Nondestructive Characterization of Materials X

N

~

X

m O

o

m o

.,

6

"~ r=

5

r

4

~2

t 9No.2 Crack | A

0"'o

3

Densily(/cm) O No.4 Crack Density(/cm.)

ip

o o.o

. . . .

>. m

u

0

0.2

0.4

0.6 Strain

No.2

0.8

I= PredictionOfNO.2 Prediction of No.4 .... i 1.2

(%)

9[(0" /Ad/Ad/90 ~ 2)s]

No.4 : [(0 ~ /Ad/2% Pre-Strained S M N A d / 9 0 ~ 2)s]

Fig. 9. Transverse crack density as a function of strain at room temperature for without (No. 2) and with (No. 4) embedded 2% pre-strained SMA films Shape memory alloy (SMA) films are embedded and used to suppress and control microscopic damages in CFRP laminates such as transverse cracks and delamination [14]. Improvement of interlaminar shear strength (ILSS) between SMA films and CFRP laminas has been investigated using spattering, sol-gel, ion-plating and anodic-oxidation. A high ILSS was obtained similar to that of CFRP laminates alone. SMA films were stretched into the plastic

16

N. Takeda

deformation region. Then, they were embedded in CFRP laminates with the deformation kept by the fixture jig during the fabrication in order to introduce the shrinking stress in 90-degree plies. Such shrinking stresses were found to suppress the evolution of transverse cracks in cross-ply laminates.

Quantitative Evaluation of Electric Properties of CFGFRP Hybrid Composites as a Maximum Strain Memory Sensors- Toray Failure of low-elongation carbon fibers in CFGFRP

hybrid

composites

under

tensile

loading causes the increase in electrical resistance and can be used to detect the maximum strain applied to composites after unloading. A systematic study was conducted to evaluate the relation between fiber failures and Fig. 10. Fiber failure in CFGFRP

electrical resistance quantitatively [15].

Figure 10 shows the fiber failures observed in CFGFRP hybrid composites in tension. Based on such quantitative observations, a Monte Carlo simulation was conducted to predict the change in electrical resistance due to tensile strain and fiber failures. As shown in Fig. 11, a good agreement was obtained between the experimental results and the prediction.

Experiments (Vf)carbon

Simulations

= 1 5 % , (Vf)glass = 4 5 %

(Vf)carbon

c,.,oos

, oic,. ,oos t

s,

15

0

~ '

io ,

Stress ~

]

.-~

0.1

/

V O r"

!,oof |

,of /

=20% , (Vf)glass = 4 0 %

/ J

1

= GmP '

"" OFRP(T700 ' ~. ) . . . .

2

strain (%)

J=

~ 3

0

4

0

0

1

2

3

strain (%)

4

U..

5

Fig. 11. Comparison between experimental results and prediction

Self-Diagnosis Function of Electrically-Conductive FRP Containing Carbon Particles - JFCC Carbon particles or flakes were dispersed into the epoxy matrix to introduce the high electrical

Structural Health Monitoring in Japan

17

conductivity in glass fiber unidirectional or textile composites (CPGFRP) [16]. Compared with CFGFRP hybrid composites, higher sensitivity could be obtained. Moreover, a linear resistance change due to tensile strain was achieved in a wider strain range, as shown in Fig. 12. The residual resistance after unloading increased with increasing applied maximum strain. This composite has an appropriate self-diagnosis function as a low-cost sensor to be embedded in concrete infrastructures. 400[ ....

, .........

~ooL~ ~.,.o,,j / / I " t " ...... ~-,oo F // ~,,o~ //

50

2

,o

1.5

'_1 . . . .

1

//

,o~

,o~~

,

'oot / / o

o.s

,o

I

Strain / %

1.8

2

~-

"

"

,

40 30

Re- 1233I'I

E5

1

20

0.50~

0$10

o

i 20

4o

i _. 6o

i 80

1oo

time / min

(a) Monotonic quasi-static loading test (b) Loading-unloading cycle test Fig. 12. Electrical resistance change in CPGFRP composites

Impact Damage Monitoring of Composite Structures Using Integrated Acoustic Emission (AE) Sensor Network- Aerospatiale Matra An integrated AE network system is being developed for practical industrial use in aircraft structures [17]. A Learning by Experience approach was developed to assess the AE behavior of a composite structure. The principle is to generate artificial AE events due to impact loading at several arbitrary points on the test specimen (Fig. 13). Recording the acquired waveform parameters enables the system to learn the structure. Then, the system can receive real AE events upon impact in real-time in order to estimate the impact amplitude and to make an accurate localization of the source of impact (Fig. 14).

Fig. 13. Specimen and AE event generator

Fig. 14. Interpolation and localization of impact

18

N. Takeda

Due to lack of space, only titles are cited for three other research themes :

(1) Integrated Global and Local Strain Measurement Using Distributed BOTDR (Brillouin Optical Time Domain Reflectmetry) and FBG Sensors- Mitsubishi Heavy Industries : Improvement of spatial resolution, temperature compensation and dynamic strain measurement are conducted. Integrated BOTDR and FBG strain measurement systems are being made for aerospace structures [ 18]. (2) Damage Detection of Transparent Composites Using Light Transmission and Reflection

Measurement for High-Speed (Magrev) Train- Hitachi : Damages such as transverse cracks, delamination and fiber failures in semi-transparent alumina fiber reinforced epoxy composites are detected using light transmission and reflection technique under severe electro-magnetic and low-temperature conditions for load-supporting structures in Magrev trains [ 19]. (3) Development of FBG Sensor Elements and Real-Time Monitoring Systems for Large-

Scale Infrastructures- Shimizu 9A multiplex FBG sensor element network system is established for structural health monitoring of infrastructures. In particular, a real-time monitoring system is installed to record strains and deformations in hysterisis dampers for urban earthquake mitigation [20].

REFERENCES

1. Takeda, N. and Ogihara, S. (1994) Comp. Sci. Tech. 52, 183. 2. Takeda, N. and Ogihara, S. (1994) Comp. Sci. Tech. 52, 309. 3. Ogihara, S. and Takeda, N. (1995) Comp. Sci. Tech. 54, 395. 4. Takeda, N., Ogihara, S. and Kobayashi, A. (1995) Composites 26, 859. 5. Takeda, N., Kosaka, T. and Okabe, Y. (1996) Sci. Engr. Comp.Mater. 5, 169. 6. Takeda, N., Niizuma, H., Ogihara, S. and Kobayashi, A. (1997) Exp. Mech. 37, 182. 7. Takeda, N., Ogihara, S., Suzuki, S. and Kobayashi, A. (1998) J. Comp. Mater. 32, 83. 8. Takeda, N., Ogihara, S., Nakata, N. and Kobayashi, A. (1998) Comp. Inter. 5, 305. 9. Takeda, N. and Ogihara, S. (1998) Composites: PartA, 29A, 1545. 10. Okabe, Y., Yashiro, S., Kosaka, T. and Takeda, N. (2000) Proc. SPIE, Vol. 3893, in press. 11. Satori, K., Ikeda, Y., Kurosawa, Y., Hongo, A. and Takeda, N. (2000) Proc. SPIE, Vol. 3893, in press. 12. Tsutsui, H., Sanda, T., Okabe, Y. and Takeda, N. (2000) Proc. SPIE, Vol. 3893, in press. 13. Kabashima, S., Ozaki, T. and Takeda, N. (2000) Proc. SPIE, Vol. 3893, in press. 14. Ogisu, T., Nomura, M., Andou, N., Takaki, J., Song, D.-Y. and Takeda, N. (2000) Proc. SPIE, Vol. 3893, in press. 15. Song, D.-Y., Takeda, N., Kitano, A., and Yoshioka, K. (2000) Proc. SPIE, Vol. 3893, in press. 16. Okuhara, Y., Shin, S.-G., Matsubara, H., Yanagida, H. and Takeda, N. (2000) Proc. SPIE, Vol. 3893, in press. 17. Saniger, J. and Reithler, L., (1999) Proc. 1st Symp. Smart Mater., pp. 119-122. 18. Yamaura, T., Inoue, Y., Kino, H. and Nagai, K. (1999) Proc. 2nd Int. Workshop on Structural Health Monitoring, pp. 533-542. 19. Aoyama, H., Tanaka, K., Watanabe, H. and Takeda, N. (1999) Proc. 6th Japan Int. SAMPE Symp., pp. 967-970. 20. Mita, A. (1999) Proc. 2nd Int. Workshop on Structural Health Monitoring, pp. 56-67.

Nondestructlve t_.haractenzatlonot Materials A Green et al. (Eds.) Published by ElsevierScience Ltd., 2001

19

SOUND VELOCITIES AND MICROSTRUCTURE

HASSEL LEDBETTER Materials. Science and Engineering Laboratory, National Institute of Standards and Technology, Boulder, Colorado 80303, USA

ABSTRACT In this brief review, I emphasize and illustrate the utility of sound velocity to study a solid's microstructure. By microstructure I mean departures from a perfect lattice, thus a wide and diverse spectrum of defects, anything from cracks to phonons (thermal vibrations). I consider several aspects of sound velocities: real and imaginary parts, dispersion, scattering, absorption (dissipation). Examples include materials ranging from composites to steels. I emphasize the importance of combining measurements with modeling-theory. KEYWORDS absorption, attenuation, composites, cracks, Debye characteristic temperature, dispersion, elastic constants, hardness, internal friction, measurement-theory, scattering, sound velocities, steels, texture, voids, waves.

INTRODUCTION By microstructure, I mean any departure from a perfect lattice. Examples considered here include the following: (1) crack, (2) second phase, (3) dislocations, (4) internal interface, (5) voids. A complete list of lattice defects would be much longer. By sound velocity v, I mean the speed with which a mechanical (acoustical) wave moves through a solid, a number that varies from 0.22 cm/laS in lead to 1.81 in diamond, with values for other typical elements being 0.37 for aluminum, 0.48 for copper, and 0.59 for iron. The sound velocities relate simply to the elastic stiffnesses C through the mass density p: C = pv 2 9

(1)

Thus, nearly always, we can interchange sound velocity and elastic stiffness. C is actually a fourth-order tensor Cijkt, and for the least-symmetrical case (triclinic) there are twenty-one independent C..qkl" The .most symmetrical case (isotropic) exhibits two independent C..tj (Voigt contracted notation). Since Stokes in 1845 [ 1], we know that the most logical pair is the bulk modulus B and the shear modulus G:

20

H. L e d b e t t e r

2

B = p(v I

4 v 2)

- .~

,

2

G = ,o v t

o

(2)

(3)

Here, subscripts l and t denote longitudinal and transverse. Sound velocities are solutions of the familiar wave equation V2u = ~)2u = ~ 1 ~92u ~)x 2

(4)

v 2 ~)t 2

Here, u denotes displacement, t time, and x the propagation direction. If we invoke Hooke's law connecting stress (y to strain e (~ = Ce), then we obtain

~)2~3

1 ~)2E

~)x 2

- ~ ~)t 2

(5)

Equation (4) is essentially the same wave equation appearing in quantum mechanics and in a wide range of solid-state phenomena [2]. Often, we get solutions to Eq. (5) by solving the Christoffel equations: det (CijklXjXk - P v 2 ~)il) = 0 .

(6)

Here ~)il denotes the Kronecker symbol, and we obtain three eigenvalues pv 2, one quasilongitudinal and two quasitransverse.

SOUND-VELOCITY CONNECTIONS WITH OTHER PROPERTIES For the cubic elements, Fig. 1 shows a diagram of shear velocity v t versus the Debye characteristic temperature O. This astonishing correlation shows that v t connects strongly with a wide variety of solid-state phenomena connected with O. Ledbetter [3] listed about fifty, ranging from atomic-vibration amplitude through melting temperature to zero-point energy. Figure 2 connects G = pv 2 with a practical plastic-deformation property: hardness. We know that hardness, in turn, connects with such properties as yield strength and ultimate strength.

IRON-COPPER STEELS" HARDNESS AND DISLOCATION DENSITY v 2) for a series of Fe-Cu Figure 3 [4] shows the Young moduli E = p v 2 [ v 2 - ( 4 / 3 ) v 2 ] / ( V l steels where hardness increases as coherent b.c.c, copper precipitates develop during heating. Although the overall change is small, about one-half percent, it is easily measured.

For the same steels, Fig. 4 shows Young modulus versus annealing time and the expected underaged, optimum-aged, and overaged regions. The figure also shows the associated internal

21

Sound Velocities and Microstructure

friction Q-1 measured by resonance-peak widths. Q-1 decreases continuously with annealing time, reflecting, probably, increased dislocation pinning.

. . . . . . . .

I

. . . . . . . .

I

1000

'

A

O.

=L

E

(.9

>

Cubicelementso / a9~ 100" i

i

i

i

i

U) :3 :3

i n

(~ 0.1

O 4)

10"

0

E

>

O ~

(3

L

I._

m q)

r

l-

.E

J:

u) 0.01 10

................... 100

0.1 104

1000

Debye 19 (K)

'!

,i

10-~ 10.2 10-1 100

Hardness

Fig. 1. For cubic elements, relationship between sound velocity v t and Debye charac-

101

102

(GPa)

Fig. 2. Relationship between shear modulus G = Pvt:2 and physical hardness.

teristic temperature | 2

212

.

.

.

.

i

.

.

.

0

9

,

5

~

2

5

.

Fe-Cu-CSteels

0=

fI ;~

D.

(9 U)

210

~

20

"O O

E 208 G

=~

c O ~"

~,,

~_.

15

208

2o6

5O

i

i

I

i

I

i

i

i

55 Hardness (R A)

,

60

Fig. 3. Young-modulus/hardness relationship for some Fe-Cu-C b.c.c, steels.

~ 207,5 0

1 50

0 100

150

200

Annealing time (h)

Fig. 4. Companion to Fig. 3. Dependence of Young modulus E and internal friction Q-1 on annealing time.

H. Ledbetter

22

1i ....

-o.ool

i ....

i ....

$5I 't1"0

"',,

0.9

7W

"'

/

0.8 7'~

=24

,4 t~

r~

3

e-

.2 2 m t}

o

1

50

52 54 56 58 60 Hardness (Rockwell A)

Fig. 6. Companion to Fig. 3. Dislocation density A obtained from internal-friction decay and from transmission-electron microscopy.

For these same steels, we now consider the dislocation density estimated in an unusual way, shown in Figs. 5 and 6 from Ogi and Ledbetter [5]. In Fig. 5, we see the attenuation change Ao~ compared with the Granato-Hikata-Lticke theory and the internal-friction Q-1 change (measured by torsion pendulum) compared with the Cottrell-Bilby-Harper theory, which assumes carbon atoms diffusing to dislocation sinks. Their relationship, of course, includes the dislocation density A, shown in Fig. 6, together with dislocation densities obtained by transmissionelectron microscopy. Considering the extreme difficulty of determining dislocation densities, the agreement of the two approaches is encouraging. Also, our results confirm Friedel's prediction that A varies as hardness squared.

A STEEL: CRACK SIZE Figures 7 and 8 show some results from Ledbetter et al. [6] who studied an artificial crack in a stainless-steel block using acoustic-resonance spectroscopy. Figure 7 shows a typical resonance spectrum. And Fig. 8 compares the measured resonance-frequency changes with those calculated by a finite-element method. Agreement is good. (The frequency shifts are enormous, up to one-hundred percent!) Thus, there arises the following question. Can we use the macroscopic-vibration-frequency spectrum to deduce various crack aspects: size, shape, location, orientation? Such questions remain to be studied.

A LAMINATED COMPOSITE: DISPERSION Figures 9 and 10 from a study by Datta, Shah, and Ledbetter [7] show some theoretical and measured results for a B-Al-matrix composite. For longitudinal waves, Fig. 9 shows the

23

Sound Velocities and Microstructure

dispersion figure calculated by a stiffness method using Floquet's theory and considering anisotropic laminae. The laminae plane is x-z, x = fiber direction, y = normal to laminae. The model predicts minimum dispersion along x, maximum along y, and intermediate along an x - y 45 ~ direction. Figure 10 confirms this behavior. Thus, even ill-defined laminae cause measurable dispersion. And, we could use the amount of dispersion to measure laminate quality.

0.8

>

0.6-

ug a

::~ 0.4I-,, __J a,,

0.0

50

Ii

'

'

,,

'

,

l

1

100

'

'

'

'

150

'

'

,

Ill '

'

I

200

'

'

'

'

,

r

,

,

-~

250

FREQUENCY

300

J

350

4(

(kHz)

Fig. 7. Resonance spectrum of a 5• 10• 14-mm stainless-steel parallelepiped containing a single artificial 0.5-cm-deep crack.

~.j

100 ~

E

80.

==

60.

401

'

'

'

'

I

20

'

'

'

'

I

40

'

"

'

"

I

60

'

'

'

'

I

80

'

'

'

% Change" Measurement

'

100

Fig. 8. Companion to Fig. 7. Predicted-versus-measured resonance-frequency shifts caused by a 0.5-cm-deep crack.

H. Ledbetter

24 0.3

'

>:,

o t-

o

i

.89

1/1

or=0 0.2

E

[7"

.87

T1 *

-

B/AI [110]

~ ~

>.-

U_

r

"O

._N 0.1 m t'cl

o Z

o

0.05

1

>

Fig. 9. Theoretical dispersion relationships for longitudinal waves in a B-A1 laminate. (x = 0 means wave-propagation vector in laminae planes and along fibers, (x = 90 means perpendicular to laminae planes and to fibers, (x = 45 means at 45 ~ to planes and to fibers.

~~.~[o~o]

"

0

0

,

!

0.1

' 0'.2 0.3 Void Content

o11

" ' ~ 00.4

Fig. 11. Effect of void content on Young moduli in compacts of sintered titanium spheres. Also shown are internal-friction changes. E mode measured by Marx oscillator. G mode measured by acousticresonance spectroscopy.

o'.2

Particle Volume Fraction

o.a

Fig. 12. Effect of volume fraction and particle shape on attenuation caused by wave scattering as predicted by Ledbetter-Datta theory based on ensemble average of scattered plane waves.

SCATTERING AND ABSORPTION The intensity decrease of a mechanical wave traveling along x can be written I - I0 e-2ctr .

(7)

The attenuation coefficient c~ contains two principal parts: = ~s + ~a "

(8)

Here subscript s denotes scattering, a denotes absorption. Dissipation, logarithmic decrement, quality factor, and internal friction are often used instead of absorption because they relate directly"

26

H. Ledbetter

Q-1 = tan7 = - 5 = AU _- M2 =__'/2 . r~ 2rtU M1 J1

(9)

Here Q denotes quality factor, Q-1 internal friction, 7 phase lag of strain behind stress, ~i logarithmic decrement, U internal energy, AU energy loss during a stress-strain cycle, M 1 and M 2 the real and imaginary elastic stiffnesses, J1 and J2 the real and imaginary elastic compliances. Also, ~a may arise from many mechanisms; thus ~a = ~ilT'ai" In some cases, one can separate the various tT,ai by varying frequency, temperature, stress, or another variable.

CONCLUSIONS Sound velocities provide a valuable probe of microstructure. Their advantages include the following: (1) polarization, (2) coupling with many defects (some nonmechanical), (3) ability to be measured accurately (at least to one part in a thousand), (4) ability to be treated accurately with models and theories, (5) real and imaginary parts, the latter coupling very strongly with lattice defects. Extension of this brief review would include the following topics: (1) measurement methods, (2) thin films, (3) many more examples of other defects, (4) surface waves.

ACKNOWLEDGMENTS Several people contributed to the above studies, especially S. Kim (NIST), S. Datta (U. Colorado), M. Dunn (U. Colorado), P. Heyliger (Colorado State U.) and H. Ogi (Osaka U.).

REFERENCES 1. Stokes, G. (1845). Trans. Cambr. Philos. Soc. 8, 287. 2. Morse, P., and Feshbach, H. (1953). Methods of Theoretical Physics. McGraw-Hill, New York. 3. Ledbetter, H. (1983). In: Materials at Low Temperatures, pp. 1-45, Amer. Soc. Metals, Metals Park, Ohio. 4. Ledbetter, H., and Kim, S. (1990). Unpublished research, National Institute of Standards and Technology, Boulder, Colorado. 5. Ogi, H., and Ledbetter, H. (2001). Mater. Metall. Trans. forthcoming. 6. Ledbetter, H., Heyliger, P., Pei, K.-C., Kim, S., and Fortunko, C. (1995). Rev. Prog. Quant. Nondest. Eval. 14, 2019. 7. Datta, S., Shah, A., and Ledbetter, H. (1983). In: Mechanics of Composite Materials, Recent Advances, pp. 207-215, Pergamon, New York. 8. Ledbetter, H., Dunn, M., Kim, S., and Fields, R. (1995). Rev. Prog. Quant. Nondest. EvaI. 14, 1633. 9. Ledbetter, H., and Datta, S. (1986). J. Acoust. Soc. Amer. 79, 239.

Nondestructive Characterizationof MaterialsX Green et al. (Eds) Crown Copyright 9 2001. Publishedby ElsevierScienceLtd. All rightsreserved.

27

RECENT DEVELOPMENTS OF THE LASER ULTRASONIC TECHNOLOGY AND OF ITS APPLICATIONS J.-P. MONCHALIN, A. BLOUIN, M. CHOQUET, D. LI~VESQUE, A. MOREAU Industrial Materials Institute National Research Council of Canada 75 de Mortagne Blvd Boucherville, Quebec, J4B 6Y4, Canada

ABSTRACT Laser-ultrasonics is an advanced ultrasonic technique based on the generation and detection of ultrasound with lasers. In this paper we present recent developments performed at the Industrial Materials Institute of the National Research Council of Canada, which includes progress made in the basic technology and applications. We discuss the recent developments on the use of photorefractive crystals in self-adapting interferometers for the detection of ultrasound and compare the performance obtained with the more established confocal FabryPerot demodulator. We present an approach based on the Synthetic Aperture Focusing Technique for enhancing the detection of small defects. Several applications that have reached various degrees of maturity are outlined, including the inspection of polymer-matrix composites, the detection of corrosion in the lap joints of aging aircraft, the characterization of the microstructure of steel and the monitoring of its evolution with temperature, the determination of the in-plane stiffness and the applied tension of a paper web.

KEYWORDS Laser-ultrasonics, laser-ultrasound, laser-based ultrasound, laser-based ultrasonics, polymer composites, ultrasonics, ultrasonic nondestructive testing, Synthetic Aperture, Synthetic Aperture Focusing Technique, SAFT, corrosion detection, paper, steel, composites, materials characterization, process monitoring, microstructure, grain size

INTRODUCTION Laser-ultrasonics, which is an advanced ultrasonic technique based on the generation and detection of ultrasound with lasers [1], has evolved during last decade from essentially a laboratory curiosity to a well-recognized industrial inspection technique. The technique, although at the present time not widespread, is however finding more and more applications in a great variety of industrial fields [2,3] and the interest is growing exponentially [4]. These applications are at various stages of development ranging from laboratory validation to real industrial use. We will first recall the principles of the technique.

28

J.-P. Monchalin et al.

