.

Fig.1 Experimental setup.

digital camera.

3. Results and discussion
2. Exper imental Setup
Fig. 1 shows a basic experimental setup for theultrasonic degassing of liquid. A continuous sinusoidal
signal generated by a function generator was
amplified by a power amplifier. Its electric output was
fed to a bolted Langevin-type transducer (resonant
frequency: 25, 46 kHz nominal maximum input
power: 55 W) or to a plate-type circularpiezoelectric ceramic transducer (resonant
frequency: 250 kHz) of 45 mm diameter. The
transducer was fixed to a stainless steel plate of 3 mm
thickness at the bottom of the vessel. The vessel was
made of a cylindrical acrylic resin of 3 mm inthickness and 74 mm in inner diameter. It was filledwith an aqueous 0.01N NaOH solution to a resonanceFig.2 shows the influence of the atmosphericpressure on the rate of degassing with or without
ultrasound of a frequency of 42.3 kHz and a power of
18 W. In the case of ultrasonic degassing, it wasobserved that many small bubbles evolved quietly andflew up orderly at the initial stage, and the acoustic
cavitation after several minutes became noisy andcaused violent motion of the liquid accompanied by
microjets. From Fig. 2, it was shown that the rate ofultrasonic degassing was drastically higher than thatof degassing without ultrasound, and was also shownthat the degassing rate was accelerated by the decrease
in atmospheric pressure in both cases. This result iscaused by the decrease in cavitation threshold withdecreased atmospheric pressure. Particularly at theinitial stage, acoustic cavitation remarkably- 308a -accelerated the rate of degassing and the rate became
moderate after an elapse of several minutes. Thesefindings suggest transition from gaseous acoustic
cavitation conditions to vaporous acoustic cavitation

solution to prepare a sugar solution of 12 mass% and0.01 N. An ultrasound of 41.9 kHz was irradiated into
the liquid with depth of 55 mm. It is found from Fig.4
that the rate of degassing was promoted by thedispersed fine particles whose surfaces might favor
nucleation of cavitation bubbles. The promotion ofdegassing with the dispersed particles was remarkableat the initial stage. This is why the surface of fineparticles can provide the effective nucleation sites
only during the period of gaseous cavitation.
Additional effect can be expected such that dispersed
particles move from around antinodes of soundpressure, where no acoustic cavitation bubbles are
generated, to low sound pressure area such as nodes ofsound pressure or surface in the vessel, where acoustic
cavitation bubbles substantially exist. Figure 4 showthat the influence of dispersed particles tended tosaturate when the particle concentration Cp exceeded
0.032 mass%.

Fig.2 Effect of atmospheric pressure on the rate of degassing
with or without irradiation of ultrasound.

ones.
Fig.3 shows the comparison of low and highfrequencies of ultrasounds on the rate of degassing.
Imposed frequencies were 42.3 kHz and 245.5 kHz
keeping a power of ultrasound as 18 W, and the liquid
depth was set in resonant condition for each working
frequency. At the initial stage, the rate of degassing
with low frequency of ultrasound was higher than thatwith high frequency, and finally the degassing ceased.
Whereas, in the case of degassing with high frequency
ultrasound, the rate of degassing was lower than thatwith low frequency of ultrasound at the initial stage ofdegassing. However the degassing continued till the
final stage of experiment. In the case of highfrequency ultrasound, acoustic streaming was
generated toward direction of ultrasound irradiated,
and acoustic cavitation multibubble was continuously
removed by the acoustic streaming.

Fig,3 Effect of frequency of ultrasound on the rate ofdegassing at constant atmospheric pressure.

Fig.4 shows the influence of fine particles dispersed
in liquid on the rate of ultrasonic degassing. In orderto adjust the density with the particles and to dispersethem uniformly, sugar was mixed with the NaOH

Fig,4 Effect of particles on the rate of degassing with irradiatedultrasound.

4. Conclusion
In this paper, the effects of atmospheric pressure,frequency and concentration of dispersed particles on
ultrasonic degassing have been investigated. The
acoustic cavitation plays an important role in thedegassing process, and the decreased atmosphericpressure, the power of ultrasound imposed, and thedispersed fine particles all accelerate the degassing
rate. The behaviors of cavitation multibubble during
ultrasonic degassing were observed by a high-speed
digital video camera in the same experimental
conditions described above. Observed behaviors of
acoustic cavitation bubbles were well correlated to therates of degassing of melt for corresponding operating
conditions.
Refer ence
[1] M. Kuwabara, S. Sano, S. Yamahira, and M. Sano: Proc.
Int. Conf. World Congress of Ultrasonic (1997) p.408.
[2] M.Kuwabara, S. Komarov and M.Sano: Proc. 8thJapan-China. Symp. Science and Technology of Iron and Steel(1997) p.175.
[3] T. Kubo, M. Kuwabara and Jian Yang: Jpn. J. Appl. Phys.
44 (2005) 4647.

- 308b -G-1

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 309-310
16-18 November, 2005

Applicability of FDTD methodon the wave propagation in the cancellous bone
◎Yoshiki NAGATANI, Hirotaka IMAIZUMI, Takashi FUKUDA,
Mami MATSUKAWA, Yoshiaki WATANABE and Takahiko OTANI (Doshisha Univ.)
1. Introductionxyx

Longitudinal wave propagating in thecancellous bone separates into fast and slow

High

y
z

waves depending on the conditions1-3). This
interesting phenomenon becomes a strong tool
for the diagnosis of osteoporosis because thewave behavior depends on the bone structure.
Low(b) Distribution of
averaged bone volume fraction.

However, there are a few studies applying (a) X-ray micro focus CT image.
FDTD (Finite Difference Time Domain) method

Figure 1 Cancellous bone.

to such complex coexistent models of liquidand solid.

averaged volume fraction of the sample.