Laser-ultrasonics involves essentially four steps. First a high power laser pulse generates ultrasound inside the material (or at its surface) by a thermoelastic mechanism or material ablation. Thermoelastic generation is often preferred since it does not affect the surface, but in cases where light penetration is small (e.g. metals), generation is weak, especially for normally propagating longitudinal waves, so a vaporization or ablation mechanism has to be used. Such a mechanism is of no consequence on hot products, the removed material being the oxide layer that is usually rapidly restored. In the cases where surface marks are of concern, damage is minimized by using a laser (generally UV) that vaporizes contaminants (such as oil) on the surface. In the cases where laser light penetrates significantly below the surface (cases of polymers and paints), an infrared laser is used and the generation of normally propagating longitudinal waves is strong. The second step involves illuminating the surface with a second laser beam. This laser could be cw or pulsed. If pulsed, its pulse length should be sufficient to capture the various echoes, so in practice pulse duration of at least 10 ps is usually required. The third step consists in collecting the light from the detection laser scattered by the surface. This light carries the ultrasonic information as a phase or frequency modulation. The fourth step involves the demodulation of this phase signal by using an optical interferometer [5]. The most useful interferometer for this purpose is the confocal Fabry-Perot, which could be used in transmission or reflection [6,7]. This system provides automatically the quasi-adaptation of the interfering wavefronts affected by speckles, which means that the system has a large 6tendue or throughput. In transmission, the system works essentially as a light filter. In reflection, at frequencies above the interferometer bandwidth, it operates essentially as a two-wave interferometer. The broad 6tendue of the confocal Fabry-Perot receiver allows coupling the scattered light through large core multi-mode fibers. A sketch of a typical laser-ultrasonic system is shown in Fig.1. In this sketch the generation laser is not fiber coupled, which would be the case of an excimer laser or a far-infrared laser, such as the CO2 laser. Other short pulse lasers in the visible or near infrared range (such as the Q-switch Nd-YAG ) could be fiber coupled up to some maximum power. The industrial use of laser-ultrasonics is motivated by essentially two characteristics of the technique. First, the technique works remotely (from centimeters to more than one meter), so there is no contact with the probed part and hot products can be tested at any temperature. Second, generation and detection are performed off the surface of the tested part, the generated wave being essentially normal to the surface (in most cases of practical interest), so contoured and complex shapes can be readily inspected without any need of robotics to provide orientation of a transducer. This feature is illustrated by the sketch shown in Fig.2, which also shows how an image would be obtained by optical scanning. This feature is at the basis of the application to the inspection of polymer-matrix composite materials, which has been commercialized and is being used in production. We present below in some details several aspects of the basic technology which have been particularly developed at IMI. First we summarize the various features of the confocal FabryPerot receiver and the two-wave-mixing photorefractive demodulator developed in collaboration with the Laboratoire Charles Fabry of the Institut d'Optique Th6orique et Appliqu6e and Action Aquitaine de recherche en apesanteur, France. We then present an approach that allows efficient detection of small defects and is based on numerical processing with the Synthetic Aperture Focusing Technique. We then discuss various applications of laser-ultrasonics pursued actively at IMI: detection of flaws in polymer-matrix composites,

Recent Developments in Laser Ultrasonics

29

detection of corrosion in aging aircraft, characterization of steel and monitoring of steel transformations, characterization of a paper web.

Fig.1. Schematic of a laser-ultrasonic pulse-echo system based on the Fabry-Perot interferometer.

Fig. 2. Principle of laser-ultrasonic imaging

OPTICAL DETECTION OF ULTRASOUND WITH THE DEMODULATOR AND THE CONFOCAL FABRY-PEROT

PHOTOREFRACTIVE

Several parameters are used to characterize an interferometric demodulator and are needed to evaluate its use for a particular application: sensitivity, response time, 6tendue and insensitivity to amplitude modulation. Among these, sensitivity or the limit of detection is by far the most important parameter. It is conveniently measured by reference to the ultimate detection limit that can be obtained by interferometric detection. This limit is obtained with the basic scheme for homodyne detection, which consists in mixing the phase-modulated signal beam with a strong reference beam or local oscillator onto a photodetector. The ultimate

30

J.-P. Monchalin et al.

detection limit for a bandwidth of 1 Hz and an incident power of the signal beam of 1 W is ~limit=(~/2/~) (hv/2rl) 1/2 , where h is the Planck constant, )~ and v are the wavelength and the optical frequency, respectively and r I is the quantum efficiency of the photodetector used [8]. For example, for 1.06 ~tm wavelength and for a quantum efficiency of 1"1= 0.3, the ultimate detection limit is 5 x 10-8 nm (W Hz) 1/2. When working with a speckled signal beam, the sensitivity is reduced well below the ultimate limit unless the interferometer also acts as a signal-to-reference wavefront adapter or only one speckle is collected.

Sensitivity of various Fabry-Perot configurations The confocal Fabry-Perot receiver is a simple optical resonator consisting of two concave mirrors that can be used under different versions, which provide different detection bandwidths and sensitivities [7]. Examples of sensitivity spectra for these various configurations (referenced to the ultimate detection limit) are shown in Fig. 3. A first configuration is obtained with mirrors of identical reflectivities, the detector being located either on the transmission side or the reflection side. The use in reflection provides high sensitivity at high ultrasonic frequencies (higher than the cavity bandwidth) except at a few rejection bands. When high sensitivity is only required in this range of frequencies, the back mirror could be made totally reflecting. This leads to improved sensitivity, as can be seen in Fig. 3. This can be explained by the fact that there are four output beams exiting a confocal cavity, two on the transmission side and another two on the reflection side, giving four independent output ports (the incoming beam has speckle and the wavefronts are not adapted between these ports). There is multiple interference on each port and each port gives an output ultrasonic signal. The signal observed in transmission or reflection is the sum of the signals collected over the two transmission or reflection ports. In the case of a totally reflecting back mirror, there are only two output ports on the reflection side, thus increasing the intensity of interfering terms, since the total intensity is shared between two ports instead of four, and leading to higher sensitivity. A further improvement can be obtained with a configuration with only a single output port on the reflection side. This can be realized with a totally reflecting back mirror and a front mirror which is totally reflecting over half of its surface, as in the original confocal Fabry-Perot design [9]. This configuration which will be designated "Fabry-Perot type Connes" realized an ideal homodyne demodulator, but since half the incoming light is discarded its detection limit (at high frequencies) is ~/2 the ultimate limit. The combination of two systems will provide a limit equal to the ultimate limit with the penalty of higher complexity. It should be also noted that the transmission scheme could operate with unpolarized light whereas the use in reflection requires polarizing optics for optimum operation. Therefore often in practice, especially if a large core multimode fiber is used to transmit light to the Fabry-Perot, the transmission configuration gains a sensitivity factor of about ~/2 with respect to the others. If the range of frequencies of interest is between 1 and 10 MHz, which is often the case in nondestructive testing, this configuration will be the one usually selected with a proper choice of mirror reflectivity and cavity length to give adequate 6tendue and frequency response. The main weakness of the Fabry-Perot demodulators, as is obvious from Fig.3, is its lack of sensitivity at low ultrasonic frequencies (below 2 MHz), which is circumvented by the devices based on two-wave mixing in photorefractive materials.

Recent Developments in Laser Ultrasonics

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Sensitivity of the two-wave mixing photorefractive interferometer In the two-wave mixing approach, wavefront adaptation is performed actively, by opposition to the confocal Fabry-Perot in which adaptation is performed by passive or linear optical components [10-12]. The technique used is also known as real-time holography. This active wavefront adaptation eliminates the need of an external stabilization device against thermal drift or ambient vibrations, required for the confocal Fabry-Perot. The basic setup of the twowave mixing interferometer is sketched in Fig.4. A signal beam which acquires phase shift and speckle after reflection on a surface in motion, is mixed in a photorefractive crystal with a pump plane wave to produce a speckle adapted reference wave that propagates in the same direction as the transmitted signal wave and interferes with it. In Fig. 5, the sensitivities of the TWM interferometer operated with different semiconductor photorefractive crystals are compared to the confocal Fabry-Perot used in transmission ( R= 85 % , 1 m length). It is shown that the sensitivity of the two-wave mixing device is comparable to the maximum sensitivity of the CFP. The best sensitivity is obtained with a CdTe:V crystal due to its higher electro-optic constant. This best sensitivity is about 3 times less than the ultimate sensitivity, and can be improved by about a factor of 2 with a better tailored CdTe:V crystal. The photorefractive device is seen to be more sensitive than the CFP for ultrasonic frequency below 1 MHz. The photorefractive interferometer also has the advantages of a fiat frequency response up to at least a few GHz without the periodic high frequency notches of the Fabry-Perot.

32

J.-P. Monchalin et al.

S(x,O Fig. 4: Basic setup of the two-wave mixing interferometer; Signal beam (S), Pump beam (P), Reference beam or local oscillator (LO).

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Etendue o f the two devices

Having a large 6tendue is very desirable since it allows a large collection solid angle and, consequently, improves sensitivity. A large 6tendue allows also detection over a sizeable spot on the surface, which is required for some applications and eliminates also the need of sharp focusing onto the surface. Large &endue also permits coupling of the collected light into a large-core multimode optical fiber. Hence, the 6tendue of the interferometer should at least be equal to the 6tendue of such optical fibers which is typically 0.4 m m 2 s r (numerical aperture of 0.39 and core diameter of lmm) to coherently process all the light collected by the fiber. The large 6tendue of the confocal Fabry-Perot follows from the confocal nature of the cavity: a ray entering the cavity travels along the same path after multiple reflections upon the mirrors. For example, a meter long cavity with mirrors with 85% reflectivityprovides an 6tendue exceeding that of the collecting optical fiber mentioned above (0.4 mm ~ sr). In a two-wave

Recent Developments in Laser Ultrasonics

33

mixing scheme, the 6tendue is fixed by the size of the crystal and the angle between the signal and pump beams and is easily made larger than the 6tendue of the fiber mentioned above.

Response times of the two-devices When the inspected part is stationary the only requirement for the response time z is to be sufficiently short to provide adaptation to ambient vibrations, which means that a low frequency cut-off of about fc ~ 1 kHz corresponding to a response time zc=l/(2rffc) ~ 160 ~ts should be in practice quite sufficient for most environments. When the part is moving a much more rapid response is required. In the case of the confocal Fabry-Perot, adaptation is performed with light and xe can be roughly evaluated as equal to (4 R/c) F, where R is the mirror radii or cavity length, c is the speed of light and F is the finesse. For a one-meter long Fabry-Perot with a finesse of 10 or a 50cm long with a finesse of 20, this gives x ~ 130 ns. In the case of the photorefractive demodulator, the response time is given by the time of creation and washout of the photorefractive grating. It is therefore advantageous to use optical materials with high photoconductivity, such as semiconductors (GaAs, InP, CdTe) and to use strong pumping (which in turn requires a high power detection laser). Response times as short as 400 ns were obtained at 1.06 lam with a power density of about 3 kW/cm / in a GaAs crystal without an applied electric field. Since the response time scales as 1/power density, it is therefore possible to obtain a response time as short as the one typically obtained with a confocal Fabry-Perot, when a high power detection laser is used. The sensitivity is however lower with GaAs than with a confocal Fabry-Perot (except at low ultrasonic frequencies). The application of an electric field increases sensitivity (see Fig. 5), but with the penalty of a longer time constant. With an InP:Fe crystal under a field of 4.9kV/cm and a power density of about 100W/cm 2, a time constant of about 2 las is obtained [ 12]. Increasing the power density on the crystal will result in a larger electrical current and a larger thermal heating of the crystal. In addition, the application of a high voltage requires a uniform illumination of the crystal to ensure a uniform electrical field. On the other hand, without applied voltage, the pump beam size is fixed by the overlap of the pump and signal beams. Hence, with the same laser power available, larger power density can be obtained without any voltage applied to the crystal. Note that the requirement regarding the response time (besides being much shorter than the pulse duration of the detection laser) is not very severe when the probe beam is normal to the surface in motion (there is only a continuous change of speckle), but becomes more difficult to fulfill in the case of motion along the line of sight. This occurs in particular when in-plane ultrasonic motion is detected, which requires the probe beam to be tilted with respect to the normal to the surface. In this case in addition to the continuous change of speckle there is a Doppler shift and the requirement on the time response can be roughly evaluated as zc less than ~,/V, where V is the projected velocity along the line-of-sight. For )~=1.06 lam and V=lm/s , this means Xc < l~ts , which is more difficult to fulfill with a photorefractive demodulator than a confocal Fabry-Perot. It should be also noted that rapid response means a high value for the low frequency cut-off, so detecting with optimum sensitivity a low ultrasonic frequency signal off a part rapidly moving is not feasible and a trade-off between sensitivity and response time has to be found. It should be further noted that when the receiver tracks the Doppler shift caused by the part motion there is no such restriction. This is the case of the confocal Fabry-Perot stabilized on the scattered light.

34

J.-P. Monchalin et al.

Insensitivity to intensity modulations

Laser sources are often affected by spurious intensity modulations (e.g. relaxation oscillations), so it is useful to have a scheme that is not sensitive to such parasitic signals. This can be satisfied by a differential or balanced configuration that is easily inserted in the photorefractive interferometer [13]. Such a scheme is also possible with the Fabry-Perot but with a greater complexity [ 14].

DETECTION OF SMALL DEFECTS BY THE SYNTHETIC APERTURE FOCUSING TECHNIQUE Since there are many applications where small defects have to be detected, it is of interest to explore their detection by laser-ultrasonics, which means finding how to focus laser-generated ultrasound. Although it is possible to do that physically [15,16], i.e. to produce a region with a higher concentration of acoustic energy or alternatively to produce a region of much higher detection sensitivity, it is advantageous to perform such focusing numerically [ 17], especially for contoured parts. We have used the Synthetic Aperture Focusing Technique (SAFT), the principle of which is explained by the sketch shown in Fig. 6. We assume that the generation and detection beams are focused at the same location onto the surface. By scanning the beams or moving the part fixed on an X-Y translation table with discrete and equal steps, a 2-D mesh of generation/detection points Mi at the surface of the specimen is obtained. As shown in figure 6, if a flaw is present at point C located at a depth z within the sample, this flaw re-radiates the acoustic field originating at point Mi. The acoustic signal S(Mi,t) received at any point Mi in the measurement mesh exhibits a peak at time t = 2di /v, where v is the acoustic velocity in the material and di the distance CMi. Consequently, by summing all the signals S(Mi,t=2di/v) we separate the points C where the signals add up coherently and flaws are present, from the points C where no coherent superposition occurs and the material is sound. Furthermore, this summation increases the signal-to-noise ratio for the detection of flaws by the factor ,J-N, where N is the number of points Mi in the measurement mesh aperture (the synthetic aperture). It can also be shown that SAFT processing leads to improved lateral resolution while maintaining depth resolution.

Fig. 6 : Principle of SAFT

Recent Developments in Laser Ultrasonics

35

This data processing approach, while straightforward in its principle and implementation, is not very efficient and is very computation intensive. For this purpose, we have developed at IMI a better approach which performs data processing in the Fourier domain and which is an improvement of previously known Fourier domain SAFT processing. This improved method (F-SAFT) includes a deconvolution algorithm to improve resolution, control of the aperture and spatial interpolation [18,19]. These last two features contribute to very significant reduction of data acquisition time and processing time, making the technique more attractive for industrial use. We present in Fig. 7 a result obtained with this approach. The data were obtained with a short pulse Q-switched Nd:YAG laser operating on its fourth harmonic for generation by slight ablation, a long pulse Nd:YAG laser and a confocal Fabry-Perot interferometer for detection. The generation and detection lasers were focused onto the fiat surface of the specimen at about the same location and the tested part was move with an X-Y translation table. The test specimen was a 7 mm thick aluminum block with two flat-bottom holes of 1 and 0.5 mm in diameter and 1.5 mm deep drilled from the surface opposite to the scan surface in order to simulate flaws. The step size of the scan was 0.1 mm and for each node Mi of the measurement mesh, an ultrasonic A-scan signal was collected, digitized and stored in the computer memory. The results obtained by F-SAFT processing are shown in Figure 7b and compared with simple bandpass filtered data (Figure 7a). The C-scans were obtained by selecting the peak-to-peak value from each A-scan in a narrow gate at depths between 5.2 and 5.7 mm corresponding to the bottom of the holes. The amplitude profiles along a line crossing the holes on these C-scans are also shown. When compared to the filtered data, the F-SAFT processed data show strong improvements of the detectivity and of the lateral resolution of the flat-bottom holes. This approach has also been demonstrated to be useful for inspecting titanium parts, laser welds and diffusion bonds.

Fig. 7: C-scans and profiles of the filtered data (a) and F-SAFT processed data (b)

36

J.-P. Monchalin et al.

APPLICATION TO THE INSPECTION OF POLYMER-MATRIX COMPOSITES For this application, a TEA CO2 laser is used for generation. This laser provides substantial light penetration and causes a normally propagating ultrasonic wave as mentioned above. A high power long pulse Nd-YAG laser is used for detection coupled, for the data shown below, to a confocal Fabry-Perot interferometer. A photorefractive demodulator could have used with the advantage of greater sensitivity to low ultrasonic frequencies as mentioned above [8]. Many results regarding this application have been previously reported by IMI in various journals [2,3,20]. We show here below the result of the inspection of very complex part, which is a composite test part representing a reduced size bulkhead of a F-22 fighter. Such a complex part with fibs, sine stiffeners and many comers could not be in practice inspected with a conventional ultrasonic technique. The laser-ultrasonic system installed at McClellan Air Force Base, Sacramento, California [21] was used to perform laser-ultrasonic inspection without any preparation in 6 different settings. Since this system includes a range finder part shape could be measured at the same time. Fig. 8 shows the amplitude scan in a 3-D view obtained by combining the 6 scans [22].

Fig. 8: 3-D amplitude scan of a F-22 bulkhead

DETECTION OF CORROSION IN AGING AIRCRAFT This application to the detection of corrosion in lap joints is being developed under support from the US Air Force and uses the same system as the one developed for the inspection of composite materials. For adequate generation with a CO2 laser the aluminum alloy plates composing the joint are painted, which corresponds to the situation usually encountered on aircraft. The pulse duration of a CO2 TEA laser being typically about 100 ns and the plates being of the order of lmm thick, the ultrasonic echoes overlap, so the analysis is performed in the frequency domain. An algorithm has been developed to retrieve both the paint thickness and the residual metal thickness. Fig. 9 shows a result obtained on a particular sample where the residual thickness measured by laser-ultrasonics is compared to the one measured by X-ray transmission and with a micrometer after joint opening and removal of corrosion by-products. As seen in Fig. 9 the laser-ultrasonic and destructive data are in good agreement and demonstrate the potential of laser-ultrasonics for corrosion detection and quantification.

Recent Developments in Laser Ultrasonics

37

Fig.9: Top: Residual thickness map of a corroded lap joint measured by laser-ultrasonics compared to the X-ray thickness map measured after joint opening, Bottom: Cross section of the residual thickness taken in the middle of the largest thickness reduction patch measured with laser-ultrasonics (thick line) and compared to X-ray data (fine line) and micrometer data (circles).

CHARACTERIZATION OF STEEL AND MONITORING OF STEEL TRANSFORMATIONS Laser-ultrasonics being performed at a distance has potentially numerous applications in the steel industry. It has been demonstrated to be applicable to the wall thickness measurement of hot seamless tubes on a production line [2,24]. Laser-ultrasonics is also a unique tool for the monitoring of microstructural transformations [4]. Using a programmable resistive furnace or a thermomechanical Gleeble simulator being equipped with windows for laser coupling, we have shown that the technique can be used to monitor phase changes, grain growth and other microstructural changes for several types of carbon steels [25,26]. Fig. 10 shows the monitoring of austenitic grain growth in a A36 steel sample obtained by measuring ultrasonic attenuation. A model was developed to relate the measured attenuation to the grain size. The

38

J.-P. Monchalin et al.

grain size measured by laser-ultrasonics is compared to the one measured by rapid quenching and metallographic examination. In other studies attenuation and velocity were monitored as steel samples were heated through the phase transformation and cooled to room temperature. These measurements allowed observing grain refinement both during heating and cooling through the phase transformation of carbon steels. We have also demonstrated the application to the measurement of anisotropy or texture in steel sheets [27]. The approach is based on an analysis in the frequency domain and the identification of longitudinal and shear resonance frequencies (see Fig. 11). Anisotropy makes the shear resonance peaks to appear as doublets. From these measurements, the orientation coefficients (precisely W400 and W420) and the sheet thickness are deduced. The average plastic strain ratio ?- is known to be proportional in good approximation to W400. This is verified by the data shown in Fig. 12 obtained with various grades of steel. Therefore laser-ultrasonics allows determining F, which is an important parameter characterizing texture and sheet formability. This approach has not only been tested in the laboratory but also on-line and has revealed texture variability along the length of the various tested production coils [27].

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The measuring method is shown in Fig. 3. The Lamb wave generated by 0.5 mm pencil lead break arrives two sensors of which distance is 10 cm in succession. Of course its front part is extensional mode. The velocity is obtained by its distance and its arrival time difference. The velocities obtained in each 5 degree interval from 0 to 90 are shown in Fig. 4. From curve fitting the following empirical equation was obtained. This equation is used for source location.

I1o=(l+c~ 2

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0< 0 1 mm interval in fact. There is no concentration in type 3 and type 4. From these results, we found that SMA wire and optical fiber had no bad effect on fracture of host material regardless they had 30-40 times diameter of carbon fiber ( 6/~m ). In addition, there is no degradation of maximum load. CONCLUSIONS 1. The measured velocity of extensional mode Lamb wave is available to the high reliable AE source location on cross ply CFRP. 2. We found that the embedded SMA wires and optical fibers did not give bad effects on the mechanical degradation of host material ( cross ply CFRP ), from the fact that no AE event was found in the vicinity of them.

60

J.H. Koo et al.

REFERENCES 1. Jang, T. S., Lee, J. J., Lee, D. C. and Huh, J. S. (1999)J. Mater. Sci. 34 5853-5860. 2. Brown, T., Wood, K., Childers, B., Cano, R., Jensen, B. and Rogowski, R. (1999) SPIE 3674,6071. 3. Chang, F. (2000) SPIE 3990-17. 4. Prosser, W. H., Gorman, M. R. and Dorighi, J. (1992)J. Compo. Mater. 26, 418-427. 5. Prosser, W. H. and Gorman, M.R. (1994) Proceedings ofthe 1994 ASNT Spring Conference, 152154. 6. Kwon, O. Y. and Joo, Y. C. (1997) Fourth Far East Conference on NDT (FENDT' 97) 219-228. 7. Ziola, S. M. and Gorman, M. R. (1991)J. Acoust. So. Am. 90, 245-251. 8. Koo, J. H., Kim, B. N., Enoki, M. and Kishi, T. (1999)J. JSND148, 283-288.

Nondestructive Characterizationof Materials X Green et al. (Eds) 9 2001 Elsevier Science Ltd. All rights reserved.

61

EVALUATION OF FRACTURE BEHAVIOR IN SIC/TI-6AL-4V C O M P O S I T E BY ACOUSTIC EMISSION TECHNIQUE

AKIO HIROSE and KOJIRO E KOBAYASHI

Department of Manufacturing Science, Graduate School of Engineering, Osaka University, 2-1, Yamadaoka, Suita, Osaka 565-0871, Japan

ABSTRACT A continuous SiC fiber reinforced Ti-6A1-4V alloy composite having excellent specific strength and moderate high temperature properties is attractive as a future light-weight structural material. However, a degradation in the tensile strength of the composite occurs after thermal exposure that results in growth of the fiber/matrix interfacial reaction zone. Especially, the strength significantly decreases when the reaction zone thickness exceeds 1~tm. An acoustic emission (AE) technique has been applied to estimate the mechanism of the degradation in the strength of the composite. It is assumed from the AE analysis during the tensile test and fractography analysis that a fiber-matrix reaction zone debonding occurs before failure of the fibers. Therefore, the degradation in the strength due to the reaction zone growth is considered to be attributed not to a notch effect of cracks in the reaction zone, but to the decrease in the strength of the fibers. KEYWORDS Metal matrix composites, SiC/Ti-6A1-4V composite, tensile test, acoustic emission, fracture behavior, fractography, fiber/matrix interfacial reaction zone, debonding

INTRODUCTION Metal matrix composites (MMCs) can achieve higher specific strength and modulus, and better creep resistance at elevated temperatures than those of conventional metallic materials, such as Ni base super alloys. Among the MMCs, a continuous SiC fiber reinforced Ti-6A1-4V alloy (hereafter described as SiC/Ti-6A1-4V) composite having excellent specific strength and moderate high temperature properties is a candidate for future light-weight structural materials. However, since the interface between the fiber and the Ti alloy matrix is thermodynamically unstable, chemical reactions occur at the interface during a thermal exposure [1 ]. The growth of the fiber/matrix interracial reaction zone (hereafter described as F/M reaction zone) results in a degradation in the tensile strength of the composite after the thermal exposure [2, 3].

A. Hirose and K.F. Kobayashi

62

In the present work, the effect of the F/M reaction zone thickness on the tensile strength of the SiC/Ti-6A1-4V composite was quantitatively evaluated. Moreover, the mechanism of the degradation in the strength of the composite was discussed on the basis of examining the fracture surfaces of the composite specimens and measuring acoustic emission (AE) during the tensile test.

EXPERIMENTAL PROCEDURE Material The SiC fiber used in this study was an SCS6 continuous fiber supplied by Textron Inc. The SCS6 fiber is produced by means of chemical vapor deposition. The fibers were approximately 1401xm in diameter. The composites were synthesized by vacuum hot pressing. Firstly, layers of the Ti6A1-4V thin foils (80lxm thickness) and SiC fibers were stacked together. The lay-ups were placed on a carbon-lot and binder was driven offby heating of 773K. The fabrication was performed at 1173K and 73.5MPa for 3.6ks, to produce a ten-ply composite with Vf=-45% and a one-ply composite with Vf=-22% in form of 20mmx80mm panel. The composite specimens were isothermally annealed in vacuum for 14.4, 32.4, 57.6, 90.0, 129.6 and 230.4ks at 1123K, and for 129.6 and 230.4ks at 1223K.

Metallographic observation

The specimens for metallographic observations were cut perpendicular to the fiber direction and were mechanically polished followed by etching with an aqueous solution of 4.4vol.%HF and 26vol.%HNO3. The samples were observed using a scanning electron microscope (SEM). Measurements of the reaction zone thickness were made on micrographs of the fibers at high enough magnification (x6000).

Tensile test

The mechanical properties of the composites and the extracted fibers from the composite were evaluated by a tensile test. The tensile test was performed on an Instron type testing machine at

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Evaluation of Fracture in SiC/Ti-6AI-4V

63

room temperature, using a cross head speed of 8.33xl 03mm/s. After the annealing, fibers were taken from the composite by solving matrix in a solution of HF+HNO3. Figure 1 shows the shapes of the tensile test specimens for the composites and the extracted fibers. AE signals were measured during the tensile test of the composite specimens. AE signals were detected by a transducer and the event counts and event energy were measured after AE signals were amplified 60dB.

RESULTS AND DISCUSSION Effect of F/M reaction zone growth on tensile strength of composite Ultimate tensile strength and Young's modulus of as-fabricated and heat-treated composites are shown in Fig. 2. Young's modulus of the composites were almost constantly 210-220GPa regard-

Fig. 2. Tensile strength and Young's modulus of as-fabricated and annealed ten-ply composites.

Fig. 3. Tensile strength of composite vs. F/M reaction zone thickness.

64

A. Hirose and K.F. Kobayashi

less of the annealing temperature and time. On the other hand, the tensile strength reduced as annealing time increased. It is considered that the degradation of tensile strength depends on the growth of interfacial reaction zone. Figure 3 shows the relation between the tensile strength and the average thickness of the reaction zone. The tensile strength reduced rapidly when the reaction zone thickness exceeded 1.0~tm. Possible reasons that can cause the degradation are as follows. (1) Effective cross-sectional area of the fiber decreases as reaction zone thickness increases. (2) Micro cracks in brittle reaction zone work as notches, and so fibers are fractured at low stress level. (3) Tensile strength of SiC fiber itself is degraded with increasing reaction zone growth. In this study, when the reaction zone thick reached 3ptm, the diameter of fiber was 21am smaller than that of the original fiber. The tensile strength degradation due to decreasing diameter of fibers is calculated to be 64MPa. The actual strength degradation is much larger than the calculated value. Thus the factor (1) is not a main reason of the strength degradation. It has been reported that tensile strength of continuous fiber reinforced MMCs is rapidly de-

Fig. 4. SEM images of fracture surfaces of as-fabricated and short-time annealed composites: (a) surface of fiber side; (b) surface of matrix side.

Fig. 5. SEM images of fracture surfaces of long-time annealed composites:(a) surface of fiber side; (b) surface of matrix side.

Evaluation of Fracture in SiC/Ti-6AI-4V

65

graded by the factor (2) [4, 5]. It is necessary for the factor (2) causing the degradation that fibers are not debonded from the reaction zones before cracks in the reaction zone reach the fibers. To consider this problem, the fracture surfaces of the tensile test specimens were observed with SEM. Figure 4 shows the fractography of the composite whose reaction zone was approximately 0.4txm. The coating layers of the fibers were debonded from the fiber, and adhered to the matrix alloy. Figure 5 shows the fractography of a composite subjected to a long time annealing. The reaction zone adhered to the matrix. That is to say, debonding occurred at the SiC fiber/reaction zone interface. It has been reported that, in SiC fiber-Ti alloy system, bonding strength of reaction zone/fiber interface was lower than that of reaction zone/matrix interface [6, 7]. Thus, It is assumed that the reaction zone was debonded from fibers before cracks in the reaction zone reached the fibers because no reaction zone adhered to the fractured fibers was observed on the fracture surface as shown in Fig. 5. To estimate when the debonding occurred, AE analysis during the tensile tests was carded out.

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0

0

1O0

200

Time (a)

300

O.

400

(sec)

AE measurements of ALFRP 5000

1000 800 b r~

600

stress

j

AE energy

4000

3000

400

2000

200

1000

0

0

1O0

200

300

400

Time

(sec)

500

>

=

o

i 0 600

(b) AE measurements of CFRP (The specimen size is the same with that of ALFRP) Fig. 6

AE measurements during the loading

Fig.6 (b) shows AE measurements of CFRP and Fig. 7 shows the distribution of the crack position. The cracks in ALFRP are observed by the transmitted-light technique and the cracks in CFRP are observed by edge replica. In ALFRP specimen, the fracture is locally and abruptly. In CFRP specimen, two or three main cracks grew and then one main crack broke the specimen. So the crack distribution is not uniform along the longitudinal direction.

NDE Techniquefor ALFRP

Fig. 7

81

Relation b e t w e e n the n u m b e r of cracks and A E event counts (Crack counts is the number o f cracks in the observation area)

Table 2

Comparison of three damage investigation methods

Detectable damage type

Method

Fiber-bragg-grating sensor

Limitation

Matrix crack

Delamination

Fiber breakage

Excellent

Medium

Bad

Difficult to detect growing damage Acoustic emission

Transmitted-light technique

of target

FBG can be mounted in the target easily.

Medium

Medium

Excellent

No limitation

Excellent

Excellent

Medium

Target has to be semi-transparent

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K. Tanaka, H. A oyama and N. Takeda

COMPARISON OF THE THREE DAMAGE DETECTION METHODS Table 2 compares the abilities of the three methods for detecting three kinds of damage. The transmitted-light technique is clearly the most suitable for detecting the damage in the load support system of the superconducting magnet even though it is limited to certain target materials.

CONCLUSION The internal damage of ALFRP was detected by three nondestructive methods: transmitted-light, FBG sensor, and acoustic emission. The transmitted-light technique has several advantages over other two: it could measure the exact size of the internal damage directly. Furthermore, it is not influenced by a high magnetic field. It is thus concluded that the transmitted-light technique is the best for estimating the degree of damage in ALFRP structure.

ACKNOWLEDGEMENTS This research was conducted as a part of the "R&D for Smart Materials Structure System" project within the Academic Institutions Centered Program supported by NEDO (New Energy and Industrial Technology Development Organization), Japan.

REFERRENCE 1. H. Aoyama, H. Watanabe and M. Terai, (1998) Proc. of 1998 ASME IMECE, Rail Transportation, pp. 77-82

ELECTROMAGNETICS AND RADIOGRAPHY

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An Alternative to Film-Based Flash Radiography Using a High Speed Camera and Intensifying Screen Dominick Salafia, Norberto L. DeLeon and James F. Outlaw Metrology and Simulation Division Material Test Center U.S. Army Yuma Proving Ground Yuma, Arizona 85365 Abstract: The field of Radiography has benefited from the digital imaging revolution that is starting to dominate many of the fields that have relied on film-based media in the past. The specialized field of Flash Radiographyhas also benefited from this same technology. Still, Flash Radiography has unique requirements and limitations that differ from the typical time/power exposures constraints. In addition, flash radiography in some cases may require a large size film image format of 14" X 51". These limitations in comparison with available and advancing digital image technology has led an effort at the U.S. Army Yuma Proving Ground installation to seek an effective method to replace film based radiographs. This has led to the concept of using phosphor intensifying screens with a sub-microsecond persistence and a high speed camera to capture an x-ray image and has the potential to result in significant long term savings over the current methods being used. This paper describes the proof of concept and the results of this preliminary test show this to be a potential feasible solution. INTRODUCTION The U.S. Army Yuma Proving Ground (YPG) is a Test and Evaluation facility of longrange munitions and weapons. One of the test functions performed at YPG is flash radiography. Flash Radiography is the practice of taking an x-ray image of an object while traveling at high speed. In our case, it is the practice of taking an x-ray snap shot of a projectile as it exits the muzzle of a large caliber cannon tube when it is fired. An x-ray image is needed this close to the muzzle since muzzle gases along with the fireball obscure the image and prevent a regular high speed video camera from being used. This paper presents the results of preliminary tests using phosphor screens to capture and digitize the image as opposed to the more traditional way of using chemical based x-ray film.

High Speed Camera/Phosphor Screen Concept The principle of operation is straightforward in concept. The idea is to expose a fast persistence phosphor-intensifying screen, which is irradiated by a triggered x-ray pulse. The x-rays impacting the phosphor screen causes it to illuminate at levels proportional to the level of energy it receives. This difference in x-ray energy levels as it passes through the projectile, is what forms an image on the phosphor screen. The high-speed camera is set to trigger at the right moment to capture the image projected on the phosphor screen that has sub-microsecond persistence. The timing of the camera is critical when considering the x-ray pulse is approximately 100 nanoseconds in duration and the highspeed camera's maximum window is one millisecond long.

D. Salafia, N.L. DeLeon and J.F. Outlaw

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Live Fire vs. Static The test was conducted using a static setup in place of a live fire test. An analysis determined that the important variables between a static test and a live fire would not affect the test results. The 100-nanosecond pulse duration is the main determining factor and it is non adjustable. The other variable is the adjustable KV power that determines the amount of x-ray penetration. It was determined that the static test could be conducted to prove the concept and achieve substantial cost savings in terms of time, materials and personnel. The test was conducted to simulate the same variables that would be present in a live fire test.

Equipment and Setup The test equipment consisted of a 5 foot by 5 foot by 18 inches high box as shown in Fig. 1. The phosphor-intensifying screen was placed on the inside of the box and the box was sealed to achieve a light tight enclosure. A front surface mirror was placed at a 45 degree angle to reflect the image onto a high speed camera located at a right angle on the outside of the box with the lens extending inside the box. The camera was initially setup with a 25mm lens then replaced with a 50mm lens to give better results. The interior of the box was painted flat to minimize glare and reflections. The camera was set up on the side of the box and protected with lead shielding. It is not possible to setup the camera inline with the source and be able to protect it from the x-rays at the same time. The x-rays will not damage the camera but would impact the CCD plates directly and affect the quality of the image. In addition, two 1-inch thick aluminum plates were placed in front of the phosphor screen, as they would be in a live fire test. In a live fire test, these aluminum plates serve to protect the film cassettes from the close proximity of the muzzle blast. The phosphor intensifying screen is a large format 14" x 51" and is placed inside the light tight enclosure box.

Figure 1. Equipmentsetup for phosphor screenand high-speedcamera.

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87

Discussion The high-speed camera used for this test has a pixel resolution of 1280 x 1024 dots per inch (DPI). It uses proprietary software to capture the image in a 12-bit grayscale resolution. In order to see the image in its full dynamic range, it is necessary to have the proper software and expensive high-resolution monitors to take full advantage of the 12bit resolution. The image in Fig. 1 has been converted to a common industry "tiff" format. The tiff format is an 8-bit image resolution. This 8-bit image translates to a resolution of 256 shades of gray instead of a 12-bit resolution with a full dynamic range of 4096 shades of gray. This difference in resolution produces images that appear lower in quality than they actually are. However, for the purpose of this proof of principal, the current equipment and the eight bits resolution is sufficient to prove the concept.

Figure 2.

8 bit x-ray image of a Projectile.

The experiment consisted of exposing the phosphor intensifying screens at two different kilovolt levels. The following images describe the voltage power levels used in this experiment. In addition, two different intensifying screens were used. The original phosphor screen is an FSL-1. The other intensifying screen is the one normally used with the film radiographs. The images illustrated are the best samples from the FSL-1 phosphor screen. The second phosphor-intensifying screen produced images of lesser quality and is not included in this discussion. Although there are numerous types and manufacturers of phosphor screens the objective of this study was to prove a concept. The authors suggest that the selection of an optimal screen be the subject of another experiment and welcome the consultation of researchers that have perused this experimentation. Aluminum plates were used to simulate the total amount of protective material that the x-rays would be required to penetrate under actual test firing conditions prior to contacting the phosphor screen. The aluminum plates are necessary to protect the x-ray source and film cassettes from the blast overpressure incurred a during live fire test. The camera was shielded with a 88 thick lead plate in front and thin flexible lead sheets around the camera. The white speckled spots on the images are due to insufficient lead shielding on the camera and can easily be eliminated with additional lead shielding.

D. Salafia, N.L. DeLeon and J.F. Outlaw

88

Figure 3

x-ray of a projectile from taken at 22 Kilovolts through 1 inch of aluminum.

Figure 4

x-ray of a projectile from taken at 27 kilovolts through 1 inch of aluminum.

Figure 5

x-ray of a projectile front taken at 27 Kilovolts through 2 inches of aluminum.

An Alternative to Film Based Flash X-Ray

Figure 6

89

x-ray of a projectile front taken at 27 kilovolts and no aluminum plates.

Figure 7 x-ray of projectile front using a larger mirror, taken at 27 kilovolts, no aluminum plates. Figures 7,8,9 and 10 were obtained after substituting the 1" aluminum plates with 1" Lexan plates. The plates were 24" x 24" thus resulting in a smaller cropped image of the same projectile repositioned.

Figure 8, 27 kilovolts no Lexan plates

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Figure 9, 27 kilovolts with l" of Lexan plates.

Figure 10,27 kilovolts ~th 2" of Lex~ plates.

Conclusion This test shows a promising potential for using the phosphor intensifying screens in flash radiography to capture and digitize the image in an effort to eliminate the need for chemical based film processing. The random white specs on the images show the need for improved shielding of the camera. The x-ray attenuation / filtering caused by the aluminum and Lexan plates was greater than expected. This needs to be addressed and investigated further and a cost/benefit study needs to be conducted on how best to approach a possible solution to this problem. This may include more powerful x-ray pulsars, a different high speed camera, different type/manufacture phosphor screens with higher sensitivity, different type material to replace the aluminum plates currently used or a combination of any of the above. The successful implementation of this concept using phosphor intensifying screens to capture and digitize large format images in flash radiography has the potential for significant long term savings. The savings will result from fewer number of personnel used to maintain, setup, operate, and post test cleanup of the equipment. In addition, environmental concerns with disposal of film developing chemicals will be eliminated.

REFERENCES 1. Cossell, Frank HandlandPhotonics, Cupertino Califonia, technicalconsultation. 2. Photo Optics Division, U.S Army Yuma Proving Ground. Technical support and consultation.

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INVESTIGATION OF L O W DENSITY CORE PROCESSES FOR P R O D U C I N G ULTRA L I G H T W E I G H T METALS USING X-RAY COMPUTED TOMOGRAPHY

W.H. GREEN US. Army Research Laboratory Weapons and Materials Research Directorate ATTN." AMSRL- WM-MD Building 4600 Aberdeen Proving Ground, Maryland, USA 21005-5069 J.M. WINTER, JR. The Johns Hopkins University Center for Nondestructive Evaluation and Department of Materials Science and Engineering Maryland Hall~3400 N. Charles Street Baltimore, Maryland, USA 21218-2681 R.E. GREEN, JR. The Johns Hopkins University Center for Nondestructive Evaluation and Department of Materials Science and Engineering Maryland Hall/3400 N. Charles Street Baltimore, Maryland, USA 21218-2681

ABSTRACT In recent years, several universities, government laboratories, and private industries have been developing specific process technologies, analytical modeling tools, and characterization methods for highly porous metals and alloys frequently collectively termed "ultra lightweight metals." The goal has been to achieve a family of metallic structures that are analogs to the organic cellular materials that exhibit high stiffness and a low specific weight. A number of quite different and distinct processes have evolved and are still largely not yet mature, including a variety of methods that start with metal powders to produce Low Density Core (LDC) metallic materials. To date, the imaging capabilities of x-ray computed tomography (CT) have not been generally employed to nondestructively examine the internal structure of the products formed by LDC methods. X-ray CT imaging reveals the sizes, shapes, and distribution of internal "cells" in a scan, or "slice." This paper will discuss CT images of LDC specimens, including some "in-situ" production images if possible. KEYWORDS Low density core, ultra lightweight metal foams, pore size, pore distribution, pore/metal composition, aluminum, steel, x-ray computed tomography.

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W.H. Green, J.M. Winter, Jr., and R.E. Green, Jr.

INTRODUCTION A number of quite different and distinct processes for producing highly porous metallic materials (i.e., metallic foams) have evolved and are still largely not yet mature, including a variety of methods that start with metal powders to produce LDC metallic materials. Applications for metallic foams include impact/blast energy absorption in vehicles, ships, and lightweight aircraft structures, thermal insulation barriers, vibration damping and sound absorption, and fluid/gas storage media. Potential Army applications include impact/blast protection in vehicle armor systems, fire retardant doors and walls in storage areas, and vibration damping in bases of gun tubes and tailbooms of helicopters. The deformation behavior and shock wave absorption capability of LDC foams at high strain rates up to ballistic ranges (105/sec) have been investigated [1]. However, nondestructive evaluation (NDE) has not been generally employed to examine the internal structure of these types of foams. Only recently have the imaging capabilities of x-ray CT been used to nondestructively examine the internal structure of LDC foams [2]. X-ray CT imaging reveals the sizes, shapes, and distribution of internal pores in a scan, or "slice." The excellent dimensional accuracy, high spatial resolution, and digital nature of CT images make it possible to obtain detailed and accurate two-dimensional (2-D) and three-dimensional (3-D) geometric information. The advantages of CT become readily apparent when geometric data unobtainable by other NDE techniques is input to engineering calculations for material characterization and production processes. This data includes pore size, pore distribution, wall thickness, and void vs. metal composition, among other structural parameters. X-ray CT is a powerful method for imaging, mapping, and measuring parameters of the complex internal structures of LDC foams. The examination of both aluminum and steel foams will be presented.

X-RAY COMPUTED TOMOGRAPHY X-ray CT is broadly applicable to any material or test object through which a beam of penetrating radiation may be passed and detected, including metals, plastics, ceramics, metallic/nonmetallic composite material, and assemblies. The principal advantage of CT is that it provides densitometric (that is, radiological density and geometry) images of thin cross sections through an object. Because of the absence of structural superimposition, images are much easier to interpret than conventional radiological images. The user can quickly learn to read CT data because images correspond more closely to the way the human mind visualizes 3D structures than 2-D projection radiology (i.e., film radiography, real-time radiography, and digital radiography). Further, because CT images are digital, the images may be enhanced, analyzed, compressed, archived, input as data to performance calculations, compared with digital data from NDE modalities, or transmitted to other locations for remote viewing, or a combination thereof.

THREE-DIMENSIONAL VISUALIZATION OF MULTIPLE TOMOGRAPHIC SCANS The excellent dimensional accuracy and the digital nature of CT images allow the accurate volume reconstruction of multiple adjacent slices. The slices are "stacked" to provide 3-D information through out the entire object or a section of the object. The two ways of visualizing volumetric data are multiplanar reconstruction (MPR) and 3-D reconstruction.

Investigation of Metal Foams UsingX-Ray CT

93

Multiplanar reconstruction (visualization) displays top, from, side, and oblique slices through the object. The orientation of the top slice is parallel to the cross-sectional image plane. The front slice is orthogonal to the top slice. The side slice is orthogonal to both the top and front slices. The oblique slice can be placed on any one of the other three slices. The MPR display is similar to an engineering drawing. However, each view (i.e., top, from, side, and oblique) is a slice with finite thickness through the object, not a 2-D projection. The top, front, and side slices can be moved anywhere in the reconstructed volume. The oblique slice can be rotated through 360 degrees. Dimensional analysis, image processing, and automated flaw detection and measurement can be performed with MPR images. The volumetric data is displayed as a 3-D solid object in 3-D reconstruction (visualization), and the orientation of the solid in space can be changed to facilitate different views. Secondly, illumination from a computer-generated light source can shadow surface irregularities from any arbitrary direction. The solid can also be "virtually" sectioned by only displaying part of the reconstructed volume, which creates a "virtual" cutting plane on the solid showing the xray CT density values on that plane. This plane may be orthogonal to the cross-sectional image plane. In effect, virtual sectioning shows the surface on the cutting plane, as it would look if the object were actually destructively sectioned along that plane.

LOW DENSITY CORE FOAM MANUFACTURING One process, invented and patented by the Fraunhofer Institute for Applied Materials Research (IFAM) in Bremen, Germany, is based on a mixture of metal powder and a suitable foaming agent. The mixture is consolidated into a high-density compact (via uniaxial or isostatic pressing, extrusion, etc.), worked into a foamable semi-finished product, and then heated to a temperature near the melting point of the metal. At the process temperature, the foaming agent decomposes, forming a gas that is trapped inside the compacted powder body. Gas bubbles create voids within the expanding body of semi-solid metal and are retained during solidification. The process results in a lightweight structure with a high degree of closed-cell porosity.