In this study, the applicability of FDTD method
to the cancellous bone is discussed focusing on the

Simulation field is 17×17×13[mm] with cube
lattice of 46[µm]. Time step is 5[ns]. As an initialpropagation of fast and slow waves.

particle velocity at the surface of the concave

2. Elastic FDTD method4)

source, a single sinusoidal wave of Fig. 2 is used.

The governing equations of elastic FDTD for the

The surface of the transmitter is star-shaped in orderto decrease the edge effect.

∂v
∂σ xx
∂v
∂v
= ( λ + 2 µ ) x + λ y + λ z (1)
∂t
∂x
∂y
∂z
∂σ xy
 ∂v ∂v y =µ x +
(2)

∂t
∂x 
 ∂y
∂vx 1  ∂σ xx ∂σ xy ∂σ zx 
= 
+
+
(3)

∂t
∂y
∂z 
ρ  ∂xwhere σij is normal and shear stresses, vi is particle
velocity, λ and µ are Lame coefficient, and ρ isdensity of the medium.

( ) (for 0 ≤ t ≤ 2π ω )sin ω t × sin 2 ω t
2v (t ) = 
2
π
for
<t
0
ω


(

Velocity [m/s]isotropic medium are;)[m s ]

0.4
0.0

-0.4
0

2

4

6

8

10

14x10

Time [s]Figure 2 Waveform of initial particle velocity.

The transformation of these equations into
difference equations enables the alternate calculationof stress and particle velocity, which realizes thenumerical simulation of the field.

yzxConcave Transmitter

Water

3. Simulation Model
The size of bovine cancellous bone used was
14×14×7[mm]. Figure 1((a)) is the x-ray micro

Focal Point

focus CT image and (b)) is the distribution ofCancellous Bone1[mm]ReceiverE-mail: nagatani@usl.doshisha.ac.jp

Figure 3 FDTD simulation model.

- 309 -

-6

Amplitude

The sample is fully immersed in water. The z
axis position of the transmitter is fixed to keep the
focal point at the half length of the sample
thickness. Changing positions of transmitter and

0

-500

-1000

sensor in the x-y plane, the distribution of wave

0

propagation was investigated. As the acousticofthecortical

bone

4

are

appropriated.

8

-6

10

14x10

(b) with samp le

Slow wave
0

4. Results and Discussion

6
Fast wave

Amplitude

values

2

100
50
0
-50
-100

constants of cancellous bone, the experimentallyobserved(a) without sample

500

2

4

6

8

10

14x10-6

Time [s]4.1. FDTD Simulation

Figure 4 Simulated waveforms.

4(
(a)
)
is the calculated pressure
waveform of the model without the sample.

FigureThe wave observed at the cross(×) point of

x

1.8µs

0.02

0.6µs

0.6
(b) Peak amplitude of slow wave.

yFig. 1(b) is shown in Fig. 4((b)). Both fast andslow waves can be seen. The fast wave isconsidered to propagate mainly in the hard
structure of bone sample. Moreover, theamplitude of the fast wave was much smaller
than the slow wave, which has also been(a) Arrival time differenceof fast wave.

1-2)

reported experimentally .

( Normalized by the wave in water. )( 0sec is the arrival time of wave in water. )Figure 5((a)) is the distribution of arrival

Figure 5 Distribution of arrival time and amplitudeof simulated waveforms.
x
2.4µs
0.1

time of the fast wave. The amplitude of slow
wave is shown in Fig. 5((b)). Figure 1(b) and
Figs. 5(a)(b) indicate that the arrival time ofyfast wave and the amplitude of slow wave havegood correlations with the distribution of bonevolume fraction in Fig. 1(b).
4.2. Experiments
Distributions of fast and slow waves areexperimentallyobservedusing(a) Arrival time differenceof fast wave.

PVDF

hydrophone. The focal length of transmitter is

1.5µs

0.5
(b) Peak amplitude of slow wave.

( 0sec is the arrival time of wave in water. )( Normalized by the wave in water. )40[mm] and the radius is 10[mm]. A single
sinusoidal wave at 1[MHz] is transmitted. The

Figure 6 Distribution of arrival time and amplitude
of experimentally observed waveforms.

radius of the receiver is 2.5[mm].

Acknowledgment
Authors would appreciate the measurement of x-ray
micro focus CT to Dr. Murata and Dr. Takada at Shiga
University of Medical Science.

The result shown in Fig. 6 shows similartendency with the FDTD data.
5. Conclusion
Using the CT model of the actual cancellous bone,
the generation of fast and slow waves is confirmed
by elastic FDTD method. This clearly shows theapplicability of FDTD method to understand thecomplicated wave propagation in cancellous bone.

References
1. A. Hosokawa, T. Otani : J. Acoust. Soc. Am, 101(1997) 558.
2. A. Hosokawa, T．Otani, T. Suzuki, Y. Kubo, S.
Takai : Jpn. J. Appl. Phys. 36 (1997) 3233.
3. T. Otani : Jpn. J. Appl. Phys. 44 (2005) 4578.
4. Y. Tanikaga, K. Takaki, Y. Watanabe, T. Sakaguchi: IEICE Technical Report, US2003-3 (2003).

- 310 -

G-2

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 311-312
16-18 November, 2005

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- 312 -

G-3

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 313-314
16-18 November, 2005

Quantifying and Qualifying Sea Bottom Backscattering Strength by
Quantitative Echo Sounder◎

1.