RESULTS AND DISCUSSION

Aluminum Foam Specimen The specimen was provided for the authors by the Fraunhofer Resource Center in Newark, Delaware, USA. It was made of 6061 aluminum alloy, with a measured final porosity of 62%. In this case, the mixed powder with its foaming agent was compacted and heated in a closed copper die. The specimen was a 6.4-mm-thick plate, 100-mm-by-100-mm square [3].

Tomographic technique. The specimen was inspected using a customized ACTIS 600/420 CT system designed and constructed by Bio-Imaging Research, Inc., and installed at the U.S. Army Research Laboratory at Aberdeen Proving Ground (APG), Maryland, USA. Fifty contiguous slices (images) were performed over a section of the specimen. The slice thickness and slice increment were 0.100 mm. A total of 4.900 mm was scanned in the through thickness direction. Each slice was scanned in offset-rotate only (RO)/130% mode using the 160 keV

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W.H. Green, J.M. Winter, Jr., and R.E. Green, Jr.

microfocus x-ray tube and the image intensifier. The wedge calibration was done through air. The cross-sectional image plane is parallel to the 100-mm-by-100-mm plane of the specimen. The source-to-object-distance (SOD) and source-to-image-distance (SID) were 397.00 mm and 648.00 mm, respectively. The tube energy and current used were 160 keV and 0.015 mA, respectively. The focal spot size was 20 microns.

Tomographic scans. Fig. 1 is a CT slice in the middle of the scanned section. It shows that the "cell" size is approximately uniform, with some larger and non-spherical pores present. The diameters of pores A and B are about 1.62 mm and 2.49 mm, respectively. The long dimension of pores C, D, and E are about 8.48 mm, 10.33 mm, and 15.76 mm, respectively. Higher density (i.e., whiter) bands of more compacted aluminum (powder) can also be seen in Fig. 1. The width of bands in the "densified" aluminum network ranges from about 1.4 mm (band F) to about 3.6 mm (band G). Fig. 2 is a CT slice 1.200 mm above the middle of the scanned section. It also shows that the cell size is approximately uniform, with the same pores as in Fig. 1, and small changes in the geometry of the bands. The slices are quite similar, indicating little change in the geometry and distribution of pores and bands over this scale. Fig. 3 is a "re-reconstructed" image of part of the top area of the slice shown in Fig. 1. Only the CT data in this particular area has been reconstructed; the CT data outside of this area has not been reconstructed. This is not the same as zooming in on the image (Fig. 1), in which the magnification eventually surpasses the spatial resolution of the image and individual pixels become apparent. Fig. 3 shows the pore geometry and distribution in significantly greater detail.

Fig. 1. Slice in middle of scanned section

Fig. 2. Slice 1.200 mm above middle

Fig 3. Re-reconstructed (image) close-up of part of top area of slice in Fig. 1

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95

Multiplanar and three-dimensional visualization. Fig. 4 is a multiplanar visualization of the scanned section (50 contiguous slices), with the top slice view parallel to the CT image plane. Dashed lines show the locations of the front and side slices relative to the top slice; the oblique slice corresponds to the diagonal dashed line on the top slice view. Fig. 4 shows that the bands of densified aluminum visible in the top slice extend through most or all of the scanned section. Fig. 5 is a 3-D visualization of the entire scanned section, in which the surface features are real and accurate. Fig. 6 is a 3-D visualization with about 1/4 of the scanned section virtually cut away perpendicular to the CT image plane. Two bands of densified aluminum are visible on the virtual surface.

Fig. 4. Multiplanar visualization with top slice view in middle of scanned section

Fig. 5. 3-D solid visualization of entire scanned section

Fig. 6. 3-D solid visualization with about 1/4 of scanned section virtually cut away

Steel Foam Specimen The specimen was provided for the authors by Ultraclad Corporation in Andover, Massachusetts, USA. It was made of carbon steel with 2.5% carbon, with a final fractional density between 58% and 64%. The porosity was between 36% and 42%. The foaming agent was chromium nitride. The specimen was an 18-mm-diameter cylinder [4].

Tomographic technique. The specimen was inspected using the same system as was used for the aluminum specimen. Fifty-one contiguous slices were performed over a section of the specimen. The slice thickness and slice increment were 0.250 mm. A total of 12.500 mm in

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the axial direction was scanned. Each slice was scanned in RO mode using the 160 keV microfocus x-ray tube and the image intensifier. The wedge calibration was done through the specimen itself, since it was cylindrical. The cross-sectional image plane is perpendicular to the axis of the specimen. The SOD and SID were 78.00 mm and 648.00 mm, respectively. The tube energy and current used were 160 keV and 0.055 mA, respectively. The focal spot size was 20 microns.

Tomographic scans. Figs. 7, 8, and 9 are CT slices at different locations in the scanned section perpendicular to the axial direction; they are approximately 1/4, 1/2, and 3/4 of the length scanned (12.500 mm) from the bottom of the scanned section, respectively. Each slice shows a high degree of non-uniformity of both pore size and "wall thickness". In fact, wall thickness is difficult to define in this type of geometric structure. Secondly, these slices show that the geometry and distribution of pores vary significantly over an mm-order-of-magnitude scale. Fig. 10 is another CT slice with pores varying in size from about 450 microns in diameter (Pore A) to about 4.59 mm by 2.48 mm (Pore B); those pores larger than about 2.0 mm in diameter are non-spherical (Pores B and D). Pore C is about 2.0 mm in diameter. Fig. 11 is the binarized image of the slice in Fig. 10, in which an appropriate CT density (i.e., gray level) was determined for the image threshold followed by setting the image width equal to two (i.e., only two levels). This produces an image in which material is white and pores are black, which allows porosity to be calculated in any region-of-interest (ROI) by taking the ratio of the number of pixels that are black to the total number of pixels in the ROI. In this particular slice (Fig. 10) the local porosity is about 26%.

Fig. 7. Slice 1/4 of scanned height from bottom

Fig. 8. Slice 1/2 of scanned height from bottom

Fig. 9. Slice 3/4 of scanned height from bottom

Multiplanar and three-dimensional visualization. Fig. 12 is a multiplanar visualization of the scanned section (50 contiguous slices), with the top slice view parallel to the CT image plane. Again, dashed lines show the locations of the front and side slices relative to the top slice; the oblique slice corresponds to the diagonal dashed line on the side slice. The oblique slice is an ellipsoidal cross-section through the center of the scanned volume from one "side" to the other, indicating (along with the other slice views) the non-uniformity of pore geometry and distribution in every direction. Fig. 13 is another multiplanar visualization, indicating how the entire scanned volume can be systematically examined by changing the locations of the views. For example, Pore A is a~proximately spherical with a diameter of about 1.37 mm and a volume of about 2.57 m m . Fig. 14 is a 3-D visualization of the entire scanned section, in which the surface features are real and accurate. Fig. 15 is a 3-D visualization with half of the

Investigation of Metal Foams UsingX-Ray CT scanned section virtually cut away parallel to the axial direction. distribution of pores on the virtual surface is readily apparent.

Fig. 10. Slice showing various pore sizes

97 The geometry and

Fig. 11. Binarized slice showing pore areas

Fig. 12. Multiplanar visualization of scanned section with ellipsoidal cross-section

Fig. 13. Multiplanar visualization of scanned section with all views through the same pore

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Fig. 14. 3-D solid visualization of entire scanned section

Fig. 15. 3-D solid visualization with 1/2 of scanned section virtually cut away

SUMMARY Aluminum and steel LDC foam specimens were evaluated using x-ray CT. Individual CT slices were analyzed to determine representative pore and, in the case of the aluminum specimen, "densified" band sizes. Individual CT slices were also analyzed to determine representative porosity levels in the steel specimen. Multiplanar and 3-D solid visualization were used extensively to determine characteristics of the internal structure of both specimens. Low Density Core production processes have not yet reached maturity. Nondestructive evaluation techniques such as x-ray CT provide an excellent way to obtain important, possibly critical, engineering data about LDC foams. X-ray CT can provide a promising tool for the process engineer to provide detailed feedback on the effect of process variables. Additionally, CT provides specific structural information to support the effort to develop computer models for predicting mechanical behavior of such products in various applications.

REFERENCES Yu, C.-J., Eifert, H., Banhart, J., and Baumeister, J. (1998) Journal of Materials Research Innovations 2, 3. Winter, J., Green, R., Waters, A., and Green, W. (1999) Research in Nondestructive Evaluation 11, 199. Yu, C.-J. (1998), private communication, Fraunhofer USA Resource Center, USA. Runkle, J. (1999), private communication, Ultraclad Corporation, USA.

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X-RAY CHARACTERIZATION OF THE STRAIN STATE IN A TENSILE DEFORMED Ti-Ni-Cu SHAPE MEMORY ALLOY Y. KISHI1, Z. YAJIMA1,K. SHIMIZU1 and M. ASAI2 1 :AMS R&D Center, Kanazawa Institute of Technology 3-1 Yatsukaho, Matto, Ishikawa 924-0838, Japan 2 :Materials Research Center, The Furukawa Electric Co., LTD. 2-4-30kano, Nishi-ku, Yokohama 220-0073, Japan ABSTRACT X-ray diffraction study was carried out to clarify the strain state in a tensile-deformed Ti-41at.%Ni-8.5at.%Cu alloy, which was solution-treated previously after some thermo-mechanical treatments. Before the X-ray study, B2--->B19 martensitic transformation start temperature, Ms', of the alloy was determined to be 338 K by a differential scanning calorimetry. The stress - strain curve obtained for the alloy at 295 + 1 K was divided into four stages, an elastic deformation of existing martensites, an elongation due to the reorientation of those martensite variants and two stages of plastic deformation. X-ray diffraction patterns and scanning electron micrographs were taken from three specimens tensile-deformed under 100, 300 and 700 MPa. Diffraction peaks between 40 and 50 degrees in 20became broader with increasing the applied stress, and these broad peaks were overlapped to each other. Slight relieves, which were due to the reorientation of martensite variants, were observed on the specimen surface deformed under 100 and 300 MPa, and slip lines and micro-cracks were observed on that deformed under 700 MPa.

KEYWORDS Martensitic transformation, Tensile deformation, Shape memory alloy, X-ray diffraction profile

INTRODUCTION Ti-Ni alloys are now very famous because of their shape memory effect and superelasticity, and they are the only alloys which are widely used for practical applications, such as thermostatic mixing valve, eyeglass frame, orthodontic wire, and others [1 ]. Recently, ternary Ti-Ni-Cu alloys are also developed to improve the shape memory and superelasticity characteristics in binary Ti-Ni alloys. In order to further develop those binary and ternary alloys for more wide range of applications, it is necessary to clarify more exactly the mechanical properties associated with the martensitic transformations in those alloys. In the present paper, stress - strain behavior of a solution-treated Ti-Ni-Cu alloy was examined at 295 + 1 K, and X-ray diffraction analysis was carried out to clarify the strain state in the tensile-deformed alloy.

1 O0

Y. Kishi et al.

EXPERIMENTAL PROCEDURE The Ti-Ni-Cu alloy used in the present work was made by high-frequency vacuum induction melting, its chemical composition being Ti-41at.%Ni-8.5at.%Cu. The ingot was subjected to cold-rolling and annealing repeatedly, and plates with 100 mm length, 20 mm width and 5 mm thickness were made and supplied for the present study. Specimens for tensile tests and X-ray diffraction analyses were prepared from the plates, whose shape is ribbon with L ( active gauge length ) = 18 mm, w ( width ) - 3.5 mm and t ( thickness ) - 1.5 mm, as shown in Fig. 1. The ribbon-shape specimens were then solution-treated at 1123 K for 1.8 ks in argon atmosphere, and were lightly polished with alumina powders in order to remove a thin oxide surface layer. The tensile tests were then carried out at 295 + 1 K by using a computer controlled servo-hydraulic type machine ( JT Toshi SVF-200/25-CPU ), the tensile speed being 8.3 x 10-6 m/s. Crystal structure and strain state in the tensile-deformed specimens were analyzed by X-ray diffraction method, where an X-ray apparatus ( Rigaku RINT 2400 ) was set-up as schematically illustrated in Fig. 4, Cu-Ka radiation, a graphite monochromator and a specimen rotation attachment being used.

T h i c k n e s s : 1.5 m m

Fig. 1.

Fig. 2.

Shape and dimension of the tensile specimen ( unit : mm ).

Schematic illustration of set-up for the X-ray diffraction measurement.

X-Ray Characterization of Shape Memory Alloy

101

RESULTS AND DISCUSSION Martensitic transformation behavior of the solution-treated specimen was first examined by a differential scanning calorimetry ( Shinku-Riko MTS9000 ). Only one DSC peak was clearly observed on both the cooling and heating curves. Similar DSC measurements for Ti-Ni-Cu alloys were reported by Shugo et al. [2] and Nam et al. [3]. According to the report [3], DSC peaks for the B19 ( orthorhombic ) ~ B19' ( monoclinic ) transformations are very diffuse and almost indiscernible, whereas those for the B2 +-~ B19 transformations are very sharp and clear. Therefore, the clear DSC peaks observed in the present work were concluded to be due to the B2 ~-~ B 19 transformations. Start and finish temperatures of the B2 ~ B 19 transformations were thus determined to be Ms' = 338 K, Mf' = 323 K, As' = 340 K and Af' = 356 K. However, it may not be denied that the B19 ~ B19' transformations were also induced in the solution-treated specimen, as will be verified later by X-ray diffraction. A typical X-ray diffraction pattern obtained from the solution-treated specimen is shown in Fig. 3. Many sharp diffraction peaks are observed, and these peaks are well indexed with

'

I

'

I

_.o 1 0 0 -

>

I

E

t~ t--

=-9

'

E

o~

50 -

O

E

t"q O

o~

'

-

E o

n,

UIO .

.

20

.

.

.,

E It-

~ O

~ O

l

30

40 50 Diffraction angle 20, deg

60

X-ray diffraction profile taken of the solution-treated alloy.

1000

~_ 800 600 400 200 0

Fig. 4.

'

ITIT=iN295 i41Cu8. K 51123ST ' I

~

Fig. 3.

I

0

20 40 Strain e, %

Stress - strain curve tested at 295 + 1 K.

60

70

102

Y. Kishi et aL

the standard diffraction peaks of orthorhombic B 19 Ti-Ni-Cu [4] and monoclinic B 19' Ti-Ni [5], as mentioned above. Stress - strain curve of the solution-treated specimen, which consists of the B 19 and B 19' martenstic phases, was recorded at 295 + 1 K, and it is shown in Fig. 4. The curve is able to be divided into four stages, an elastic deformation of existing martensites, an elongation due to the reorientation of those martensite variants and two stages of plastic deformation, although the boundaries between successive stages are not so clear. Saburi et al. [6] also reported a similar stress - strain curve for a Ti-40.5at.%Ni-10at.%Cu shape memory alloy, but the boundaries between successive stages on their curve were fairly clear compared to ours. Such difference between their and our curves is supposed to be caused by the difference in Cu content, because tensile properties of Ti-Ni-Cu alloys have been well known to be affected by Cu content.

Fig. 5. X-ray diffraction patterns and scanning electron micrographs of surface obtained from the tensile deformed specimens.

morphology

X-Ray Characterization of Shape Memory Alloy

(c) Fig. 5.

103

Applied stress = 700 MPa

Continued.

X-ray diffraction patterns and scanning electron micrographs of surface morphology were taken from three specimens, which were tensile-deformed under 100, 300, 700 MPa and so subjected to plastic deformation of an early and a later of the first stage and a middle of the second stage, respectively, and the patterns and SEM images are shown in Fig. 5. The peaks between 40 and 50 degree in 20 become broader with increasing the applied stress, as clearly known from a comparison with those in Fig. 3, and these broad peaks are overlapped over a wide range of angle. Intensities of diffraction peaks from the orthorhombic B19 martensite seem to decrease with increasing the applied stress value. In the X-ray diffraction pattern taken of the specimen deformed under 700 MPa, peaks between 40 and 50 degree in 20were completely overlapped so that respective peaks can not be distinguished at all. Diffraction peaks from a uniformly strained specimen are known to shift to lower or higher angle slightly. However, such a peak shift was not observed in the present X-ray diffraction patterns taken of tensile-deformed specimens. This seems to mean that the strain in the tensile-deformed Ti-Ni-Cu alloy is not uniform. Slight relieves, which seem to be due to the reorientation of martensite variants, are observed on the specimen surface tensile-deformed under 100 and 300 MPa, and the surface relief area on the specimen deformed under 300 MPa is larger than that under i 00 MPa. On the other hand, slip lines and micro-crack observed on the specimen surface deformed under 700 MPa. These surface relief morphologies observed by SEM seem to correspond to the peak broadening in X-ray diffraction patterns. Therefore, crystallographic strain state may be quantitatively known from the analysis of X-ray diffraction profiles as well as surface relieves, being now carried out.

CONCLUSIONS Tensile deformation behavior of the solution-treated Ti-Ni-Cu alloy was examined at 295 + 1 K, and X-ray diffraction analysis was performed for the tensile deformed alloy. The stress strain curve obtained was divided into four stages, an elastic deformation of existing martensites, an elongation due to the reorientation of martensite variants and two stages of plastic deformation. The X-ray diffraction peaks between 40 and 50 degree in 20 became -

104

Y. Kishi et al.

broader with increasing the applied stress, and these broad peaks are overlapped over a wide range of angle. Surface relief of the tensile deformed specimens was observed by a scanning electron microscope. Slight relieves, which seem to be due to the reorientation of martensite variants, are observed on the specimen surface deformed under 100 and 300 MPa, and slip lines and micro-cracks were observed on that deformed under 700 MPa. These surface relieves observed by SEM seem to correspond to the peak broadening in X-ray diffraction patterns. Therefore, crystallographic strain state may be quantitatively known from the analysis of X-ray diffraction profiles as well as surface relieves, being now carried out.

ACKNOWLEDGEMENT This research was partially supported by Grant-in-Aid for Encouragement of Young Scientists (A), 11750577, and Scientific Research (C) 12650663 and 12650703 of the Ministry of Education, Science, Sports and Culture, Japan.

REFERENCES 1. Asai, M., and Suzuki, Y. (2000) Proc. of the international Symposium and Exhibition on Shape Memory Materials ( SMM'99 ), 17. 2. Shugo, Y., Hasegawa, H. and Honma, T. (1981) Bull. Res. Inst. Mineral Dress. Metall., Tohoku Univ., 37, 79. 3. Nam, T. H., Saburi, T. and Shimizu, K. (1990) Materials Transactions, JIM, 31,959. 4. ICDD card No. 44-0114 5. ICDD card No. 27-0344 6. Saburi, T., Takagi, T., Nenno, S. and Koshino, K. (1989) Proc. of the MRS International Meeting on Advanced Materials, 9, 147.

Nondestructive Characterizationof Materials X Green et al. (Eds) (c)2001 Elsevier Science Ltd. All rights reserved.

105

A NEW THEORY OF X-RAY STRESS MEASUREMENT WITH ITS APPLICATIONS Masanori KURITA Department of Mechanical Engineering, Nagaoka University of Technology Nagaoka, 940-2188 Japan ABSTRACT The method of x-ray stress measurement can nondestructively measure residual stress in a localized thin surface layer of polycrystalline materials. This method is basically based on the theory of elasticity for isotropic materials. A new theory of x-ray stress measurement is proposed which is applicable to both isotropic and anisotropic (textured) materials. This theory is based only on the premise that the lattice strain varies in proportion to the stress without using any other theory of elasticity. This theory was applied to the measurement of the stress of ap-split material which has a nonlinear sin2ap diagram. KEYWORDS Residual stress measurement, X-ray diffraction, Theory of elasticity, Steels, Nondestructive testing, Experimental stress analysis.

INTRODUCTION The method of x-ray stress measurement is useful technique to nondestructively measure residual stress in a localized thin surface layer of polycrystalline materials. Basically, this method is based on the theory of elasticity for isotropic materials. Originally, the plane stress state is assumed in this method because the stress in thin surface layer is measured; in this case the variation in the peak position of a diffraction line profile as a function of sin2ap, called the sinZap diagram, is linear, where ap is the angle between the specimen normal and the diffraction plane normal. Later, the three-dimensional stress analysis for isotropic materials was reported by Dolle [1]. According to this theory, the existence of the shearing stress components -Cxz and r~yz in the direction of the specimen normal will produce a different values of the strain or the peak position on the + and -~p sides; this phenomenon is called the ap-split. In textured materials, it is well known that the peak position oscillates in the sinZap diagram [1-3]. A new theory of x-ray stress measurement is proposed; since this theory is based only on the premise that the strain varies in proportion to the stress without using any other theories of elasticity, it is applicable to any elastic materials irrespective of isotropic and anisotropic (textured) materials [2,3]. In the present paper, this theory was applied to a ap-split material which has a curved sin21p diagram.