Henry M. Manik, Masahiko Furusawa, and Kazuo Amakasu(Tokyo University of Marine Science and Technology)

Introduction

Underwater acoustic technologies are mosteffective and useful for characterization of sea
bottom due to their reliability and versatility.
They are based on the measurement, analysis, andinterpretation of characteristics of signal reflectedor scattered by the sea bottom.
One of the effective and promising acoustic
methods to measure bottom surface backscatteringstrength (SS) is to use a quantitative echo sounder(QES) originally developed for fisheries surveys.
The QES can rather easily measure echoes due
primarily to scattering from the water-bottom
interface. A sea bottom backscattering strength
theory was developed by one of the authors 1 forchecking the QES performance.
In our previous study 2 we developed a ring
surface scattering (RSS) model for sea bottom
scattering as an extension of the above theory andcompared the SS values obtained by the RSS
model with the results from the bottom echo
integration model to confirm the reliability of thetheory. In this study, we apply our bottom echo
integration model and RSS model extensively to abottom survey by QES off Jawa Island andexamine the effectiveness of our models.
2.

Methods

A QES (KaijoSonic, KFC3000) installed
onboard the research and training vessel″Umitaka-maru″ of Tokyo University of Marine
Science and Technology was used to measure SS.
The simplified block diagram of the QES tomeasure SS is shown in Fig. 1. The QES wascalibrated with a 38.1 mm φ tungsten carbide
sphere following the standard method.3
In the bottom echo integration method, 1 weaverage the bottom echoes for a predefined depth
layer (r to r + rw) including the bottom echoes andfor a ping sequence or integration period to obtain
the average bottom echo integration strength("bottom SV") as
ETB

2

(1)
2
KM
where ETB is the time varied gain (TVG) amplifier
SVB =email : md02207@edu.s.kaiyodai.ac.jp- 313 -

output, KM is the multiple echo coefficient, and< · > exhibits the averaging. The simple relationship between the average SS and the bottom SV isgiven by 1

S s = rw SVB(2)

where rw is the integration width.

Fig. 1

Trans mitter

Pulse
Gen.

PreAmp.

20 log r
TVG

Echo
Processing

SVB ,
Ss

Simplified block diagram of QES tomeasure SS

Using the RSS Model we have a relationshipbetween the raw SV value of the bottom echo, SVB ,and the raw surface scattering strength, Ss, as

Ss Φ
(3)
Ψ ( c τ/ 2)
where Φ and Ψ are the equivalent beam angle forSVB =surface and volume scattering, respectively, c is thesound speed, and τ is the pulse width. At the peakof the bottom echo, we haveΦ ≃ Φo(4)where Φo is the asymptotic value of the equivalentbeam angle for surface scattering. In the case of a
sharp beam we have 4

Φo ≅ Ψ .

(5)

Introducing Eqs. (4) and (5) into Eq. (3) gives

S s = ( c τ / 2) SVB(6)
and we can easily convert the raw SV to raw SS.
A bottom survey was conducted from 9 to 24
December 2003 in the Indian Ocean, southerncoast of Jawa, Indonesia in conjunction with theoceanographic and exploratory fishing.

It is shown that the QES can rather easily
measure echoes by reflection of sounding pulse
from the bottom. The QES is especially usefulwhen both fish and sea bottom should bequantified. Figure 4 shows simultaneously thescattering from fish (SV) close to the sea bottom
and the SS values by Eq. (6) of the fixed andexpanded bottom.

3. Results and Discussions
An example of the average SS map obtained by
the echo integration of bottom echo withintegration period of 0.1 nmi of bottom echo isshown in Fig. 2 with depth contour deduced fromthe QES data ; the SS of this area are ranging from-23 to -12 dB. We observe an evident correlation
between the bottom slope and the SS value : SSvalues are generally high where the bottom slope islarge as seen at the south-west corner of Fig. 2.
The instantaneous SS derived from RSS modelat point A of Fig. 2 is shown in Fig. 3.
SS [dB]
7.98
-12

8

110
110

-14

110

-16

0
11

8.04

110

120

LAT [deg]

8.02

-18

120

Fig. 4

120

8.06
120

-20

130
14
0

8.08

A
8.1
108.6

130

130

We can correlate the bottom material estimatedby the SS value and bottom fish. From themeasured SS value, the bottom type is estimated as
sand. The SV values of fish close to the sea
bottom are ranging from -60 to -30 dB. From thislike display, the assessment of fish and their habitatby bottom material is possible.

-22
150

140

108.65108.7

108.75108.8

108.85108.9

LON [deg]

Fig. 2

The average SS map along the acoustic
track
SS [dB]

-5

-10

-15

-20

-25

-30

We summarize our study as follows :

-35

151

The SS values can be measured by the bottom echo
integration method and the RSS model.

152

Depth [m]Measured scattering strength of fish (SV)close to the sea bottom and the SS sea
bottom

153
154
155

The efficiency of QES to measure scatterings bothfrom fish (SV) and bottom (SS) was shown.

156
157

Bottom Fix Range [m]21

4. Acknowledgment

0

We acknowledge JSPS-DGHE Core University
Program for Fisheries Science for the collaborative
research between Tokyo University of Marine
Science and Technology and Bogor Agricultural
University (Indonesia).

1

2
150

170

190

210

230

250

Ping Number

Fig. 3The instantaneous SS of point A (upper)
with the fixed and expanded bottom(lower)

References
1. C. Aoyama, E. Hamada and M. Furusawa :Nippon Suisan Gakkaishi. 65 (1999) 78-85.
2. H.M. Manik, M. Furusawa and K. Amakasu :J. Fish. Sci (submitted).
3. Y. Miyanohana, K. Ishii and M. Furusawa :Nippon Suisan Gakkaishi. 59 (1993) 933-942.
4. C. Aoyama, E. Hamada and M. Furusawa :Nippon Suisan Gakkaishi. 63 (1997) 570-577.