106

M. K u r i t a

A NEW THEORY OF X-RAY STRESS MEASUREMENT

Let a straight line fitted to n peak positions p in the sin2ap diagram by the least squares method be, irrespective of the linearity of the diagram p - Mx + N (1) where the slope M and the intercept N of the straight line are given by M = ZaiP i (2) U = Zbip i (3) n

and Z denotes ._~ , and nxi --~i~i - (~1~7i)2

(4)

a i = n~,xi 2

~-'3('i2 -- Xi~l('i hi -~ n ~ c i 2 - (~-J('i

)2

(5)

= sine ~i The following equations hold for Eqs.(4) and (5). xi

~a i

= 0

(6)

Yb i

=1

(7)

For materials having a nonlinear sin2ap diagram, Eq.(1) represents average variation of the peak position or the strain with sin2ap. For elastic materials, irrespective of isotropic and anisotropic, the strain will vary in proportion to the stress, so that the peak position p i of a diffraction line with its change Api caused by applied and residual stresses is given by Pi = Po + l ~ i

"- Po + ki (era + ~ 0 )

(8)

where po is a peak position of a stress-free specimen, Oa and Cro are the applied and residual stresses, respectively, and ki is a coefficient depending on the angle ap, between the specimen normal and diffraction plane normal, i.e. the direction of the strain. Equation (8) is the only one premise used here in deriving the new theory. Substituting Eq.(8) into Eqs.(2) and (3) and using Eqs.(6) and (7), we obtain m

"- n ( o " a

"["O'0)

N = p,, + C(cr a + Cro)

(9)

(10)

where B -- Y a i k i

(11)

(12) Equations (9) and (10) show that the slope M and the intercept N of the straight line fitted to n peak positions in the sin2ap diagram by using the least squares method will vary in proportion to the stress even for materials having a nonlinear sin2ap diagram and that the stress can be determined from the slope M and the intercept N. A comparison of Eqs.(ll) and (12) with Eqs.(2) and (3), respectively, shows that B and C are the slope and the intercept of the ki-x (x = sin2~p ) diagram, respectively. Equation (9) shows that the value of B is also the slope of the M-Oa diagram. From Eq.(9), we obtain O"a "4" O"0 = K ' M (13) C - Ybik i

N e w Theory o f X-Ray Stress Measurement

107

where K ' is the stress constant given by K'=I/B (14) Equation (13) shows that the stress is proportional to the slope M of the straight line in the sin2xp diagram that is determined by the process described previously and that it can be determined from M regardless of the linearity of the diagram. The stress constant K' can be determined experimentally from the slope B of the M-oa diagram that is obtained from the slopes M for various applied stresses Oa. Substitution of Eqs.(9) and (10) into Eq.(1) gives p = B(O.. + o.o)X + Po + C(o.a "["O'0) (15) = (o.. + o.o)(Bx + C) + Po

Equation (15) shows that a set of straight lines in the sin2ap diagram obtained by applying various stresses intersect at a point given by C (--~, Po) (16)

THREE-DIMENSIONAL STRESS ANALYSIS FOR ISOTROPIC MATERIALS Three-dimensional stress analysis for isotropic materials in x-ray stress measurement was carried out by Dolle [1]. Rearranging the equation proposed by Dolle, the strain e ~ in the direction of OP in Fig.1 is derived from the theory of elasticity as I+V

s,~, =----~- (o.x - o.33)sin2 ~P + e33 1+~' + E (~ cosr + 023 sin r sin 2~p

(17)

3 0"33 E33

R

where o.x =o.11c~162 +0"12sin2r +o.22sin2r (18) E33 --~-1 [o.33 - v ( o . n + o.2z)]

(19)

and E and v are Young's modulus and Poison's ratio, respectively. Transforming the strain E,, in Eq.(17) to the 2 peak position p of a diffraction line, we obtain P - Po - c e ~ (20) where 360tan0o z c -= ~ (21) zc F i g . l Stress state on the surface of specimen. and 00 is the Bragg angle of a stress free specimen. Substitution of Eq.(17) into Eq.(20) gives 1 E (22) e33 + ~-(o',3 cosr + 123 sin r 2~p P = Po + ~(1 K,,o.x - O.33) sin 2 lp + ~ + K(1 1/) where K is the theoretical stress constant given by K--- 7d~cot0 o = E (23) 360(1 +v) cO+v )

108

M. Kurita

kj VALUE FOR ISOTROPIC MATERIALS For isotropic materials, the value of ki in Eq.(8) can be determined theoretically as follows. From Eq.(8) ki ._ Op

(24)

0o" a

When a stress Oa is applied to a specimen in the direction of the axis 1 in Fig.l, replacing Eqs.(18), (19) and (22) with Oal+ (ia, we obtain 4

ki -~ Op

1

_~ g(COS2 Csin2~p _

v__L_)

(Ill

in

(25)

a0. a l+v When a stress (ia is applied in the x direction in Fig.l, substituting ~ = 0 into Eq.(25) we obtain

l(sinZ~p_ v ) (26) ki = K l+v Equations (25) and (26) hold for both the two- and three-dimensional stress analyses. These equations show that the ki values determined by using peak positions at the same absolute values of xp on both sides will agree, so that each B and K' in Eqs.(ll) and (14) determined from peak positions of either side should agree; that is, the constant K' can be determined from peak positions on either side. From the same reasoning, C in Eq.(12) determined from peak positions on either side should agree. From Eq.(26), for (27) l+v ki equals zero and pi = po from Eq.(8), so that a set of straight lines in the sinZap diagram for various applied stresses (ia will intersect at a point given by sin 2/]3

__--- 'V

(v__L_, 1 +v

Po)

(28)

Since the point given by Eq.(28) corresponds to ki=O for isotropic materials, every straight lines determined by the least squares method in the sinZap diagram will pass the point in Eq.(28) irrespective of applied and residual stresses.

STRESS MEASUREMENT FOR Y-SPLIT MATERIAL Although the six stress components of ap-split materials can be totally determined from the equations given by Dolle [1], it takes a lot of time. For simplicity, the stress Ox to be measured can be determined as follows. Following the definition given by Dolle, let al be 1 1 e (29) al - -~(P§ + P-~,)- P0 +-~-(0.x - 0.33)sin21p + g(1+v'------~e33 where p+, and p., are the peak positions on the +ap and -ap sides, respectively. From Eq.(29), the slope m of the al- sin2ap diagram is given by 1 m -- --(0.x -0"33) (30) K Since the stress 033 in Eq.(30) is zero at the surface of a specimen and it is usually small in the thin surface layer measured by x-rays, the stress ox can approximately be obtained from 0"~ ~Km (31)

New Theory of X-Ray Stress Measurement

109

TEST PROCEDURES

Table 1

A specimen of the size of 101x22x5 mm was prepared from a structural rolled steel JIS type SS400. It was annealed at 9000(; for 1 h in vacuum and ground. Various stresses were applied to the specimens using a four points bending loading device and the stress at the surface of the specimen was measured by x-ray diffraction using the condition shown in Table 1.

Characteristic x-rays Diffraction plane Filter Tube voltage, current Preset time Irradiated area

TEST RESULTS AND DISCUSSIONS

Conditions of x-ray stress measurement.

Before l o a d i n g

Chromium Ka (211) plane Vanadium foil 30kV, 9mA 1s 5 x 15 mm 2 Perpendicular direction

156.6

Figure 2 shows the sin2ap diagram of the specimen before loading. In the direction "~ ~_..~" Grinding 156.4 perpendicular to the grinding direction, the .~ . r a : .... :on peak positions on the +ap and -ap sides almost :~ agreed. In the grinding direction, however, the 8 peak positions on each side were split as shown in Fig.2. Figures 3 and 4 show the ~ 156.2 9- r variation of the peak positions on the +ap and - 00 ~b>0 -~p sides, respectively, in the grinding direction 95% confidence with sin2ap for various applied stresses Oa. The 156 _ interval l , I , I , , I straight lines determined by the least squares 0 0.2 0.4 0.6 method intersect at a point as predicted by sin2~ Eqs.(16) and (28). Fig.2 Sin2~diagram before loading.

Applied stress 1 6 [ ~b_~ 0 55[-.

-48

9 150 o 203

g

~,

a a , MP~

f

i 156-1~ ~X~'~"~ [ 95%confidence \ x , , "-.t. interval k,,,,~ [ I Maximum ",%" 155.9[- I Minimum \ l l , I , I , I 0 0.2 0.4 0.6 sin2~ Fig.3 Sin2apdiagram on +ap side for various applied stresses Oa.

Applied stress

l-

/

0

156-51- ~ .

/~'~o

/

]_ " ~ - - ~ ~

qL$

9 48 . 99 o

" 150

o

~., 156.1 | 95% confidence interval [ I Maximum 155.9~ i I Mlmmum , I 0 0.2

~

!

'

0'6

0.4

sin2~

Fig.4 Sin2apdiagram on -ap side for various applied stresses Oa.

110

M. Kurita

Figure 5 shows the variation of the slopes M of the straight lines in the sinZap diagrams for the +ap and -ap sides shown in Figs.3 and 4 as a function of the applied stress Oa, that is, the M-Oa diagram. The slopes M for the +ap and -ap sides vary linearly with the applied stress Oa and the slopes B of the two straight lines in Fig.5 almost agree as predicted from the theory described previously. This shows that the stress constants K' determined from data points on either side will agree because Eq.(14) holds. Figure 6 shows the variation of the intercept N of the straight lines in the sin2ap diagram on the +ap and -ap sides in Figs.3 and 4 as a function of the applied stresses Oa. Similarly to the variation of the slope M shown in Fig.5, the intercepts N also varies linearly with the applied stress % and the slopes C of the two straight lines in Fig.6 almost agreed as described previous section. angle, deg 156.5 O

0

- ~"~ -0.2 - ~ - "L..... -0.4 _

-o,

~

O o

._q

-0.6 -0.8 -1

_

~ 95% confidence interval

" ~ ~I~.

_

Y

~

-.- ~b-_-0 - -.,- 4 , ~ 0

-6

30

~

0

-27

O

27

.

"~ 156.3

-39

o

0 .,-i r/l 0

_359

r 156.1

"'--L

95% confidence Gteivai I Maximum I Minimum

155.9

160 1~o 26o

Applied stress o a, MPa

I

3o

o

* 51

16o 1~o

Applied stress cr a, MPa Fig.5

M-% diagram on +ap and -ap sides. Fig.7 Variationof peak positionp with Oa for fixed ap angles. 2 X 10-3:

O

156.5 -~...

gh

_ Theoretical value f o r isotropic materials

/

'-I.../t

o

" ",. "' ..

"13

o~

--~ 156.4

0

.

-..,,,.

cr 0

~ 95% confidence interval -.- ~b_~ 0 ~ - ~b_~0

~o 156.3 !

o

3o

1~o

1~o 26o

Applied stress cr a, MPa Fig.6

N-oa diagram on +ap and -ap sides.

-1 -

~ 95% confidence interval 9

-2

_

. I

0

"--l.

~ < 0 ~b_~0 I

"

I

0.2

I

I

0.4

I

I

0;6

sin2r Fig.8 Variationof ki value with sin2ap.

New Theory of X-Ray Stress Measurement Table 2

111

95% confidence limits of measured stress constant K' (MPa).

ap_~O

ap~_O

-334 _ 20

-326 + 19

ap_~O and ap~_O -330 +_ 14

Figure 7 shows the variation of the peak position p at fixed ap angles as a function of applied stress Oa. The peak position varies linearly as predicted by Eq.(8). The ki value in Eq.(8), determined as the slope of the straight line in Fig.7, is shown in Fig.8 as a function of sin2ap. Figure 8 shows that the ki values on the +ap and -ap sides agreed and they fall on the theoretical straight line given by Eq.(26). For textured materials, the measured ki value oscillated around the theoretical line of Eq.(8) for isotropic materials [3]. Table 2 shows the 95% confidence limits [4] of the measured stress constants K' determined from the peak positions on the +% -ap and both sides. The three constants K' almost agreed as predicted by the theory in the previous section. The 95% confidence limits of the stress Ox calculated from Eq.(31) was 63 --- 7 MPa, while the stress components calculated from equations given by Dolle [1] were Ox(=O11)=61 MPa and 033=-2 MPa, showing that Eq.(31) holds. CONCLUSION A new theory of x-ray stress measurement is proposed which is applicable to elastic materials irrespective of the linearity of the sinZlp diagram. This theory was applied to a 1p-split material. A straight line was fitted to n peak positions p in the sin2ap diagram by using the least squares method. (1) The slope M and the intercept N of the straight line in the sin2ap diagram vary in proportion to the stress, and the stress can be determined from Eq.(13). (2) A set of straight lines in the sin21p diagram for various applied stresses Oa intersect at the point given by Eq.(16). For isotropic materials this point is given by Eq.(28). (3) For isotropic materials, the coefficient ki in Eq.(8) is given by Eqs.(25) and (26). (4) The stress constant K' needed for determining the stress is obtained from Eq.(14) as the reciprocal of the slope B of the straight line in the M-oa diagram. (5) The stress constants K' can be determined from peak positions on either (+ap and -ap) side. REFERENCES 1. 2. 3. 4.

Dolle, H., (1979). Journal of Applied Crystallography 12, 489. Kurita, M. and Saito, Y. (1997), JSME International Journal, Series A 40, 135. Kurita, M. and Sato, Y. (1998), Transactions of Japan Society of Mechanical Engineers, Series A 64, 1778 [in Japanese]. Kurita, M. (1989),Advances in X-Ray Analysis 32, 377.

This Page Intentionally Left Blank

Nondestructive Characterizationof Materials X Green et al. (Eds) 9 2001 Elsevier Science Ltd. All rights reserved.

113

EDDY CURRENT NONDESTRUCTIVE

TESTING OF WELD ZONE

USING UNIFORM EDDY CURRENT PROBE

K. KOYAMA, H. HOSHIKAWA and N. TANIYAMA Nihon University, College of Industrial Technology 1-2-1 Izumicho, Narashino, Chiba 275-8575, Japan

ABSTRACT The conventional eddy current testing using pancake probe can hardly detect flaws in the weld zone because the weld causes large noise.

The authors have tried to detect flaws in the weld zone

using the uniform eddy current probe that generates no lit~-offnoise in principle.

The experimental

results have indicated that the uniform eddy current probe can detect both parallel flaws and perpendicular flaws to the weld line with far less noise than the conventional eddy current pancake probe. KEYWORDS Eddy current testing, Weld inspection, Flaw detection, Uniform eddy currem probe, Differential surface tangential probe

INTRODUCTION Eddy current testing has the advantage of non-contact and fast test method over other nondestructive testing methods.

However, one of the disadvantages of eddy current testing is that it

tends to generate large noise by variations of many factors such as probe liit-off and electromagnetic characteristics of the test material.

As a result, when the conventional eddy current testing using

pancake probe is applied to maintenance inspection of the weld zone, it is difficult to detect flaws in the weld zone because of the noise generated by the shape change and electromagnetic characteristic change of the weld zone.

114

K. Koyama, H. Hoshikawa and N. Taniyama

The authors have developed a uniform eddy current probe [1-5]. The uniform eddy current probe is much less influenced by variations of the probe lift-off and the electromagnetic characteristics of the test material than the conventional pancake probe.

The authors improved the

uniform eddy current probe in order to detect the flaws at the weld zone. The new uniform eddy current probe consists of a wide tangential exciting coil induces uniform eddy current in the test material and differential tangential detector coils pick up the local variation of the eddy current by flaws.

The authors have tried the eddy current testing of the weld zone using the new uniform eddy

current probe in order to reduce the noise by the weld zone.

The experimental results have proved

that the uniform eddy current probe has a high signal-to-noise ratio compared to the conventional pancake probe when it is applied to detecting flaws in weld zone. EDDY CURRENT TESTING OF THE WELD ZONE The welded parts of material tend to have some deformations and variations of electromagnetic characteristics.

These variations cause large noise and make it difficult for flaws in the weld zone

to be detected by the conventional eddy current testing.

Therefore, the eddy current probe that does

not generate the noise from the weld zone is necessary for the maintenance inspection of the weld zone by eddy current testing. The authors have developed a new uniform eddy current probe in order to detect the flaws at the weld zone.

The uniform eddy current probe consists of a large wide tangential exciting coil and

differential tangential detector coils.

The large wide tangential exciting coil induces uniform eddy

current in the test material as shown in Figure 1. The uniform eddy current is induced perpendicularly to a flaw as shown in Figure 1 (a) in order to detect the flaw at high sensitivity. But when the uniform eddy current is induced parallel to a flaw as shown in Figure 1 (b), it is difficult to detect the flaw. Thus, the direction of the uniform eddy current has to be changed depending on the flaw direction.

- "

9

._,

n

flaw ._,

eddy current eddy current (a) Perpendicular to the flaw (b) parallel to the flaw Figure 1 Uniform eddy current induced in the test material

WeldInspection by Uniform Eddy Currentprobe

115

Figure 2 shows the two types of uniform eddy current probe that consist of a large wide tangential exciting coil and differential tangential detector coils. Figure 2 (a) shows the structure of the uniform eddy current probe Type 1 to detect the flaw parallel to the weld line. Figure 2 (b) shows the structure of the uniform eddy current probe Type 2 to detect the flaw perpendicular to the weld line. The exciting coil induces uniform eddy current in the test material and the differential tangential detector coils detects only local eddy current variations of the eddy current by flaws.

The

uniform eddy current probe is differential and liit-offnoise free. The uniform eddy current probe Type 1 is applied to inspect the weld zone inducing the eddy current perpendicular to the weld line as shown in Figure 3 (a). Two tangential detector coils are connected for the differential, thus the probe causes very little noise from the weld zone.

If there is

a flaw parallel to the weld line, the eddy current in the test material is perturbed locally by the flaw. The detector consisting of differential coils generates a flaw signal due to the local variation of the eddy current around the flaw. The uniform eddy current probe Type 2 is applied to inspect the weld zone inducing the eddy current parallel to the weld line as shown in Figure 3 (b). Two tangential detector coils are connected for the differential, thus the probe generates very little noise from the weld zone.

If there

is a flaw perpendicular to the weld line, the eddy current in the test material is perturbed locally by the flaw. The detector consisting of differential coils generates a flaw signal due to the local variation of the eddy current around the flaw. As a result, the uniform eddy current probe can detect flaws in weld zone with very little disrupting noise.

Figure 2

Structure of uniform eddy current probe

K. Koyama, H. Hoshikawa and N. Taniyama

116

Figure 3 Eddy current testing of weld zone by uniform eddy current probe

EXPERIMENTAL SETUP Eddy current testing of weld zone was conducted with stainless steel plates of SUS 304 welded by TIG welding. The thickness of the plate is 1.9mm and the weld width is about 5mm. The plate has slit flaws parallel and perpendicular to the weld line made by electric discharge machining. The size of flaw is 5mm in length, 0.2mm in width, and 75% or 90% depth to the plate thickness. The uniform eddy current probe Type 1 was arranged to induce uniform eddy current perpendicular to the weld line as shown in Figure 3 (a) in order to detect flaws parallel to the weld line. And the uniform eddy current probe Tyep2 was arranged to induce uniform eddy current parallel to the weld line as shown in Figure 3 (b) in order to detect flaws perpendicular to the weld line. The size of the exciting coil of uniform eddy current probe is 30mm in width, 40mm in length, and 30mm in height. The sizes of the differential tangential detector coils are lmm in width, 4mm in length, and 7mm in height. The differential tangential detector coils alone were also used as a differential surface probe. The test frequency of 70kHz was chosen so as to make the skin depth of eddy current equal to the plate thickness. EXPER/MENTAL RESULTS Figure 4 shows the three-dimensional display of absolute value of the eddy current signal obtained when the conventional surface pancake probe that is 10 mm in outer diameter scans over the weld zone.

Figure 4 (a) shows the signal by the flaw parallel to the weld line. Figure 4 (b)

shows the signal by the flaw perpendicular to the weld line. The figures indicate that the weld zone generates quite large noise and makes it difficult for the pancake probe to detect the flaw. The authors conclude that the large noise in the figure is caused by the probe 1LR-offdueto the weld zone

Weld Inspection by Uniform Eddy Currentprobe

117

Figure 4 Three-dimensional display of absolute value of the eddy current signal obtained by the conventional surface pancake probe deformation of the plate and by change of electromagnetic characteristics of the weld zone. Thus the conventional eddy current testing using staface pancake probe can hardly detect flaws in weld zone. The authors have tried the eddy current testing of weld zone using uniform eddy current probe in order to reduce the noise by weld zone. The uniform eddy current probe is differential, and lift-off noise flee. From these features, the following fact can be expected, when the uniform eddy current probe is applied to inspection of the weld zone, the noise by weld zone is small. Figure 5 shows the three-dimensional display of quadrature component of the eddy current signal obtained when the eddy current probes scan over the flaw parallel to the weld line. Figure 5 (a) shows the signal by differential tangential probe and Figure 5 (b) shows the signal by the uniform eddy current probe Typel. Figure 6 shows the three-dimensional display of quadrature component of the eddy current signal obtained when the eddy current probes scan over the flaw perpendicular to the weld line. Figure 6 (a) shows the signal by differential tangential probe and Figure 6 (b) shows the signal by the uniform eddy cun~nt probe Type2. These figures (b) show that the weld zone generates only low levels of noise and the large flaw signals are clearly seen. The authors conclude that the uniform eddy current probe does not suffer from large noise because of its features of a lift-off noise free and self-differential.

K. Koyama, H. Hoshikawa and N. Taniyama

118

Figure 5 Three-dimensional display of eddy current signal when probes scan over the flaw parallel to the weld line Signal amplitude

Figure 6 Three-dimensional display of eddy current signal when probes scan over the flaw perpendicular to the weld line

Weld Inspection by UniformEddy Currentprobe

119

Figure 7 shows patterns of the flaw signal and the weld noise when the eddy current probes scan over the flaw parallel to the weld line. Figure 7 (a) shows the signal by the differential probe. Figure 7 (b) shows the signal by the uniform eddy current probe Typel. Figure 8 shows patterns of the flaw signal and the weld noise when the probes scan over the flaw perpendicular to the weld line. Figure 8 (a) shows the signal by the differential probe. Figure 8 Co) shows the signal by the uniform eddy current probe Type2. These figures (b) show that the flaw signal is large enough to detect the flaw compared to the weld noise. Thus the uniform eddy current probe can detect flaws in the weld zone with high signal-to-noise ratio. I

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-

120

K. Koyama, H. Hoshikawa and N. Taniyama

Figure 9 shows the change of normalized flaw signal amplitude with respected to the tiff-off change. Figure 9 (a) shows the signal by the flaw parallel to the weld line and Figure 9 (b) shows the signal by the flaw perpendicular to the weld line. The proportion of the change of the flaw signal amplitude of the uniform eddy current probe is smaller than that of the differential probe. '

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CONCLUSION The experimental results have shown that the uniform eddy current probe can detect flaws parallel and perpendicular to the weld line with high signal-to-noise ratio compared to the conventional smface pancake probe. In this study, the test material is the stainless steel plate of non-ferromagnetic. Further study is needed to detect the flaws in ferromagnetic test material. REFERENCES 1. H.Hoshikawa and K.Koyama : "Eddy Current Testing by Uniform Eddy Current Probe", Proceedings of 4th Far East Conference on NDT, pp43-52 (1997) 2. H.Hoshikawa and K.Koyama : "Uniform Eddy Current Probe with Little Disrupting Noise", Review of Progress in Quantitative Nondestructive Testing, Vol.17, pp1059-1066 (1998) 3. H.Hoshikawa and K.Koyama : "A New Eddy Current Probe using Uniform Rotating Eddy Current", Materials Evaluation, Vol.56, No. 1, pp85-89 (1998) 4. H.Hoshikawa and K.Koyama : "Eddy Current Testing by Uniform eddy Current Probe", Proceeding of The 2nd Asian Joint Seminar on Applied Electromagnetics, pp 1-6 (1998) 5. H.Hoshikawa and K.Koyama : "Eddy Current Testing of Weld", Proceeding of The 2nd Japan-US Symposium on Advances in NDT, pp232-235 (1999)

NondestructiveCharacterizationof MaterialsX Green et al. (Eds) 9 2001 ElsevierScienceLtd. All rightsreserved.

121

A NEW EDDY CURRENT PROBE WITHOUT LIFT-OFF NOISE H. HOSHIKAWA, K. KOYAMA, and H. KARASAWA Nihon University Izumicho Narashino Chiba 275-8575, Japan

ABSTRACT The authors have devised a new surface eddy current probe that generates no lift-off noise in principle. The probe is lift-off noise free because it picks up the eddy current that is generated only by a flaw but not by the exciting coil and not by the lift-off of the probe from the test material. The minimal lift-off noise of the probe makes it possible for eddy current testing to utilize the phase of flaw signals to evaluate flaw depths. Thus the probe improves the quantitative evaluation of surface flaw depth by eddy current testing.

KEYWORDS Eddy current testing, Surface probe, Lift-off noise, Signal phase, Surface flaw depth

INTRODUCTION Eddy current testing has been used to detect surface flaws in metal products. The depth evaluation of surface flaws is very important because material breakage is connected with the depth rather than with their length and width. Thus eddy current testing has always been required to be more quantitative and reliable in evaluating depth of surface flaws. The conventional eddy current testing using the pancake coil probe suffers from the large noise by the variation of the probe lift-off from the test material [1 ]. Since the noise changes signal phase much, the signal phase can hardly be used to evaluate flaws. As a result, most eddy current testing by surface probes uses only the signal amplitude to evaluate flaws because the phase information is eliminated in the process of signal processing to suppress the lift-off noise. However, the signal amplitude changes not only by the depth of flaws but also

122

H. Hoshikawa, K. Koyama and 11. Karasawa

by the length and width. Consequently, eddy current testing by surface probe has not been considered as a quantitative method of evaluating the depth of surface flaws. The authors have devised a new surface probe for eddy current testing that generates no lift-off noise in principle. Experimental results have indicated that the probe provides the phase information on surface flaw depth and that the signal phase changes by the depth of flaws but not by the length and width. Thus the probe makes it possible for eddy current testing to evaluate flaw depth based on flaw signal phase without much influence from flaw length and width. The authors hope that the new probe utilizing flaw signal amplitude and phase will make the eddy current testing more quantitative and reliable to evaluate flaws than the conventional probes that only use the flaw signal amplitude.

CONVENTIONAL PANCAKE COIL PROBE Conventional pancake coil probes pick up the variation of the eddy current by flaws to detect them in test materials. Lots of work have been done on test coil impedance analysis and have contributed a great deal to the development of eddy current testing [2,3]. The lift-off of pancake coil probe from the test material reduces the eddy current in the material and also decreases the magnetic coupling between the pick-up coil and the eddy current. As a result, large noise caused by the lift-off variation is unavoidable so long as the probes pick up the change of the eddy current induced by the test coil or the test coil impedance. Since the large lift-off noise causes the phase of the probe signal to change a great deal, the signal phase can hardly have been utilized to evaluate flaws in the eddy current testing using the conventional surface probes. The only exception of using signal phase to evaluate flaw depth is the tube inspection by inner bobbin probe where probe lift-off does not change much. Thus the conventional eddy current testing using surface probes usually eliminates the phase information in the process of signal processing to suppress the lift-off noise. Thus only the flaw signal amplitude has been used to evaluate flaws by surface probe. However, since the signal amplitude changes not only by the flaw depth but also by flaw length and width, the conventional eddy current testing has not been considered as a quantitative and reliable method of evaluating depth of flaws.

A NEW EDDY CURRENT PROBE Lift-off noise is unavoidable so long as the probe picks up the eddy current induced by exciting coil. The authors have thought of two notions in order to design a new probe that generates no lift-off noise and picks up surface flaws.

A New Eddy Current Probe without Lift-off Noise

123

1) One of the methods to eliminate lift-off noise in eddy current testing is to develop a probe that picks up the eddy current generated only by flaws but not by exiting coil and the probe lift-off. 2) Each small part of detecting coil windings picks up the parallel component of eddy current to itself [4]. With the above two notions in mind, the authors have devised a new eddy current surface probe that is composed of a pancake exciting coil and a tangential detecting coil as shown in Figure 1. The circular exciting coil is adopted because it induces eddy current most efficiently in the test material. The exciting coil induces axi-symmetric circular eddy current in the test material with no eddy current circulating across the exciting coil circle when there is no flaw in the test material as shown in Figure 2. When there is a flaw crossing the circle, some eddy current circulates along the flaw. Since each small part of the detecting coil windings picks up the parallel eddy current component to the part, the tangential detecting coil picks up only the eddy current circulating across the circle as shown in Figure 3. As the new probe scans over a flaw, the detecting coil generates a signal depicting a figure eight like pattern. The new probe is lift-off noise free because the lift-off of the probe from the material does not cause any eddy current to circulate crossing the exciting coil circle. Thus lift-off noise can be eliminated by picking up only the newly generated eddy current by flaws and by not detecting the eddy current induced by the exciting coil when there is no flaw in the test material. The probe is self-nulling because the detecting coil generates a signal only when a flaw causes some eddy current to circulate across the circle.

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124

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Since the probe actually generates minimal lift-off noise, the authors have also thought that the probe lift-off does not influence much to the flaw signal and that the signal phase can be used for evaluating the depth of surface flaws. The feature makes the probe more reliable to estimate flaw depth than the conventional eddy current probes. The conventional probes estimate flaw depth based on the signal amplitude and the signal amplitude changes not only by flaw depth but also by flaw length and width. If the probe has two tangential detecting coils wound perpendicular to each other, it can detect all flaws in any orientation. The impedance of the exciting coil can also be used to monitor the probe lift-off in order to avoid the probe not detecting flaws in the material.

EXPERIMENTAL SETUP Figure 4 shows the sizes of the new probe and brass plates with an electric discharge machined slit flaw of different depths, lengths, and widths used for the experiments. The test frequency of 32 kHz has been chosen to make the skin depth of the eddy current induced in the material equal to the plate thickness. The exciting coil alone has also been used to conduct experiments by the traditional pancake coil probe.

EXPERIMENTAL RESULTS Figure 5 shows the experimental results of signals by front surface flaws and lift-off noises. Figure 5(a) indicates that the conventional pancake coil probe generates far larger lift-off noise than the flaw signals. On the other hand, Figure 5(b) indicates that the new probe generates far larger flaw signals than lift-off noise. Thus it is obvious that the new probe generates flaw signals with far higher signal to noise ratio than the conventional pancake coil probe. The result of the minimal lift-off noise shown in figure 5(b) indicates that the signal phase can be used to evaluate the depth of surface flaws.

A New Eddy Current Probe without Lift-off Noise

125

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Figure 6 shows signal pattems by front surface flaws and back surface flaws with different depths obtained by the new probe. The figure indicates that the amplitude and the phase of the flaw signals change depending on the depth of front surface flaws and back surface flaws. Figure 7 shows flaw signal pattems obtained by the new probe with different flaw lengths. The flaw length changes the amplitude of signals a lot but keeps the phase almost constant. Figure 8 shows flaw signal patterns obtained by the new probe with different flaw widths. Again, the flaw width changes the amplitude of signals a lot but keeps the phase almost

H. Hoshikawa, 1s Koyama and H. Karasawa

126

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constant. Figure 7 and Figure 8 indicate that it is not proper to use the signal amplitude in order to evaluate the depth of surface flaws in the eddy current testing. Figure 9 shows the signal pattems by front surface flaws and back surface flaws with the normalized amplitude. The figure indicates that the depth of front surface flaw lags the flaw signal phase and the depth of back surface flaw leads the phase.

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A New Eddy Current Probe without Lift-off Noise

127

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From the signal patterns shown in Figure 9, the authors have derived the flaw depth evaluation curve based on the signal phase as shown in Figure 10. Thus flaw depth can be evaluated by applying flaw signal phase to the curve in Figure 10 without much influence from the variations of flaw length and width. The relation between signal phase and flaw depth is just the same as the one used in the tube inspection by inner bobbin coil probe. The authors believe that the flaw evaluation method based on the signal phase improves the evaluation accuracy of flaw depth in eddy current testing.

Test frequency 932kHz, Flaw 9length 15mm, width 0.5mm I

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0 and overline indicates the complex conjugate. The function r is called mother wavelet which satisfies a certain mathematical condition. In this study, Gabor wavelet (or Morley wavelet) of the following form was employed: ~(t) = ~ lV - - ~w~ -exp

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where w0 and 7 (= 7r~/2/ln2 ~ 5.336 in this study) are constants. The parameter a is related to the angular frequency by w = 27r/a if w0 = 27r while the parameter b represents the time. Therefore, the wavelet transform provides a time-frequency distribution of the signal f(t). In this study, parameters a and b were discretized as a = 2 m/s and b = nat, respectively, where m and n are integers and At is the sampling interval of the signal f(t). When using Gabor wavelet, the integral in the definition of the wavelet transform can be evaluated effectively by using the algorithm of the fast Fourier transform (FFT). Kishimoto et al. [7] showed that the arrival time of each frequency component of a pulse can be determined by the wavelet transform of detected signal. Therefore, the time of flight of each frequency component of a pulse traveling between two locations can be determined by taking the wavelet transforms of two signals detected at these locations [8, 9] and, consequently, the wave velocity can be determined at each frequency. Note that the velocity determined by this technique is the group velocity. This technique has been applied to the analysis of Lamb or Rayleigh-like waves by, for example, Cho et al. [10], Futatsugi et al. [11], Wu and Chen [12]. In what follows, this technique is referred to as Method A. Another technique was developed to reduce the computational task for the wavelet transform. Consider two signals (say fl(t) and f2(t)) detected by two receivers at distances dl and d2 (> dl) from the transmitter, respectively. The impulse response function between these two can be identified by taking the Laplace transforms of f~(t) and f2(t), dividing the Laplace transform of f2(t) by that of f~(t), and taking its inverse Laplace transform. The Laplace transformation and inversion can be efficiently computed by using FFT [13]. Since the group delay at each frequency can be determined by the wavelet transform of the impulse response function, the time of flight between these two locations is readily determined. In practice, the impulse response function obtained from measured signals is often very noisy, the step response function obtained by integrating the impulse response function is used instead of the impulse response function. This technique will be called Method B. Figure 4(a) and (b) show dispersion relations of the a0 mode Lamb wave for 1-mm and 0.3-mm thick aluminum plates, respectively. Solid curve shows theoretical prediction based on the mechanical properties listed in Table 1. The measured group velocities coincide well with the predicted ones over a fairly wide range of frequency. There are few differences between the results obtained by Method A and Method B. It can be said that the group velocity can be determined successfully by both methods.

H. Inoue, K. Kishimoto and T. Shibuya

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(b) 0.3-mm thick aluminum plate (dl = 10 mm, d2 = 40 mm) Fig. 4. Dispersion relations of the a0 mode Lamb wave in aluminum plates obtained by Method A (left) and Method B (right). Figure 5(a), (b) and (c) show dispersion relations of the Rayleigh-like wave for 3-mm, 1-mm and 0.5-mm thick titanium coatings on steel substrate, respectively. Discrepancy between the measured and predicted results is more apparent than in the previous case. In addition, the discrepancy becomes more significant as the thickness of the coating decreases. A primary reason for this can be considered that the thickness of the coating was assumed to be uniform in the theoretical prediction whereas the actual interface between the coating and the substrate was wavy due to explosive welding. Discrepancy between the results obtained by Method A and Method B is less significant than the discrepancy between the measured and predicted results.

ESTIMATION OF ELASTIC CONSTANTS OF PLATES AND COATINGS The elastic constants of plates and coatings were estimated on the basis of the dispersion relation obtained by the wavelet transform. The group velocity of Lamb wave in the plate can be predicted theoretically if elastic constants, mass density and thickness of a plate are given. Therefore, if the mass density and thickness of the plate are known in advance, the elastic constants can be estimated by minimizing a functional

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i where ~9(coi; E, u) and cg(c0i) are predicted and measured group velocities at a discrete frequency wi, respectively. An iterative optimization technique with initial guesses for Young's modulus E [GPa] = 24p [g/cma] (see [14]) and Poisson's ratio u = 0.3 is applied to minimize the functional. The elastic constants of a coating can be also estimated in the same manner if elastic constants and mass density of the substrate are given. In this study, optimization

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150

H. Inoue, K. Kishimoto and T. Shibuya

Table 2. Results of estimation of elastic constants. Method A1 plate (1-mm thick) A1 plate (0.3-mm thick) Ti coating (3-mm thick) Ti coating (1-mm thick) Ti coating (0.5-mm thick)

A B A B A B A B A B

Young's modulus [GPa] Rel. Error (%) 65.7 -0.6 67.O +1.4 71.7 +0.7 67.4 5.3 120 +7.1 120 +7.1 107 -4.5 125 +12 95 15 94 16

Poisson's ratio Rel. Error (%) 0.29 -6.5 0.33 +6.5 0.44 +38 0.31 3.1 0.38 +19 0.38 +19 0.18 -44 0.51 +59 -0.04 113 -0.06 119

CONCLUSIONS In this paper, a new contact-type line-focused transducer has been developed for ultrasonic NDE of plates and coatings. It has been shown that the developed transducer can be used to excite and detect Lamb waves in plates or Rayleigh-like waves in coatings. It has been also shown that two techniques utilizing the wavelet transform are effective to determine the dispersion relation of the group velocity from experimentally detected signals. For further application of the developed transducer, it is necessary to investigate the characteristics of the transducer in detail.

REFERENCES 1. 2. 3. 4. 5. 6.

Achenbach, J.D. (2000). Int. J. Solids Struct. 37, 13. Monkhouse, R.S.C., Wilcox, P.D. and Cawley, P. (1997). Ultrasonics 35, 489. Castaings, M. and Hosten, B. (1998). Ultrasonics 36, 361. Zurn, B. and Mantell, S.C. (1999). Proc. SPIE 3589, 149. Yamanaka, K., Nagata, Y. and Koda, T. (1991). Appl. Phys. Left. 58, 1591. Nishino, H., Tsukahara, Y., Nagata, Y., Koda, T. and Yamanaka, K. (1993). Appl. Phys. Lett. 62, 2036. 7. Kishimoto, K., Inoue, H., Hamada, M. and Shibuya, T. (1995). Trans. ASME, J. Appl. Mech. 62, 841. 8. Inoue, H., Kishimoto, K. and Shibuya, T. (1996). Exp. Mech. 36, 212. 9. Inoue, H., Kishimoto, K., Nakanishi, T., Hori, J., Arai, M. and Shibuya, T. (1997). J. JSNDI 46, 206. 10. Cho, H., Ogawa, S. and Takemoto, M. (1996). NDT ~ E Int. 29, 301. 11. Futatsugi, T., Ogawa, S., Takemoto, M., Yanaka, M. and Tsukahara, Y. (1996). NDT E Int. 29, 307. 12. Wu, T.-T. and Chen, Y.-Y. (1999). Trans. ASME, J. Appl. Mech. 66, 507. 13. Inoue, H., Kamibayashi, M., Kishimoto, K., Shibuya, T. and Koizumi, T. (1992). JSME Int. J. 35, Set. I, 319. 14. Rogers, W.P. (1995). Res. Nondestr. Eval. 6, 185.

Nondestructive Characterizationof Materials X Green et al. (Eds) 9 2001 Elsevier Science Ltd. All rights reserved.

151

ADVANCED ON-LINE LAMB WAVE INSPECTION SYSTEM USING REAL-TIME SSP TECHNIQUE Y. NAGATA, H. YAMADA, Y. KONNO and S. NAITO

Plant Engineering & Technology Center, Nippon Steel Corporation 20-1 Futtu-shi, Chiba 293-8511, Japan

ABSTRACT In order to improve the defect detection capability of the Lamb wave inspection system conventionally used for on-line detection of internal defects of steel strip, the application of SSP (split spectrum processing) to Lamb waves and the optimization of the SSP parameters were investigated. Sample strips with artificial defects were used to evaluate the performance of SSP in this research. As a result, the enhancement of the detection capability of the inspection system was confirmed by optimum selection of SSP parameters. Furthermore, in order to realize real-time SSP calculations, the number of SSP calculations was drastically reduced. The newly developed Lamb wave inspection system can implement real-time SSP and other tasks such as image processing by using parallel data processing. The system can perform at a repetitive frequency of 500 Hz which is a sufficient frequency for the on-line Lamb wave inspection system. KEYWORDS Ultrasonic testing, Lamb wave, signal processing, non-linear processing.

INTRODUCTION The Lamb wave inspection system using one or two tire-type probes is conventionally used for on-line detection of internal defects (bubbles and inclusions) of steel strip. This is a system for detecting the presence of defects by projecting ultrasonic waves obliquely into strip, thereby generating Lamb waves in the transverse direction of the strip and detecting the signals reflected from the defects using one or two probes which are sealed in tires [1]. The on-line Lamb wave inspection systems now in use detect defects using narrow-band toneburst waves at frequencies of several MHz. Such narrow-band waves are used to eliminate the influence of the velocity dispersiveness of Lamb waves by using only Lamb waves of some specific mode and frequency so as to increase the detection capability. The detection capability of the on-line system is largely influenced by the grain echo, which is caused by the reflection from the grain boundary, and the electric noise. The electric noise can be removed by appropriate hardware or software filters, but the grain echo is a characteristic noise which is produced when an ultrasonic wave propagates in steel. Since it uses narrow-band

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Y. Nagata et al.

toneburst waves, the Lamb wave inspection system can considerably reduce both grain echo and electric noise by the use of band-pass filters which pass only the bands for received signals. Recently, however, there has been an increasing need for detection of finer defects by improving the detection capability of the system, and the development of a new technique is indispensable. Conceivable means of improving the detection capability are: (1) increasing the frequency of transmitting waves, (2) detecting defects by the use of two or more Lamb wave modes at the same time and (3) improving the detection capability by signal processing. Regarding (1), attenuation due to propagation remains a large actual problem partly because the Lamb wave propagation distance increases to about 3 m, for example, in the case of detection for the whole width of strip using a single tire-type probe. Regarding (2), it was reported that the detection capability for each of several Lamb wave modes was analyzed by numerical simulations assuming the depth and size of defects in steel and also experimentally verified, and that it was possible to improve the detection capability by the use of several modes [2,3]. This method is promising indeed, but in order to adopt it in the on-line inspection system, cost and other problems should be cleared because the use of several modes requires the paralleling of the probes. Therefore, if the detection capability can be improved by signal processing such as SSP technique, which had been reported as an effective means of removing grain echo, it is possible to commercially use the advanced on-line ultrasonic inspection system at relatively low cost. Minimization algorithm (Min) [4,5], polarity thresholding algorithm (PT) [6,7,8], a combination of both (Min+PT) [9] and geometric mean filtering (GM) [10] have been proposed as actual techniques for SSP. Their evaluations to ultrasonic waveforms and theoretical analysis of their effectiveness have been conducted. Also, studies on the sophistication of SSP have been conducted recently, for example, by modeling the grain echo and studying the optimization of the SSP techniques using the models [11,12]. Furthermore, studies have been made on a method of grain echo reduction by shifting the position of the ultrasonic probe, thereby collecting several waveforms and nonlinearly processing them in the same manner as SSP and on the application of this method [13,14]. Although very interesting, this method involves the drawback that the hardware and cost must be increased to realize its commercial on-line application because it requires the mechanism of shifting and paralleling the probe positions and the function of collecting a lot of data. For the application of SSP to the lamb wave inspection system it is necessary to clarify the questions of whether the effectiveness of SSP can be obtained if noise having the same frequency as narrow-band Lamb waves is generated, and how real-time SSP can be realized, since no report had ever been made on this matter.

SIGNAL PROCESSING ALGORITHMS In the SSP techniques, n waveform data r/t)(j=l,...,n) are first obtained by passing ultrasonic wave received signal r(t) through n filters which are adjoined by the pass bands as shown in Fig. 1, and these data are subjected to nonlinear processing for rj(z-) at time z- to obtain final output y ( ~ ) . Eqs. (1)-(3) present three types of processing as typical methods of nonlinear processing. In the case of echoes from defects, the probability is high so that n waveform data rj(~) are the same in phase and all the waveforms are positive or negative. On the other hand, in the case of noise signals such as grain echo, the probability is high so that the data are various in phase and both positive and negative waveforms coexist. The principle of SSP is the removal

Real-Time SSP

153

of noise by utilizing this difference. As SSP parameters there are the number of filters n, each filter band b, the center frequency of the first filter ]'1 and center frequency interval/9/as shown in Fig. 1. The optimization of these SSP parameters must be studied. (1) Min (1)

y ( r ) = rk ( r ) where Irk(r) I = rain{ b ( r)l, j=l,"',n} (2) PT y(z') - r(z") if all values of rj(z) are positive or negative y(r) = 0 in other cases (3) Min+PT y ( r ) = rk ( r ) if all values of rj(r) are positive or negative y(z') = 0 in other cases where Irk( r)l = min{ b ( r)l, j=l,"',n)

(2)

(3)

This section explains signal processing in the time domain, not in the frequency domain, as real-time SSP will be described later. First, the impulse response of the n filters adjoined by the pass bands is expressed by Eq. (4) [15]. In this equation, which assumes Gaussian-type band pass filters, b is the filter band, .~ is the center frequency of the jth filter and fj=fi+(j-1)D s.

h~(t) = 2x/-~b e -( '*~

rcfjO

(4)

Actually, digital signals which are made discrete by sampling time T, are processed. Therefore, signals after passing through each filter are expressed by Eq. (5), and they are subjected to nonlinear processing as shown by Eqs. (1)-(3) to obtain the final output [15]. Also, hi(t) in Eq. (4) is represented in the finite time domain, and the function is sampled by sampling time T, so that it has discrete data from the -Lth to the Lth. L

r,(nTs ) = i.~_ 2xf-~bTse-r ~b,r, )2COS(2 rcf , iT s )r[(n - i)Ts )1

(5)

EXPERIMENTAL RESULTS OF SSP APPLICATION TO LAMB WAVES The on-line Lamb wave inspection system used in this experiment performed defect detection using toneburst waves at about 2.25 MHz, and the effectiveness of the application of SSP to the system was studied. It was so planned that this experimental system was capable of taking out Lamb wave signals, trigger signals which correspond to transmitting waves generation, etc., from the existing inspection equipment as shown in Fig. 2. After making an 8-bit AD conversion of the waveforms, the digital data were saved in the personal computer and then SSP was performed. Since toneburst waves at about 2.25 MHz were used for detection, sampling was performed at 20 MHz. First, optimization of the SSP parameters was studied, that is, the number of filters n, each filter band b, the center frequency of the first filter fl and center frequency interval D I. Boring holes 0.3 mm in diameter, as artificial defects, through strip 830 mm wide, 490 mm long and 1.5 mm

154

Y. Nagata et al.

Transducer spectrum

~I

/ .........r(t): / _k"" r~e.ived / D//" si~..\ J / ~

fi ..fk! ~,, r,(t) Non-linear processing

/f~i ki(t) -f~i ~

, ,,

~

(t):

existingequipment

Lamb wave II inspection [[system Ipr~

experimentalsystem

~ signal//~ i PR~signa~ ~ AD converter I I "7 (20MHz) [ [ [

Generated/receivedI

Processed

I

| I ii I T~,~m~ ~ . - : : V e /

output

~Steell ...............I 1 ~ ~

I

SP Data

save

r"B:Bandwidth ,-1 Fig. 1. Processing flow of SSP technique.