The upper figure is the sea bottom image with
undulation caused by the ship motion and thelower figure is the bottom image after bottom
fixing and expansion. The SS value read fromthis figure is about -20 dB and the value by bottom
echo integration methods seen in Fig. 2 are in good
agreement.
- 314 -

G-4

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 315-316
16-18 November, 2005

Measurement of Fish School Volume Using
Omnidirectional Multi-Beam Sonar, -Scanning Mode and Algorithm-○Yong Tang, Kohji Iida, Tohru Mukai, (Univ. Hokkaido, Fish. Sci.) Yasushi Nishimori (Furuno Co., LTD）
For the circular mode DA-3, since the fish school is scanned

1. Introduction

with a conic plane, as shown in Fig. 2b, the angle αx which

Since the pelagic fishes show features that tend to gather into

between nx of dS and x-axis can be described aslarger stocks and swim near sea surface, it is difficult toinvestigate by using a quantitative echo sounder with a sharpcos α x = cos θ b cos( π / 2 − β )vertical beam, because a smaller sampling and fish avoid

and the volume can be expressed as

approaching vessels [1].

Therefore, the quantitative sonar

x2

θb 2

r2

b1

r1

(2)

(Furuno FSV30 research version) which have omni directional

V H = cos β sin β ∫multi-beam was developed, and makes it’s possible to measure

At anchored modes DA-4 or 5, the volume can be estimatedx1∫θ ∫ r cos θ drdθ d x .
b

b

(3)

by volume integration method (Fig. 2c), with the formulaboth the volume backscattering strength (SV) and volume of fish
school simultaneously. In this study, the method for estimating

V BT = ∫

θb 2

θ b1

the volume of fish school is proposed using various scanning

β2

r2

1

r1

∫ β ∫ r | sin(π / 2 − β ) | drdβdθ
2

b

(4)

where the β and θb are zenith and azimuth angle respectively.

modes of sonar. A computer simulation is conducted to clear
the characteristic of measured volume, and the valid selection is

Also these formulas could be used for biplane mode in terms

discussed for estimating the volume of pelagic fish school.

of the states of ship at curious or anchor.

2. Quantitative sonar and measurement method of fish school

3. Simulation of measuring fish school volume

The quantitative sonar is constructed by FSV30 sonar

A simulation was conducted for measuring the volume of(24kHz), a data recording unit and analysis tool [2]. There arespherical scatterer. The parameters of acoustic and fish school

six modes for investigating fish school radiating 64 beams fromare showed on Table 1. The insonified acoustic pressure to

the spherical transducer, including on cruise scan (DA-1,2,3),transducer by individual scatterer of fish school which given abeam scan (DA-4,5) and biplane scan (DA-6)(Fig. 1).

random orientation, and for each beam, can be expressed using

The volume can be estimated by integrating the plural

the sonar equation as

projective of echo area on y-z plane along x-axis assigned to the

Prij2 ( t ) = P02 ri − 4 10 − 0 .2 α ri D sij2 D rij2 T Si u 2 ( t −course of ship at mode DA-2(Fig. 2a), expressed as

VS = ∫

x2
x1

θb 2

r2

b1

r1

∫θ ∫ r cos α drd θ dx
x

2 ri
)
c

(5)

(1)

b

where x is cruising distance, r and θb are distance and azimuth ofecho respectively, αx is the angle of normal line nx of scanningplane to the x-axis. Otherwise, at mode DA-1 αx is zero.
DA-1

DA-2

DA-3

DA-4

DA-5

DA-6

Fig. 1 Schematic diagram of six measurement modes of FSV30 sonar

Fig. 2 Integration method for various modes of sonar. a shows integral area atmode DA-1,2, b is for mode DA-3, c shows integral volume of echo at mode
DA-4,5.

Yong TANG , tang@echo.fish.hokudai.ac.jp- 315 -

Table 1 Parameters of sonar and spherical scatterer for computer simulation
Measure mode

mode DA-2,3,4,5

Beam widthtransmission
reception

180°×12.4° (360°×11.9° at DA-3,5)13.6° (11.9° at DA-2, 4)

Pulse width

2ms

Source level

216dB

Displayed angular resolution

2.86° (5.71°at DA-3, 5)

Displayed range resolution

0.375mdistance at ping ratio

4m

Target strength-40dB (normal distribution)

Diameter of spherical scatterer

20m

Interval of individual scatterer

1m (random distribution)

Threshold of signal level-75dB, -70dB, -65dB

Fig. 3 Measured volumes of spherical scatterer at 30m depth using various tilt
angles and threshold of signal level at mode DA-2, 3

where P0 is source level, r is distance, α is absorption coefficient,
Ds and Dr are directivity functions of transmission and reception
respectively, u is the function of wave shape [3], i and j areindexes of scatterer and beam respectively. So, the insonified
acoustical pressure for each beam could be synthesized by all
scatterers, and it can be estimated using the amplitude of carrier
signal expressed as
n

n

i =1

i =1

PMj2 (t ) = (∑ Prij (t ) cos γ j ) 2 + (∑ Prij (t ) sin γ j ) 2(6)

where γ is the phase with a random distribution between 0~2π.
The directivity functions of linear and circular transducer are

Fig. 4 Measured volumes of spherical scatterer at 30m depth and various
azimuths using 30° tilt angle at mode DA-2 ,3used for transmission and reception approximately. Finally, thePM is amplified with 20logr TVG to generate the raw SV signal
and be displayed the same at a polar coordinate as a sonar.
4. Results
As shown in Fig. 3, we estimated the volume of sphericalscatterer which ahead of ship at 30m depth using mode DA-2
and DA-3. The measured volumes normalized theoreticalvalue show the stabilizing feature between 30°~90° tilt angle atmode DA-2 and 30°~60° at mode DA-3. So the small tilt is
not useful for measuring the fish ship volume.

Also theazimuthal characteristics were examined using 30° tilt angle.
Fig. 5 Measured volume of spherical scatterer at 30m depth and various

The volume can be measured within 30° azimuth from ship

distances using mode DA-4 and DA-5

head for mode DA-3, and within 70° for mode DA-2(Fig. 4).
So mode DA-2 is more available to estimate the volume ofpelagic fish school, although mode DA-3 can search for fish
school at a broad circular area.
For anchor mode, Fig. 5 shows the precision of estimatedvolume is higher at mode DA-4 than that at mode DA-5.
Furthermore, the anchor mode has a better precision than that at
cruise mode (Fig. 3, 4, 5).

Reference
1. F.Gerlotto, J. Castillo, A. Saavedra, M.Barbieri, M.Espejo and
P.Cotel: ICES J. Mar. Sci. 61 (2004)1120.
2. K. Iida, Y. Tang, T. Mukai and Y. Nisimori: OCEANS’04,
Proceeding, (2004) 401.
3. M.Furusawa, K. Isii and Y. Maniwa: J. Acoustical Soc. Jpn.
42 (1986) 2. (in Japanese with English captain)- 316 -

Proc.
3 Symp. Ultrason. Electron., Vol. 26, (2005) pp. 317-318
16-18 November, 2005

H-1

Integrated Ultrasonic Transducers
Made by Sol-Gel Spray Techniqueတ Makiko Kobayashi and Cheng-Kuei Jen(Industrial Materials Institute, National Research Council of Canada)
surface temperature of the substrate was kept around 300 and

1. Introduction
In-situ structural health monitoring, on-line and120°C, respectively. A high positive voltage supplied from a 28

off-line nondestructive testing (NDT), and on-line real-time

kV DC power supply was fed into a thin and sharp needle locateddiagnosis of industrial material processing are desired toseveral centimeters above the films coated on the substrates,

identify, characterize, assess the voids, defects and damages inwhich serves as the ground electrode. The poling time was about

many materials and structures for applications in aerospace

10 minutes. Finally, silver paste painting was used to form thevehicles, marine vessels, automobiles, nuclear power plants, gas,

top electrode at room temperature. This convenient approachwater and oil pipelines, etc and also ensure the quality of

makes the selection of electrode size, i.e. the sensor size, simple.

manufactured products of polymer and metal extrusion and

The silver paste had been tested and its operating temperaturemolding (or casting), etc [1-3]. Ultrasonic techniques are

could be up to at least 200°C. It is noted that PZT sol-gel solution

frequently used for these purposes because of their subsurface

contributed as bonding material between the BIT or PZT powder

inspection capabilities, fast inspection speed, simplicity, cost

and the substrates. The dielectric constants of the BIT/PZT andeffectiveness and ease of operation. The common limitations of

PZT/PZT films were 15-90 and 60-340, respectively, measured

the use of the conventional and non-integratable UTs are (1) the

by a Hewlett Packard 4192A LF Impedance Analyzer at 1 kHz

requirement of couplant, (2) complicated for curved surfaces, (3)and the difference in each UT comes from the heating methods

difficult for pulse-echo mode and (4) not efficient forused. The measured piezoelectric coupling constant kt was 4 andtemperature higher than 60°C. In this investigation, our

24% for BIT/PZT and PZT/PZT film UT, respectively.

technology focuses on optimizing the fabrication process of
3. Ultrasonic performances of UTs

handheld equipments and no furnaces that are significantly

Top ElectrodeUT1 2

different from the fabrication method reported in [4] in order to

BIT/PZT Film34integrate them onto flat and curved surfaces of metal and5

Feeding
Zone

graphite/epoxy (Gr/Ep) composites at the desired sensor
locations without the displacement of and harming substrates.

6

7

HeatingZone
Exit

2. Fabrication and characterization of UTs(a)Amplitude (arb. unit)

sol-gel sprayed piezoelectric thick (>40 µm) film UTs with

L1

2

260°C
o

L2

4

at 440 C

L3

6Time Delay (µs)

L4

8

(b)

Bismuth titanate (BIT) or lead zirconate titanate (PZT)

Fig. 1 (a) Integrated 90µm BIT/PZT composite film UT arrays

powders were purchased with a particle size distribution of abouton barrel for diagnostics of micro-molding and (b) its ultrasonic

80 and 2 µm, respectively. BIT is chosen because of high 675°C

performance in time domain at 260°C.

Curie temperature (Tc) and reasonable piezoelectric strength,
which are suitable for high (>350°C) temperature operations.

Figure 1a shows seven BIT/PZT composite film UTs

PZT is chosen because of relatively high Tc ~350°C, high(UT1-7) fabricated directly on the external surface of thepiezoelectric strength and dielectric constant. The powders werecylindrical barrel for a micro-molding device; Microsystem50

dispersed into PZT sol-gel solution by ball milling method tofrom Battenfeld, Austria. The diameter of the top electrode was 5

achieve the paint. The final BIT or PZT powders were estimated

mm, which is the active area size of the UT. The length of theof sub-micron size. An airbrush was then used to spray thesteel barrel was 265mm, the internal diameter was 14 mm, andsol-gel composite directly onto steel, stainless steel (SS), Gr/Ep

the external diameters at the areas of UT1-3, 5-7 and at that ofsubstrates, etc. In this way the films can be produced at desired

UT4 were 40 mm and 30 mm, respectively. In Fig.1b L1, L2, L3

locations through a shadow mask onto the samples. After sprayand L4 are respectively the 1st, 2nd, 3rd and 4th round-trip reflectedcoating, thermal treatments such as drying and firing were

echoes through the thickness of the barrel wall at UT6 location.

carried out using a heat gun or a gas torch. Multiple layers were

The pulse-echo mode was used. The measured center frequency,

made in order to reach the desired thickness and frequency (1 –6 dB bandwidth and signal-to-noise ratio (SNR) of the L1, by the

30 MHz) of interest. The films were then electrically poled using

UT6 at 260°C, were 13MHz, 13MHz and 31dB, respectively.