Fig. 2. Block diagram of experimental system.

thick, the waveforms were saved in the personal computer when the signal-to-noise ratio (SNR) of the raw signals was 0.6 in the case where the distance between the probe and the hole was 565 mm. The SNR was defined as the defect signal amplitude/maximum noise amplitude in the range excluding the dead zones which correspond to the strip edges. The dead zone on the transmitting side was set at 80/is to remove the waves reflected from the tire-steel contact surface and the dead zone at the other edge was set at 20/zs to remove the reflected waves from the edge. Fig. 3 shows the results of the SNR improvement using the nonlinear SSP shown by each of Eqs. (1)-(3) by changing both the number of filters and the number of filter coefficients. Particularly, the results shown in Fig. 3(b) for Min and PT techniques by changing the number of filters proved to show almost the same trend as the results of the experiment and the theoretical analysis already reported in the past regarding the improvement of SNR by SSP [5,6]. Based on the results in Fig. 3, the Min+PT technique which is high in SNR improvement, the number of filters n=10 and the number of coefficients of each filter m = 2L+1 = 500 was adopted in the experiments described follows. The frequency band for SSP was decided based on the results of the frequency characteristics of the received waveforms shown in Fig. 4. This was in the case where holes 0.5 mm in diameter were bored, as artificial defects, through strip 830 mm wide, 490 mm long and 3.0 mm thick and the distance between the probe and each hole was 620 mm. Fig. 5 shows the frequency characteristics of defect echo waveforms (500520/zs in Fig. 4) and the noise waveforms obviously reflected from grain-boundary (for example, 280-300/zs in Fig. 4). As seen from Fig. 5, the narrow-band Lamb waves obviously fall on the frequency band of the grain echo, making it difficult to remove the noise by simple band-pass filters. Based on Fig. 3 and Fig. 5, the Lamb wave transmitting frequency range is determined to cover from 2.1 to 2.4 MHz with 10 filters and b = 4/9/is adopted for the relationship between each filter band b and center frequency interval/9/ [4,5]. After determining the SSP parameters, a sample strip, the 830 mm wide, 490 mm long and 1.5 mm thick, having bored holes 0.5 mm in diameter at positions 50, 40, 30 and 20 mm from one edge of the strip was used. A tire-type Lamb wave transmitting and receiving probe was set at a position 150 mm from the other edge as shown in Fig. 2, and then Lamb waves were received by rolling the tire-type probe. This experiment was conducted installing the sample strip in a calibration place provided at the side of the on-line system. Fig. 6 shows the detected waveforms and the processed waveforms for the artificial defect 40 mm from the strip edge. The raw waveforms are displayed after being subjected to rectification, and the processed waveforms are displayed after being subjected to SSP and further to rectification.

155

Real-Time SSP SNR of the 3 processed 2.5

signal

2.5

SNR of the processed 2

,~ .....

2

sign,~

;~-" "'-.'k:

- .........

'

1

0.,5

k,~a......-

4.~, ~:,,_.'-..

d

0.5 0

0

200 400 600 number of filer coefficients

800

0

I0

20 30 number of filters

(a)

40

(b)

Fig. 3. (a) SNR after SSP in the case of the number of filter coefficients changing, (b) SNR after SSP in the case of the number of filters changing. (Solid lines represent PT, dashed lines represent Min+PT and dash-dot lines represent Min.) Edge

Defect 0.2

R,.,

0.15

~ 2.0

":'

-=0.05 ~-o,

,~

~ 1.5

9

-o.15 -0.2

0

-100

200

300

400

500

0.0

0.0

Fig. 4. Typical detected Lamb waves.

'~mrll ' ' ]]~v: p, raw signal

ltlltl.

I

Ampfltude

loo

(arbitrary

,

9

!l! '|

,

2o0

~lJ

~

zoo

400

Time(ps)

t!11'

0.5

'

Processed signal

il

'

L,I.

500

6oo

elect

1 ~

oLIll"l I~ ...... ~,~..,....,..,~...a.,..,. .... ,.ill[ 0 lOO 200 300 4oo Time(~)

1.0

2.0

3.0

4.0

5.0

Frequency(MHz)

6.0

Fig. 5. FFT spectrum of the defect echo and the noise signal included in the Fig.4 waveform.(solid line: defect signal, dashed line" noise signal)

edge signal i[

1.0 0.5

Time( ~

unit) [HN§ 1

...................................................................................

3.5

~ 3.0

",-- o

.~
,, c ,,i,,,a

C

~,

U

,

~,

100

I

200

Position (Pixel)

I

I'-"

. . . . .

300

I"

1 O0

Intensity distribution

,

~_A

I

I

. . . . . .

I

200

Position (Pixel)

I

300

Deconvoluted waveform

Fig. 5 Deconvolution procedure of intensity distribution To reconstruct the original waveform Y(t) from the intensity distribution D(t), the analytical deconvolution procedure of Jacobi method [11] shown in Fig.4 is employed after measuring the stroboscope flashing waveform H(t)using a photo diode.

In this procedure, an unknown function

Y0(t) is assumed for D(t) and a tentative D'(t) is calculated by the convolution of Y0(t)* H(t). After comparing estimated D'(t) with D(t), Y0(t) is modified to Y~(t) by adding a function proportional to the difference between D(t) and D'(t) and the same procedure is repeated.

If the

diffeience between estimated D'(t) and D(t) becomes small enough, estimated Yn(t) is considered to be Y(t).

The accelerate coefficient k in the deconvolution procedure shown in Fig.4

was chosen adequate value below 1 for the convergence of the procedure.

An example of the

waveform analysis from a visualized image is shown in Fig.5. The contrast in visualized image was improved as seen by comparing the waveforms before and

Photoelastic Visualization of Ultrasonics after the procedure.

163

Note that the original sound pressure is proportional to the square root of the

intensity as seen in eq. (4).

Recovery of the original polarity The visualized image loses the polarity because of eq.(4).

In the field of the static photoelastic

imaging, a loading technique is generally applied to determine a stress state for a fringe pattern [ 12]. This technique is also effective for a photoelastic ultrasonic visualization [6].

The principle of this

technique for a longitudinal wave transmission is shown in Fig.6.

Fig.6 Principle of polarization recovery of waveform When a compressional load is applied to the specimen with a dead load, the intensity of the compressional stress part is increased and on the contrary that of the tensional stress part is decreased.

We applied a dead load of 2 kg on the specimen to examine the polarity of the original

waveform.

The applied load dependence of the intensity in the visualized image is shown in Fig.7. ,.

1 -

,okg ..... 2kg

E 0 Z v

c-

04O

60

80 100 Position (Pixel)

120

140

Fig. 7 Intensity distribution of visualized image depending on applied load An increase and a decrease of the intensity are observed at B and A, respectively.

From this

experiment we can decide A as a tensional stress part and B as a compressional stress part.

In all

measurements of visualized images, the dead load is applied and the polarity is recovered for longitudinal wave.

As for a shear wave measurement, the loading direction must be changed to be

45 degree from the ultrasonic transmission direction. An example of the waveform reconstruction process with the above two procedures is shown in Fig.8.

164

T. Mihara, T. Musashi and K. Yamanaka

....1 E

2

0

~ e=

_.E

0

a 0

b

-

100

2()0

Position (pixel) Distribution of Intensity

"~~ 9

0

J

. . . . . I

0

100

200

- " Compressional part =

+

+ " Tensional part

0

+ "

0

o

!

i

lOO

2

+

~ - 7 ....... , 200

-2

|

0

i

lOO

i

2oo

!

Reconstructed Wave Shape Fig.8 Reconstruction process of waveform from visualized image

EXPERIMENT AND RESULTS Shape and dimensions of the glass specimen and the experimental setup are shown in Fig.9. The narrow-band commercial probe of 5 MHz in frequency was used.

Ultrasonic visualized

images before (position 1 and 2) and after (position 3 and 4) the back wall reflection for a longitudinal wave were measured and the intensity distribution of each image was reconstructed

Fig. 9 Model specimen for visualization and displacement measurement by laser interferometer using the above mentioned procedures.

To confirm the availability of the waveform

165

Photoelastic Visualization o f Ultrasonics

reconstruction process, a laser interferometer was applied to measure the displacement at the back wall surface of the specimen.

The surface of the back wall glass specimen was coated with

evaporated Ag film to measure the absolute displacement accurately. A received RF echo also measured by the probe was compared with others.

The experimental results for the narrow-band 5

MHz probe are shown in Fig. 10. 0.5-

+

-

~

o ......

....,

N

Position(~)

Position( 9

-

,_,

-

-0.5 0

!

!

I

2

3

4

_

~, o 5 -

5

Position(~)

o

o-

.~-~

1

I

I

I

2

3

4

' _

5

Position@

--/

_

~

0

-0.5 I 2

i 0

1

I 3

I 4

l ~ r - - - - - - - ~ - - - - - - - - - ~

5

0

1

2

4O

E

4

5

80

zo

"2 E

3

Time ( tz s)

Time ( U s) (a) Reconstructed waveform ~,

60

>

4

o

~. -zo -40

0

I

I

I

I

1

2

3

4

5

0

1

I

I

I

2

3

4

5

Time (/z s) Time ( U s) (b) Displacement by laser interferometer (c) RF signal by probel Fig. 10 Comparison among reconstructed waveform, displacement waveform and recieved signal by the probe (5MHz, narrow-band transducer) First four waveforms (a) show the reconstructed intensity distribution from visualized image of each traveling position in the specimen.

Only the phase inversion occurred at a reflection at the

back wall and other signal shape remained almost the same and was independent of ultrasonic traveling.

Figures 10(b) and (c) show the time dependence of the surface displacement by the

laser interferometer and the RF echo received by the probe. position and area for these techniques were different.

Note that the effective measurement

That is, the laser interferometer detect the

back-surface displacement of 1 mm in diameter, the RF echo by probe detect the average sound pressure of 20mm in diameter at specimen surface and the intensity of visualized image is the projection of two dimensional sound pressure across the specimen thickness.

Comparing (a)~,/

166

T. Mihara, T. Musashi and K. Yamanaka

with (b), it is found that the reconstructed waveform resembles the surface displacement though a slight difference existed.

On the contrary, the RF signal measured by the probe shows a difference

with an increased number of carriers due to resonance of the probe.

These results show that the

developed reconstruction process using the visualized images is useful for accurate waveform analysis at any point during the ultrasonic transmission.

The reconstructed waveform in this study

still contained some noise shown in Fig.10 (a) due to the limited dynamic range in the 8 bit CCD camera, the A/D converter in image processor and brightness of stroboscope.

Though the image

averaging process up to 100 times was employed in this experiment for reducing the noise in image, other procedure are also required for more accurate analysis in future.

CONCLUSIONS A new algorithm for deconvolution of ultrasonic visualized images using the Jacobi method combined

with the experimental polarization determination was developed for accurate

reconstruction of ultrasonic waveform.

Comparing the surface displacement by a laser

interferometer with the reconstructed waveform from visualized image, a good agreement was obtained. This result shows that the developed process is useful for accurate experimental waveform analysis at any point during the ultrasonic transmission.

We also applied this process to the

waveform analysis of other types of transducers and to the waveform analysis of reflection echo from artificial flaw tip.

About these results, we would like to discuss elsewhere[ 13].

REFERENCES 1. K. G. Hall and P. G. Farley, Brit. J. NDT, (1978), 171 2. H. P. Rossmanith and J.W. Dally, Strain, (1983), 7 3. C. F. Ying, S. Y. Zhang and J. Z. Shen, J. Nondest. Eval., 4, (1984), 65 4. M. M. Marsh, Research Tech. in Nondestructive Testing, (Ed. by R. S. Shape, Academic Press, London and New York Vol.2, (1973), 333 5. K. Negishi, Jpn. J. Appl. Phys., 23, (1984), 23 6. K. Date and H. Shimada, Jpn. J. Appl. Phys., 27, (1988), 206 7. K. Date and Y. Udagawa, Jpn. J. Appl. Phys., 28, (1989), 197 8. T. Mihara, M. Obata and H. Shimada, SEM 6th International Congress, (1988), 221 9. T. Mihara, K. Hagiwara, and T. Furukawa, Jpn. J. Appl. Phys., 37,5B (1998), 3030 10. T. Mihara, Y. Otsuka, H. Cho and K. Yamanaka, Nondestructive characterization of materials IX (1999), 168 11. H. C. Burger and P. H. V. Cittert, Zeitschrift fur Physik, 79, (1932), 722 12. E. G. Coker, Phil. Mag. 20, (1910), 740 13. T.Mihara, T. Musashi, K.Yamanaka, (to be published)

Nondestructive Characterization of Materials X Green et al. (Eds) Published by Elsevier Science Ltd., 2001

167

NONDESTRUCTIVE TEST FIELD SURVEY FOR ASSESSING THE EXTENT OF ETTRINGITE-RELATED DAMAGE IN CONCRETE BRIDGES R. A. Livingston

Office of Infrastructure R&D, Federal Highway Administration, McLean VA 22101 USA A.M. Amde

Civil Engineering Dept., University of Maryland, College Park MD USA ABSTRACT A specific type of microcracking deterioration of concrete is associated with the formation of the mineral ettringite. It is hypothesized that this damage is caused by high potassium content in the Portland cement. To test this hypothesis, a survey of bridges is planned using nondestructive characterization methods. The concentration and distribution of potassium in the concrete will be measured by autoradiography of natural 4~ using storage phosphor imaging plates. The extent of microcracking will be evaluated using a modification of the impact-echo ultrasonic method to measure the log decrement or attenuation properties. The occurrence of this damage may not be random; instead it is thought to be a function of the type of Portland cement manufacturing process, the age of the structure and the method of casting the concrete (pre-cast vs cast-in-place). To control for these factors, it wil! be necessary to develop a stratified random sampling design, based on data in the National Bridge Inventory. KEYWORDS Concrete, ettringite, potassium, autoradiography, impact-echo, bridges, stratified sampling INTRODUCTION Recently there has been concern over the potential for damage to concrete from micro-cracking deterioration associated with the presence of the calcium aluminum sulfate mineral ettringite (3CaO 'A1203 "3CaSO4"32H20 ). A typical case is illustrated in Fig. 1. However, many aspects of the process remain controversial[l]. Ettringite that forms at early times, on the order of hours, after the mixing of the concrete, is considered beneficial because it prevents false setting. Normally, after several more hours, the ettringite then breaks down into another sulfate phase, monosulfate ( 3CaO AlzO3"CaSO4" 9 12H20 ). However, in cores from severely damaged concrete several years old, ettringite is observed growing in the cracks and air voids. This is the basis for the name "delayed ettringite formation" or DEF for this type of damage. Among the unresolved issues is the role of the ettringite itself in the damage process. Another expansive phase, alkali-silica reaction (ASR) gel, is often found in combination with the ettringite, leading some researchers to propose that the ASR gel is the real agent of damage, and hence the ettringite is only a side product. This distinction is important because of its implications for cement manufacturers vs mixers of concrete. If ettringite is the actual agent, then the cement-making process may have to be changed. On the other hand, if ASR is the problem, then the choice of aggregates and

168

R.A. Livingston and A.M. Amde

Fig. 1: Pier supporting a bridge on the Washington DC Beltway. Arrow points to typical network of cracks. The inset shows the characteristic features of the cracks. the concrete mixing and placement processes may have to be altered. Since this issue remains unresolved, the term used here is DEF-associated damage. Another issue is the mechanism that controls the timing of the DEF. Several processes have been proposed, including the presence of sequestered sulfate phases within silicate phases in the cement. It is assumed here that the chemical reaction takes the form: MONOSULFATE + 2Ca 2++ 2SO42- + 20H20 ~ ETTRINGITE Hence the only processes that can shift the equilibrium from monosulfate to ettringite are the addition of Ca 2+, SO42- or H20. However, in this case, the concrete is a closed system with respect to the first two components, which leaves water from the atmosphere as the only possible factor. Nevertheless, different batches of concrete exposed to the same atmospheric conditions show different rates of expansion, which indicates that other factors beside environmental conditions must be taken into account. As discussed below, the working hypothesis is that the critical factor is the potassium content of the cement. A further issue is the role of curing temperature in the DEF process. The first cases of DEF appeared in concrete railroad ties that had been steam-cured at up to 80 ~ and some researchers maintain that such high temperatures are necessary to initiate the reaction. Elevated temperatures can occur in precast concrete temperatures, but would not be expected in structures that are cast in place. A final issue is the geographic extent and the severity of the DEF-associated damage. One view holds that the problem is localized to a few sites where some batches of cement from a misoperating plant were used. Others have observed characteristic damage across the country. Questions have also been raised concerning the amount of damage on a given structure, since, as shown in Fig.l, it typically has a patchy distribution, rather than being uniformly distributed throughout.

NDT Survey of Ettringite-Related Damage in Concrete

169

POTASSIUM EFFECTS The evidence for potassium as a significant factor in the damage process comes from several sources. Field studies have indicated a significant correlation between degree of distress and potassium content [2]. Laboratory studies have also shown correlations between potassium content and expansive damage [3]. The specific mechanism of damage has not yet been confirmed. The working hypothesis for this study is that the hygroscopic nature of potassium salts causes the problem. Potassium is found in unhydrated Portland cement primarily in the form of potassium sulfate [4]. It usually comes from illitic clay used as a raw material in the manufacture of the cement. Once the cement is mixed with water, the potassium sulfate undergoes an exchange reaction with calcium and hydroxide ions to produce a dilute potassium hydroxide solution and calcium sulfate: K2SO4 + Ca 2++ 2OH- -+ 2K + + 2 O H + CaSO 4 t Over time the solution reacts with the atmosphere to produce potassium carbonate: 2K + + CO3= ~ K2CO3 which is has a very low critical relative humidity, for deliquescence (RH = 43 %). Consequently, under prevailing temperate climate conditions, the compound would typically be in a saturated solution rather than in solid form. This can explain why the crack network usually appears wet, or even dripping with water, as in illustrated in Fig. 1. The frequent exposure to water would in turn drive the equilibrium in Equation 1 toward ettringite. It would also promote ASR if reactive aggregates are present. Thus the timing of DEF can depend on the rate of carbonation of the concrete and the number of relative humidity cycles.

SURVEY OBJECTIVES AND METHODS The planned field survey has the main objective of collecting data on the correlation between potassium content and degree of microcracking using the NDT methods described below. It would also provide information on the frequency of occurrence and the age distribution. In addition to the two main NDT methods, visual inspection and digital imaging would also be used. The latter permits the application of digital image processing techniques such as spatial Fourier transform to automate the detection and quantification of damage. Finally, a limited number of cores would have to be taken by drilling for petrographic analysis for the presence of ettringite and/or ASR gel. IMPACT-ECHO FOR DISTRIBUTED DAMAGE In the basic impact-echo method, a transient stress wave is launched into the test object by means of a mechanical impact on the surface, which is made by tapping a small (usually less than 15 mm in diameter) steel ball against a concrete or masonry surface. The stress pulse generated by the impact propagates into the test object as spherical compression (P) and shear (S) waves, as well along the surface as a Rayleigh (R), or surface wave. The waves are reflected by internal flaws, or interfaces, and by the external boundaries of the object. A displacement transducer placed near the impact point is used to monitor the reflected stress waves. Previous applications of the impact-echo method have generally focused on the detection of a few discrete flaws in concrete structures. In contrast, the distributed damage mechanisms such as DEF produce micro- cracking widely distributed in concrete. The presence of many, small cracks in concrete structnres causes scattering of the propagating stress waves, rather than causing prominent reflections of these waves. Thus to quantify the extent of micro-cracking, the amount of stress wave

170

R.A. L i v i n g s t o n a n d A . M . A m d e

scattering must be determined. The scattering results in the attenuation of the signal over successive reflections. Consequently, the attenuation can be, used as a parameter for scattering and hence for degree of microcracking. Figure 2 shows a typical waveform from a test slab of concrete. [5]. The amplitude is measured at the peaks associated with the resonant frequency. These data are then fitted to an exponential decay model: xl0

-4

I t = I o exp(-c~t)

4.

where It is the intensity at time t, Io is the initial stress wave intensity and a~is a decay constant. Thus a is a measure of attenuation. Kesner et al.[6] have evaluated this method in the laboratory on a concrete specimen undergoing accelerated microcracking damage. The degree of micro-cracking was quantified by digital image analysis of neutron radiographs.

i .......

ii:i :ii :ii i: iii: i

-4 0

0.5

1

i 1.5

2

2.5

Time (milliseconds)

Fig. 2: Filtered impact-echo waveform from damaged concrete slab. Arrows indicate peaks used to calculate decay constant

Neutron radiography uses neutrons instead of X-rays to produce 2-dimensional images of 3-dimensional objects. In the case of concrete, neutrons have the advantage of higher resolution than X-rays for imaging cracks when a suitable contrast agent such as gadolinium is applied [7]. A typical neutron radiograph of microcracked concrete is presented in Fig. 3.

Fig. 3: Neutron radiograph of microcracked concrete. Arrows indicate typical microcracks. Crack density is 0.6%.

NDT Survey of Ettringite-Related Damage in Concrete

171

In this study, cores from the damaged concrete slab were removed and cut into 3.8 mm slices which were then treated with a gadolinium nitrate solution. The slices were then irradiated at the TRIGA Mark II fiuclear reactor at Cornell U. with a neutron flux of approximately lxl 06 neutrons/cm2-s. The resulting neutron radiographs were scanned, and the digital image processed to highlight the cracks as shown in Fig. 3. The area of each crack was determined in units of pixels, and the crack density was calculated ratio of the sum of the crack pixels to the total number of pixels in the image [6]. The impact- echo test results obtained on the laboratory slab specimen showed an excellent correlation with the crack densities as measured by the neutron radiographs. Table I compares the results obtained from the neutron radiographs with the corresponding decay constants calculated from impact-echo signals for the 130 mm thick laboratory specimen with a P-wave speed of 3920 rn/sec and an impact duration of 30 gsec. The results presented in Table 1 are also consistent with results obtained from numerical simulations of impact-echo tests on 130 rnm thick plates.

Table I: Damage Classification Using Impact-Echo and Neutron Radiography Data

Decay Constant

Crack Density

Less than 3100

< 0.5%

3100 to 4000

0.5% to 1%

4000 to 5000

1% to 2%

Greater than 5000

> 2%

POTASSIUM AUTORADIOGRAPHY Since the potassium content of the concrete may be a significant factor, it is important to be able to measure this nondestructively. Such a method has been developed that uses autoradiography of the natural radioactivity of the element [8]. The particular radioisotope is 4~ which has an abundance of 0.0117%. It decays primarily by emitting a beta particle with a maximum energy of 1.314 MeV. Alternatively it also decays by electron capture roughly 11% of the time which yields a gamma ray with energy of 1.461 MeV. The rate of decay is slow, with a half-life of 1.26 X 109 years. The conventional method for producing autoradiographic images has used the X-ray film, but this application uses electronic imaging plates or storage photostimulable phosphor (SPP) plates [9]. The operating principle involves the transfer of energy from the radiation to the storage phosphor, which is usually a BaFBr scintillator doped with europium. This energy remains trapped until it is read out by stimulating the phosphor with a scanning laser beam, which results in the release of the energy as photons of light. These are then detected by a photomultiplier tube. The result is a digital image with a pixel size as small as 50 gm with a dynamic range > 105. The image plates, which come in sizes up to 35 x 45 cm, are reusable.

172

R.A. Livingston and A.M. Amde

The 4~ autoradiography method was demonstrated on four cores taken from precast concrete bridge beams [8]. The cores, which were supplied by the Texas Department of Transportation, came in pairs. One pair was taken from a beam that was in sound condition, the other pair from a beam that showed significant cracking from DEF-associated expansion. The cores were exposed simply by standing them on end on an 20 x 50 cm SPP imaging plate (Fuji BAS-5000)* at the National Institute of Standards and Technology in Gaithersburg, MD. The image, produced primarily by beta particles, after an exposure of 24 hours is given in Fig. 4. All four cores produced recognizable patterns. By visual inspection, the images of two of the cores ( Nos. 1 & 4) appear darker than the others, indicating a greater radioactivity and hence a higher potassium content. This is confirmed quantitatively by summing the pixel signal level (PSL) over the area of each of the images. After correction for background, these two samples have radioactivity count rates roughly 50% higher than the other two. Direct visual inspection shows that the high count rate samples appear to have a greater degree of cracking than the other two. This implies that the damage is related to the amount of potassium present. In addition to estimating the bulk average potassium content per sample, the SPP image can provide further information by examination of the spatial distribution of the radioactivity. Analysis of the data shows Fig. 4: Autoradiography of cores from concrete box beam clusters of values that can be girders used to distinguish between regions of cement paste and the aggregates. Also, the distribution of the 4~ is not uniform. A qualitative visual comparison of the 4~ distribution Versus the position of cracks and mottled areas on the cores suggests that they are correlated. This would be consistent with a redistribution of the potassium, which would be initially uniformly distributed in the cement matrix, after experiencing a number of relative humidity cycles..