the corona discharging technique. For corona poling, the top

The SNR is defined as the ratio of the amplitude of the first echo

- 317 -

film UT was fabricated with all the thermal treatment carried out

echoes traversing back and forth in the substrate. The amplitude

by a heat-gun, which produced gentle heat in order not to harmand time delay variation of L1 can be used for monitoring thethe Gr/Ep substrate. A 70µm thick PZT/PZT film UT depositedproperties such as viscosity, filler composition, melt quality,

on such a composite of 12.7 mm thick is shown in Fig.3a.This

average temperature etc. of molten polymer in the gap between

Gr/Ep has sufficient conductivity to serve as the bottom

the inner barrel wall and the rotation screw inside the barrel. The

electrode of the UT. Figures 3b shows the ultrasonicperformance of UT1-7 array was almost the same. The BIT/PZT

performance of this integrated UT operated in pulse-echo mode

UT can operate up to more than 440°C [3]. The integrated UTs

in the time domain at 22°C. The measured SNR, center

have also been deposited onto aluminum and titanium substrates

frequency and 6 dB bandwidth of ultrasonic signal L1 were 15

and their performance is equally to that shown in Fig.1b.

dB, 1.6 MHz and 1.5 MHz, respectively. The low center

Amplitude (arb. unit)

L1 over that of the signals, which are undesired, between the

LL
22frequency and bandwidth were caused by the high ultrasonic

160°C
L
L4
4

attenuation within the thick specimen. The result demonstrates

LL6

the capability of integrated UT for in-situ real-time NDT

6

10

20

30

Time Delay (µs)

(a)

UT

(b)

Fig. 2 (a) Integrated 40µm PZT/PZT composite UT on a 75µm
thick SS membrane and (b) its ultrasonic performance in time

Ampliture (arb. unit)

application of Gr/Ep substrates.
L1

22°C
L2

10

L3

20

30

L4

40

Time Delay (µs)

domain at 160°C when attached to a steel pipe with outer
diameter of 102 mm.

(a)

(b)

Fig. 3 (a) Integrated 70 µm thick PZT/PZT composite film UT
Flexible ultrasonic transducers (UTs) consisting of ametal foil, a piezoelectric ceramic film and a top electrode areon a 12.7 mm thick Gr/Ep composite and (b) its ultrasonicperformance in time domain at 22°C.

also developed. Here, a 75µm thick SS, a 40 µm thick PZT/PZT
composite film and silver paste was chosen as metal foil,

4. Conclusions

piezoelectric film and top electrode materials, respectively. The

Integrated UTs have been developed for in-situ

SS foil serves both as substrate and bottom electrode. The

structural health monitoring, on-line and off-line NDT andflexibility is realized owing to the porosity of piezoelectric film

on-line real-time manufacturing process monitoring. Fabricationand the thinness of metal foil as shown in Fig.2a. This flexible

techniques focused on the use of handheld equipments to

UT was attached by a mechanical holder onto the steel pipe withperform sol-gel spray coating including using a torch or heat gun

outer diameter of 102 mm, inner diameter of 46 mm, and a length

to carry out drying and firing, poling and electrode fabrication inof 205 mm. A viscous oil ultrasonic couplant was placed

the production of these UTs. These UTs can be fabricated in

between the SS foil and the outer surface of the pipe. The

arrays and operated at desired NDT sites even of curved surfaces

pulse-echo mode was used. The steel pipe was heated up by a hot

at high temperatures. Metal and Gr/Ep composites substrates

plate. The ultrasonic performance in the time domains at 160°C

were used.

is shown in Fig.2b and L2, L4 and L6 are respectively the 1st, 2nd
and 3rd round-trip reflected echoes through the thin SS foil,

Acknowledgment

couplant and the wall thickness of the steel pipe. The measured

The authors are grateful to H. Hébert, Y. Ono, S.B. Nguyen and

SNR of L2 was 28 dB. Thus in-situ or off-line NDT using such

J.-F. Moisan for their technical assistance.

flexible UTs attached or glued to substrates are feasible andconvenient. In addition, the test results show that this particular

References

flexible UT can be bended into a curvature of a radius of 15 mm

1. T. Arakawa, K. Yoshikawa, S. Chiba, K. Muto, and Y. Atsuta

without damaging the ultrasonic performance. It is expected that

: Nondestr. Test. Eval., 7 (1992) 263.

further reduction of the thickness of metal foil and piezoelectric

2. H. Karasawa, M. Izumi, T. Suzuki, S. Nagai, M. Tamura, andcomposite film will enhance the flexibility.

S. Fujimori : Journal of Nuclear Science and Technology, 37

Gr/Ep composites are now increasingly used for large(2000) 769.

structures due to their high strength over weight ratio. In the

3. M. Kobayashi and C.-K. Jen : Smart Mater. Struct., 13 (2004)

manufacturer's recommended curing technique the highest

271.

temperature in the curing stage was only 176°C and the duration

4. D. Barrow, T.E. Petroff, R.P. Tandon, R.P. and M. Sayer : J.

was two hours. For this kind of substrate the PZT/PZT composite

Apply. Phys., 81 (1997) 876.

- 318 -

H-2

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 319-320
16-18 November, 2005

Evaluation of closed cracks by nonlinear ultrasonic phased array◎Yoshikazu Ohara, Ryouta Sasaki, Toshihiro Ogata, Tsuyoshi Mihara and Kazushi Yamanaka(Tohoku Univ.)

1. Introduction
Although closed cracks can propagate underapplied stress resulting in catastrophic accidents e.g.
in atomic power plants, it is difficult to detect themsince the intensity of scattering waves is much
lower than that from open cracks. Recently,
nonlinear ultrasound is expected to have a potentialof detecting and evaluating them, and subharmonic
wave [1,2] with half-frequency of input signal isparticularly useful because of its selectivity forcracks. Thus far, we have observed subharmonics
in closed cracks and developed a model toanalytically reproduce the observed waveform [3].
Also, we observed a peculiar wave at the tail part ofreceived wave and explained it as a free oscillationof cracks [4]. However, fast and reliable test was
not realized due to the lack of imaging capability.
In this study, we propose a novel nonlinear
ultrasonic phased array (PA) to image closed andopen cracks using digital signal processing (DSP)
for filtering at linear and nonlinear frequency. To
interpret the results and ascertain the mechanism ofsubharmonic resonance [5,6], we examine thefrequency dependence of subharmonic generationin closed cracks.
2. Nonlinear Ultrasonic Phased Array
As a fast and reliable method for the evaluationof closed cracks, we constructed nonlinear
ultrasonic PA with DSP, as shown in Fig. 1. Sincetransducers in conventional PA can not generate an
intense ultrasound required in nonlinear testing of
cracks, we developed a LiNbO3 single crystal
transmitter, which can endure against high voltage
input, with an appropriate wedge made ofpolyimide. It is driven by a short tone burst
amplified by a gated amplifier, whereas the PA with
32 elements at 5 MHz center frequency is utilized
as a receiver. Before the process of imaging, DSP
extracts linear or nonlinear components. For
example, closed cracks can be imaged by DSP
which passes only subharmonic or superharmoniccomponent. On the other hand, open cracks canbe imaged by DSP which passes only fundamental
components.
To realize such a measurement system, thecalculation of a time shift of waveforms arrivingfrom focal point is required, since the transmitterand receiver are separated contrary to conventional
PA. The formulation of the propagation time andtime delay has been already reported in Ref. [7].
Signal Generator

Personal
Computer

Gated Amplifier

Programmable
Phased Array

LiNbO3 single crystal

Phased array sensor(32 element)

Polyimide
52
52

85
85

22

3 10

4038
19
tt 30

170

Fig. 1. Nonlinear ultrasonic phased array.
3. Results and Discussions
We prepared well-defined fatigue cracks in analuminum alloy (A7075). The crack was extended
from a notch in a three-point bending fatigue test
under two conditions in Table I [8].
Table I. Maximum and minimum stress intensity factor
for fatigue test.
Kmin (kgf/mm3/2)
Kmax (kgf/mm3/2)
Sample A
17
2
Sample B
142

Figure 2 shows a linear image for sample A,
where scan area is from 14 to 50 mm (2 mm step)in depth and from 10 to 60° (2° step) in angle. It
is confirmed that the position of crack tip in theimage agreed with that of the actual crack tip.
Therefore, the validity of the developed ultrasonic
PA was verified.
Next, we examined the frequency dependence ofsubharmonic generation in closed crack of sample
B by sweeping the input frequency.

- 319 -

(a)

(b)

(c)

Fundamentals(Input 7.2 MHz)

Subharmonics(Input 7.2 MHz)

Subharmonics(Input 10.2MHz)

Absent5 mm

Fig. 2. Comparison of linear image obtained by
nonlinear ultrasonic PA with actual fatigue crack ofsample A.
A tone-burst of 3 cycles with peak-to-peakdisplacement amplitude of 20 nm was used. The
parameters of the band pass filter are shown in
Table II.
As shown in Figs. 3(b) and 3(c), a clear
difference between 7.2 and 10.2 MHz was observed,
whereas crack tip was not observed in linear imageof Fig. 3(a) probably because the crack is closed.
In Fig. 3(c), the subharmonic waves were generated
not only at the crack tip but also at a part below
crack tip. On the other hand, it was generated onlyat crack tip in Fig. 3(b). Thus, the frequency
dependence of subharmonic generation in closed
cracks was observed for the first time by applying
the nonlinear ultrasonic PA to the well-definedcrack. As a means of interpreting the frequency
dependence of subharmonic generation in closed
cracks, we proposed a new concept [6] as follows.
In medical ultrasonic imaging, it has beenclarified that the subharmonic generation at micro
bubble is maximized when the input frequency isclose to twice the resonant frequency of bubble [5].
Accordingly, in the case of subharmonic generationin closed cracks, the intensity will take a maximum
when the input frequency is close to twice theresonant frequency of closed crack.
In linear ultrasound, the frequency dependence of
scattered wave from a crack was examined by theboundary integral equation method and the resonantcondition was shown that crack length is close to
the odd multiple of a quarter Rayleigh wavelength[9]. In the case of closed cracks, although thesame resonant vibration as the above is not excited,the subharmonic vibration can be resonantly excited
if the input frequency is close to twice the resonant
frequency of the crack. On the basis of this concept,
the local variation in Figs. 3(b) and 3(c) may beinterpreted as the frequency dependence ofsubharmonic excitation of a part of the closed crackwith dimension comparable to the Rayleighwavelength in submillimeter to millimeter range.

NewfeatureFig. 3. Images of fatigued crack tips in sample B
obtained by nonlinear ultrasonic PA with DSP.
Table II.
Cutoff frequency, attenuation rate andstopband attenuation of selected band pass filters.
Frequency
CutoffAttenuation
Stopbandto extract
frequency
Rate
Attenuation
(MHz)
(MHz)
(dB/octave)
(dB)
7.2
6.2
8.2
281
261
60
3.6
2.6
4.6
170
121
60
5.1
4.1
6.1
217
177
60

4. Conclusions
We proposed a novel nonlinear ultrasonic phasedarray. We expect that this approach is significantly
useful in detecting and quantitatively evaluatingclosed cracks.
In our future works, we willquantitatively verify the model for subharmonic
generation.
References
1. A. Moussatov, V. Gusev and B. Castagnede：Phys.
Rev. Lett. 90-12 (2003) 124301-1.
2. I. Solodov, J. Wackerl, K. Pfleiderer and G.
Busse：Appl. Phys. Lett. 84 (2004) 5386.
3. K. Yamanaka, T. Mihara and T. Tsuji：Jpn. J. Appl.
Phys. 43-5B (2004) 3082.
4. R. Sasaki, T. Ogata, Y. Ohara, T. Mihara and K.
Yamanaka：Jpn. J. Appl. Phys. 43-5B (2004)
3082.
5. P. M. Shanker, P. D. Krishna and V. L.
Newhouse：J. Acoust. Soc. Am. 106 (1999) 2104.
6. K. Yamanaka, R. Sasaki, T. Ogata, Y. Ohara and
T. Mihara : Rev. Prog. QNDE (2006),
submitted.
7. Y. Ohara, T. Mihara and K. Yamanaka : The third
US-Jpn. Symp. in NDE, (2005) 229.
8. M. Akino, T. Mihara and K. Yamanaka：Rev.
Prog. QNDE 23B (2004) 1256.
9. D. E. Budreck and J. D. Achenbach : J. Appl.
Mech. 55 (1988) 405.

- 320 -

H-3

Proc. Symp. Ultrason. Electron., Vol. 26, (2005) pp. 321-322
16-18 November, 2005

On the Photoacoustic Nondestructive Instrumentation with aLine-Focus Laser Beam and a Planar Specimen Combination
Tsutomu Hoshimiya1, Mika Hatake-yama, Naoki Ohtaki and Haruo Endoh○(Tohoku Gakuin Univ., Fac. Eng.)
The specimen used was a pair of welded steel plates

1. Introduction
In the field of photoacoustic (PA) imaging, apoint-focus (PF) laser beam is dominantly used tomake thermal or acoustic image for inspection of thespecimens such as welded steel plates [1]. However, aline-focus (LF) laser beam is also useful to achievewith a thickness of 1.5 mm (maximum turn-on currentof welding: 8kA). A drill hole was made by drillingwith a diameter 1.5mmφ along interface surface of
the welded two plates from the edge of one side to thecenter of the specimen.

quick inspection of the solid specimen, as pointed by
the first author as a Japanese patent [2].
In this paper, the inspection of the welded steel
plates by rotating a LF laser beam around the center ofthe welded region was done. The PA signal
dependence on rotation angle is measured forinspection [3]. The feasibility to apply the LF laser
beam to achieve photoacoustic tomography (PAT) [4]is also described.
Fig. 1 Basic experimental setup.

2. Experimental apparatus
The PA imaging system is similar to that of the

3. Experimental results

Photoacoustic microscope (PAM). The PA
signal was detected by a highly sensitive condenser
microphone (Brewer & Kaejer 4166), and it wasprocessed with a lock-in amplifier, reconstructed as
amplitude and phase images, and displayed andrecorded by a computer. Amplitude and phase PA
images were stored as spread-sheet data, and those areavailable to be processed such as integration or noise
reduction easily.
A laser diode (LD) pumped YAG laser wasused as an optical excitation source with a LF beam
shape. Its beam was previously diffused with aconcave lens and then focus as a line image by acylindrical lens. Laser beam size was 8 mm in length
was 8 mm in length and 1 mm in width.

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e-mail:tpth@tjcc.tohoku-gakuin.ac.jp

The step of the rotation angle and the timeneeded for a measurement were one degree and sevenminutes, respectively. Laser power and modulation
frequency were 28mW and 8 Hz, respectively. The
thermal diffusion length was about 77μm such that itis enough for the thermal wave to reach the drilledhole made at the welded interface.
The time needed for inspection of the welded
specimen by the use of a point-focus laser beam withthe resolution of 50 x 50 pixels was about 30 minutes
so that a simple inspection using a LF laser beamshortened measurement time in about four times.
The dependence of the obtained PA signal on
the rotation angle was shown in Fig. 2. In Fig. 2 (a),PA signal dependence of the specimen without an
artificial undersurface defect is shown, whereas that- 321 -

2005-206177（2004.3）.

corresponding to the specimen with undersurface
defect is shown in Fig. 2 (b). It shows the PA signal[3] T. Hoshimiya and M. Suzuki : Abstract of 2005

maximum around the angle of 90 degrees, at whichIEEE

Ultrasonics

the overlap of the LF laser beam and a drill hole

Rotterdam).

Symposium,(2005.

9,[4] X. Wang, Y. Pang, G.. Ku, G.. Stoica, and L. V.

maximizes.

Wang: Opt. Lett., 28, (2003) 1739.

4. Discussion
The PA signal generated in the solid specimens
in most case is proportional to the absorption
coefficient of the specimen α and laser intensity I0 ,andinversely

frequency f.

proportional

to

the

modulation

If we irradiate plane solid specimen

with a LF laser beam, the generated PA signal is
proportional to the summation of the absorption
coefficient of the irradiated area so that thiscombinations useful to realize two-dimensional (2D)CT PA imaging or 2D-PAT as pointed in ref [2].
The feasibility to use a LF laser beam to realize
non-contact photothermal radiometry, such as active
thermography, is also pointed out in the patent [2].

5. Conclusion
Simple nondestructive inspection (NDI) of thewelded specimen with an artificially made wasachieved by rotating a LF laser beam. Combination of
LF laser beam with CT technique is the next step
purpose of the current study.

Fig. 2PA signal dependence on rotation

angle: (a) specimen without undersurface
defect(upper), (b) specimen with undersurface

Acknowledgement
The first author (TH) is grateful to Suzuki
Motor Company for the financial support and sample
preparation.

References[1]K. Miyamoto, M. Suzuki, T. San-nomiya, T.
Hoshimiya and N. Chubachi: Jpn. J. Appl. Phys., 43,3107 (2004).
[2]T. Hoshimiya: Japanese patent, submitted No.

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defect (lower).