SAMPLING DESIGN The survey is not intended to examine all bridges on the highway system, which number on the order of 560,000. Instead, it will investigate a statistically significant sample of this population for frequency of DEF. However, this type of damage is not equally probable for all bridges. In the first place, a major proportion of them are not constructed of concrete. Of those that are, the incidence of DEFassociated damage may vary significantly with age of construction. This reflects historical changes in Portland cement chemistry and physical properties [ 10] as well as in practices of mixing the concrete. This is complicated by the fact that the number of bridges constructed per year has varied considerably especially during the decades when the Interstate Highway System was being built. Therefore, the probability of finding a bridge with DEF-associated damage would involve a multinomial distribution rather than a simple binomial one.

*Commercial names are provided for information only. No product endorsement is implied.

NDT Survey of Ettringite-Related Damage in Concrete

173

Efficient use of limited resources for survey requires that a stratified sampling method be used. This will be developed using age and construction-type data available from the National Bridge Inventory. Figure 5 presents a possible tree diagram for developing such a plan. Even with only three levels: type of casting (2), decade of construction (4) and potassium level (3), 24 categories would have to be sampled. In this case, the potassium level would have to estimated from data on Portland cement analyses performed by individual state DOT materials laboratories. This may not always be available.

INVENTORY

PRECAST 1960s 1970s 1980s 1990s LOK

CAST

IN PLACE

!960S~ 970S 1980S 1990S

MIDK HIK

Fig. 5: Preliminary tree diagram for stratified random sampling plan. K = potassium.

CONCLUSIONS Nondestructive methods for characterizing DEF-associated damage in concrete have been developed to quantify the amount of microcracking and the associated potassium level. These can be combined in a field survey to investigate the relationship between the two variables. However, these methods do not measure directly either the ettringite content or the ASR gel. This can only be done by taking cores and performing a petrographic analysis. Nevertheless, the NDT survey can be used to identify the most significant locatiens for taking cores, thereby minimizing the amount of destructive testing required. It should be noted that petrographic analysis can often be ambiguous about the cause and effect relationship between DEF and ASR. Ultimately, it may not be necessary to distinguish between the two, because they are both influenced by the potassium level, which may be the critical factor in this type of damage. Finally, given the large number (over 550,000) and diversity of bridges in the US, it is not possible to survey the entire population. Instead a stratified random sampling approach is required.

R.A. Livingston and A.M. Amde

174 REFERENCES

10.

Taylor, H. F. W. (1997). Cement Chemistry, 2nd Edition, Thomas Telford, London. Gress, D. (1997). Early Distress in Concrete Pavements. FHWA-SA-97-045, Federal Highway Administration, Washington, DC. Ramadan, E. O., Amde, A. M., and Livingston, R. A. (2000) ACIMat. J., (In press). Miller, F. M., and Tang, F. J. (1996) Cem. Contr. Res., 26(12), 1821. Kesner, K. (!998). MS Thesis, Cornel! lJ., Ithaca, NY. Kesner, K., Sansalone, M., and Poston, R. W. (1998). FHWA Conference on Nondestructive Testing of Concrete." SPIE Vol. 3400, San Antonio. Najjar, W., Aderhold, H., and Hover, K. (1986) Cem. Concr. Agg., Winter, 103. Livingston, R. A., Aderhold, H. C., Hobbs, S. V., Hover, K. C., and Cheng, Y. T. (2000) Cem. Concr. Agg., 22(1), 37. Cheng, Y. T., Soodprasert, T., and Hutchinson, J. M. R. (1996) App. Rad. Isotop., 47(9/10), 1023. Neville, A. M. (1998). Properties of Concrete, John Wiley & Sons, New York.

Nondestructive Characterization of Materials X Green et al. (Eds) 9 2001 Elsevier Science Ltd. All rights reserved.

175

CHARACTERIZATION OF MICROSTRUCTURE AND INTERFACIAL PROPERTIES OF ADVANCED CERAMIC COMPOSITES BY ULTRASONICS, SCANNING ACOUSTIC MICROSCOPY AND RAMAN SPECTROSCOPY M. H. M A N G H N A N I , M. P R A S A D , P. V. ZININ, Y. W A N G , J. B A L O G H S. K. DEB, R. L E M O R

Hawaii Institute of Geophysics and Planetology, University of Hawaii, Honolulu, Hawaii 96822, USA ABSTRACT The variations in the elastic properties (modulus and impedance contrast) and microstructure in three types of SiC(Nicalon)fiber/SiCmatrixcontinuous ceramic fiber composites (CCFC) with different fiber orientations have been characterized using the Ernst Leitz scanning acoustic microscope (ELSAM) at - 0.8 - 1.0 GHz frequency, with a resolution of about 1.5 - 2.0 lam. The technique, especially in conjunction with micro-Raman scattering, allows us to image and study the variations in the elastic properties, impedance contrast, and crystal-chemical characteristics of the fibers, coatings and matrix, leading to a better understanding of the interfacial properties, interfacial microstructure and porosity in the CFCC. KEYWORDS Continuous fiber-reinforced ceramic composites (CFCC), acoustic microscopy, Raman spectroscopy, INTRODUCTION Microstructure and interfacial properties play important roles in the performance of advanced ceramic composites. Characterization of these properties should therefore lead to better understanding of not only the performance, evaluation and modeling of such composites, but also the fabrication processes. The application of scanning acoustic microscopy (SAM) for microstructure characterization of ceramic composites is well documented [1]. The main objective of this paper is to demonstrate that by combining the three experimental techniques, namely the scanning acoustic microscopy, ultrasonic pulse transmission technique and microRaman spectroscopy, unique opportunity is provided for studing the complex microstructure of the CFCC and for correlating between microstructure and their elastic properties. TECHNIQUES

A. Scanning Acoustic Microscopy The high-frequency acoustic microscope, with a spatial resolution of about 1 ktm, is the acoustic equivalent of an optical microscope [1]. In the acoustic microscope a monochromatic acoustical signal is focused onto the specimen through the sapphire rod and the couplant (usually water) between the lens and the specimen. Acoustic images are obtained when the acoustic microscope mechanically scans the specimen in a plane parallel to the specimen surface. Variation of the mechanical properties with depth can be studied by scanning the specimen at various lens positions. The acoustic wave propagates through the coupling fluid to the specimen surface and is reflected by the surface of the specimen with the reflectance function R(O). When the acoustic wave is defocused into the specimen by a distance z, it experiences a phase delay of 2zkcosO. The total output signal V(z) received at the transducer is a function of the depth of defocus z, and is calculated by integrating the wavefront over the surface of the transducer, P ( O ) R ( O ) e -iEzkc~176sin0cos0d0

V(z) = ; 0

(1)

176

M.H.Manghnanietal.

where cx is the half-aperture of the lens, k is wave number, P(O) is the combined pupil function, and R(O)is the reflection coefficient. The contrast in the acoustic microscope represents the variation of the acoustical properties of the specimen and is dependent on the defocusing distance z. In the present study we used an Ernst Leitz scanning acoustic microscope (ELSAM) with a high frequency lens (about 1 GHz) to investigate fine structure of the interfaces and a low frequency acoustic microscope (SAM-50) operating at frequency 50 MHz to investigate fatigue damage in the fiber composite.

V(z) Transducer

~

. ~ ,, , Sapphlrelrod

', ,

'

I(2

1' D, IE?F Defocus z j

[

Solid

\Tf)_~XZ ",,.

Fig. 1. Schematic geometry of the defocused acoustic lens. BE is the trajectory of specular wave. ADFC is the trajectory of leaky Rayleigh surface acoustic wave. In the ray model the leaky Rayleigh wave is excited by ray AD, striking the surface at the angle OR = Vw/Vn. Here Vw is the velocity of the longitudinal wave in coupling liquid and Vn is the velocity of Rayleigh wave. B. R a m a n Spectroscopy

Spectra were obtained accros the fibers and interfaces in selected CFCC specimens, using a Dilor XY spectrometer with confocal aperture and an Olympus microscope (50x). The excitation source was an Ar § laser operating at 514.5nm with a power of 100 mW; approximately 2 mW was incident on the specimen. The spot size of the laser was about 2 ~xn in diameter. The Raman signal was collected using a liquid nitrogen cooled CCD detector (EG&G 1530 C) integrating over 200 to 300 s. C. Ultrasonic Pulse Transmission Technique

Compressional and shear wave velocities (Vp, Vs, respectively) measurements were made with a commercially available transducer (5 MHz for Vp and 1.0 MHz for Vs) with fused quartz buffer rods as the delay line. Vp was measured in one propagation direction with two different directions of wave polarization, in order to fully evaluate elastic anisotropy of the CCFC.

Continuous Ceramic Fiber Composites

177

SPECIMENS Elastic properties of 7 SiC(Nicalon)r,ber/SiCmatrix CCFC specimens obtained from DupontLanxide Composites (DLC) (Table 1) were investigated.

Table 1. Description and densities of CFCC specimens investigated. CFCC specimen 1 2 3 4 5 6 7

Type Standard Enhanced Enhanced Thermally treated Enhanced Enhanced Fatigue tested

Bulk density (kg/m 3) 2220 2320 2190 2457 Low density High density N/A

RESULTS A. Ultrasonic measurements The velocities measured by ultrasonic pulse transmission technique and calculated elastic moduli of the CFCC specimens are presented in Table 2. We first assume that fibers are oriented randomly in the 1-2 plane with respect to 1-2-3 orthogonal coordinate system. This would mean that the elastic properties of the specimens are uniform in the 1-2 plane, but different in the direction of axis 3 (axis 3 is directed perpendicular to the fibers), and the CCFC specimens may be assumed to be transversely isotropic. The longitudinal wave propagating in the plane containing the fibers is denoted as Vpl, and the one propagating in the direction normal to the fibers is designated as Vp3. Vsl is the velocity of the shear wave traveling in the 1-2 plane a with polarization direction perpendicular to the fiber orientation. Vsz denotes the shear wave with polarization direction that is parallel to the fiber orientation. All four specimens display transverse isotropy, with minimum velocity values measured at 90 ~ to fiber orientation. Results are shown in Table 2. Table 2. Measured longitudinal Vp, shear velocityVs, and the calculated elastic moduli.

Specimen Number

Vm (m/s)

Vp3 (m/s)

Vsl (m/s)

Vs2 (m/s)

Ell

C33

C44

C66

C12

(GPa)

(GPa)

(GPa)

(GPa)

(GPa)

1

8529

5208

2520

4023

161.4

60.21

14.0

35.92

89.6

2

7995

6199

3136

3871

148.2

89.15

22.8

34.76

78.6

3

8073

2938

3859

142.7

67.46

18.9

32.6

77.5

4

8867

5555 0 8124

3133

4192

193.41

162.36

24.2

43.2

108.

As seen in Table 2, the velocity anisotropy for both Vp and Vs is significantly high in all the four specimens, ranging from approximately 10 to 45% for Vp and approximately 21 to 55% for Vs, and can be attributed to the alignment of fibers and the elastic anisotropy of the SiC fibers themselves. The two longitudinal velocities (Vp1 and Vr,2) measured in the plane containing the axes of the fibers are nearly equal, that is, the specimen is isotropic in that plane (transverse isotropy). A transversely isotropic material has five independent elastic constants. In the co-ordinate system used in the present work, with axis 3 being the axis of symmetry, the constants extracted from the model using the experimental measurements as input are Cll (= C22 ), C33, C13 (- C23 - C31 = C32 ) and C44 (- C55). The fifth independent constant, the in-plane shear modulus, is C66-0.5(CllC12). Table 2 shows elastic moduli calculated from the experimental data. It is interesting to note that the range of values of the elastic moduli Cll and C33 of "thermally treated" specimen (exposed to Pittsburgh 8 coal ash for 1500 hours at 2300 F), specimen #4, is significantly higher than the values for both the standard and "enhanced" CCFC specimens. The moduli Cll and C33 determine

M.H. Manghnani et al.

178

the bulk velocities normal to the fiber (Cll = pVpl 2) and parallel to the fibers (C33 = pVp32), respectively. In contrast, the moduli C44 (pVs12) and C66 (pVs22) do not show dependence on the density and the synthesis process of composites. Differences between the various synthesis processes are clearly seen in the anisotropic modulus values among the four specimens: standard (specimen #1), enhanced (specimens # 2, 3), and heat treated (specimen #4). The enhancment does not change individual fiber property, Cll; however, it does change the interfacial properties markedly. The change is seen by the C33 values, measured in direction perpendicular to the fiber orientation, which are higher in the enhanced specimens. Heat treatment enhances composite the strength (moduli) of CCFCs in both directions. Also, the moduli are higher than the values measured in standard and enhanced specimens in both directions. Heat treatment also appears to increase the strength of the composite. This is likely to be a result of coalesence due to removal of coating. In effect, there is a significant decrease in the porosity and anisotropy in specimen # 4. Figure 2 shows the elastic moduli versus density for the four specimens.

#4

200 160

j

#1

180

e7 ~

#3

140

9

e

011

C33

#2

r,3 120 _= 100-

-~ o

80 60

40

v~v

A~...A~A.

20 2.15

'

I

2.20

'

I

2.25

'

~

v~ I

2.30

'

v

~ 66

A C l

2.35

'

I

2.40

'

I

2.45

'

44 I

2.50

Density (g/cm 3) Figure 2. Elastic moduli of CCFC specimens as a function of density. B. Microstructure of CCFC by Acoustic Microscopy Microstructure of the CCFC was evaluated with a scanning acoustic microscope operating at 0.81.0 GHz. Image resolution obtained at this frequency is 1.5-2 ~ n in water, which was used as a coupling medium. Interfacial bonds between fibers, between fibers and matrix, matrix porosity, and fiber configurations were mapped with this technique and compared to the ultrasonic pulse transmission results. Figures 3 and 4 show well-oriented fibers and matrix (very light shade) of a standard CCFC specimen. The impedance contrast between fibers and matrix indicates the matrix has slightly higher modulus. The darker area commonly around the fibers and less often between the fibers (in matrix) indicates porous (debonded) coating and lower elastic modulus or porosity, respectively. The two concentric darkest zones around the fibers indicate lower modulus of the coating (i.e. graphite), compared to that of the core of SiC fiber, and some possible porosity or debonding at the two interfaces (Fig. 4).

Continuous Ceramic Fiber Composites

Figure 3. Surface SAM image (312x312 ktrn) at 1 GHz of a standard CCFC specimen (#1).

Figure 5. Surface SAM image (312 x312ktm) at 1 GHz of an enhanced CCFC specimen (#2).

179

Figure 4. Surface SAM image (100x100 ~trn) at I GHz, showing closer view of the coating of the fibers (specimen # 1).

Figure 6. Surface SAM image (312x312 ~xrn) at 1 GHz of an enhanced CCFC specimen (#3).

Compared to the standard specimen (Fig. 3), the impedance contrast in the enhanced CCFC specimens (Figs. 5 and 6) is noticeably different. The contrast between fibers and matrix indicates that the matrix has slightly lower modulus. Microporosity in the matrix is seen as black dots; and fiber-plucks are observed as black areas between the intact fibers (Figs. 5 and 6). The fibers in heattreated CCFC specimen appear to be very well bonded and their interfaces are well coalesced (Fig. 7). The graphitic coating on the fibers is extremely thin or absent. Porosity is negligible and hence the density is much higher than the "standard" and "enhanced" CCFC. Impedance contrast between fibers and matrix is small: yet it clearly shows that the matrix has higher elastic modulus than that of fibers as in Figs. 7 and 8. The impedance contrast (Fig. 8) and its implications are the same as discussed above under Figs. 3 and 4. Black areas represent fiber-plucks Acoustical contrast in the low density, enhanced CCFC specimens is very similar to enhanced CFFC specimens (#2,#3), except that the low-density specimen (#5) contains two types of matrix distinctly observed in both acoustical (Fig.9) and optical images (Fig. 10). Chemical composition of the components will be discussed in the section describing Raman scattering data. The high-density specimen contains only one type of matrix. The difference in the contrast between matrix and fibers reveals (Fig 9) that matrix is less rigid than the fibers.

180

M.H. Manghnani et al.

Figure 7. Surface SAM image (312x312~xn) of heat-treated CCFC specimen (#4) made at 1 GHz.

Figure 9. Surface SAM image (lxl mm) of low density, enhanced CCFC specimen (#5) made at 1 GHz.

Figure 8. Surface SAM image (312x312~n) of heat-treated CCFC specimen (#4) made at 1 GHz.

Figure 10. Optical image of the same specimen (#5). Field of view is 700x700 ~n.

The boundaries of the fibers of "high density" specimen #6 (Fig. 11 and 12) are eroded as compared to the boundaries of "low density" specimens #5 (Fig.9). This erosion is more clearly seen in the acoustical image (Fig. 11) than in the optical one (Fig. 12). ACOUSTIC MICROSCOPY OF DAMAGED FIBER SPECIMEN

The damage in a DuPont-Lanxide CFFC specimen produced by mechanical fatigue was inspected by KSI 50 acoustic microscope. The low frequency acoustic microscope was chosen to study the internal fatigue damage in specimens without any preparation for SAM investigation. A timeresolved image of the damaged area of the fiber CFCC bar (Fig. 13 and 14) exhibits a strong reflection from the internal crack, besides the reflection from the surface.

Continuous Ceramic Fiber Composites

Figure 11. Surface image (312x312gm) of high density, enhanced CCFC specimen (#6) made at 1 GHz.

181

Figure 12. Optical image of the same specimen (#6), high density, enhanced. Field of view is 700x700 ~tm.

In the time-resolved microscopy the imaging of the structure at different depths from the surface can be done by choosing appropriate gate positions. Figures 13 and 14 show the images of the composite bar taken from top and middle surfaces.

Figure 13. SAM image of the top surface of the test CFCC bar at 50 MHz. (scale: 30 m m x 30mm.

Figure 14. SAM image of the surface of the middle part of the test CFCC bar at 50 MHz. (scale: 30 mm x 30 mm.

Figure 13 shows the texture of the fiber bundles. No surface-breaking cracks have been observed. Image of the middle surface (Fig. 14) reveals the edge of the internal crack propagating parallel to specimen surface. RAMAN SPECTROSCOPY Raman spectra show that the fibers appear to be similar in the CFCC specimens studied. The matrix is heterogeneous, consisting of nanocrystalline and crystalline silicon carbide. The silicon carbide crystallines are observed to be in the m-phase. Matrices of two types (see Figs. 9 and 10) can be distinguished in low density, enhanced specimen #5: low modulus (dark shade) and high modulus (light shade). Micro-Raman measurements were conducted in order to identify these two matrices and to investigate the crystal-chemical characteristics of these matrices.

M.H. Manghnani et al.

182

The spectra shown as Fig. 15 (a) and (b) are from ample #5 in the corresponding region arc shown Fig. 9 (a) the white portion and (b) from the black portion. Both the spectra show sharp transverse optical (TO) and longitudinal optical (LO) bands of SiC and their Raman shift frequencies (cm -1) correspond to those for the 3C-SiC (13-SIC) phase. The small width of the bands corresponds to crystallite sizes > 1 ~xrn. The shoulder on the low frequency side of the 795 cm -1 line shows existence of 6H-SiC polytype in both the white and black portions. In addition to these sharp lines of SiC, there are broad features over the range 300 - 600 cm -1 and also between the TO and LO bands. These are typical of disorder known to be present in SiC matrix due to the presence of oxygen. The bands at 1350 cm -1 and 1600 cm -~ seen in the fibers are, however, absent in the Raman spectrum of the matrix. The different appearance of the light region (a) and dark region (b) in the optical image may be due to presence of pores in the region (b).

r

,.-., (/)

\ (a)

~.

k

t.--

:D

< t/) t.-(D

_=

r

200

4O0

0O0

(r

t ,..~.,...:~i

000

1000

Raman shift (cm -1)

t200

Figure 15. Raman spectra of low density, enhanced CCFC matrix of the specimen (#5) taken at point (a) and (b) Fig. 9.

200

400

1500

Raman

800

I D00

1200

1400

1600

shi~t ( crn ')

Figure 16. Raman spectra of low density, enhanced CCFC specimen (#5) taken at fibers (a), (b) and (c) (acoustical image Fig. 9).

Three spectra shown Fig. 16, a, b, c are recorded for the regions of specimen #5 marked (a), (b) and (c) in the lower part of the acoustical image (Fig. 9); (a) and (c) have been recorded from the center of the two nearby fibers and are identical. The spectrum (b) in Fig. 16 is obtained from the fiber region close to the fiber edge and exhibits quite different features. The SiC bands at 784 cm -1 (TO) and 969 cm -1 (LO) can be clearly seen. The strong band near 1355 cm -1 and 1600 cm -1 can actually be decomposed into three components at 1352 cm 1, 1500 cm -1 and 1590 cm -1. The contribution of the band at 1500 cm -1 is however, much smaller than the other two bands. Those three bands can be assigned to sp 3 bonded carbon at 1352 cm -1, sp 2 bonded carbon with heteroatoms (especially O) at 1500 cm -1, and pure sp 2 bonded carbon at 1600 cm -1 respectively. The absence of SiC lines indicates amorphous and glassy nature of the SiC in the fibers. Further the Raman scattering cross-section of SiC is much smaller than carbon and hence is not seen in the spectra. The. amorphous nature is also seen in the large bandwidth (170 cm -1 and 90 cm -1 for 1352 cm -1 and 1600 cm -1 , respectively). Asymmetry of the TO band indicates presence of SiC in the 6H-SiC form. The presence of these sharp bands indicates existence of fairly large size (> 10 nm) crystallites.

REFERENCES 1.

Briggs, A., Acoustic Microscopy. 1992, Oxford: Clarendon Press.

THIN FILMS AND COATINGS

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Nondestructive Characterizationof Materials X Green et al. (Eds) 9 2001 Elsevier Science Ltd. All rights reserved.

185

ULTRASONIC EVALUATION OF REMELTED ZONE THICKNESS IN ALUMINUM ALLOY CASTINGS

Guoxin Jiang ~, Hiroshi Katol, Yuji Yoshida 2 and Tadashi Komai 2 'Department of Mechanical Engineering, Saitama University, 255 Shimo-Okubo, Urawa, Saitama 338-8570, Japan 2Nissan Motor Co., LTD

ABSTRACT

Aluminum alloy castings for automobiles are partially strengthened by remelting their surfaces to a depth of several millimeters with a TIG welding machine. In the present work, a nondestructive method was proposed to evaluate the thickness and the width of the remelted zone by applying a phenomenon that the intensity of backscattering ultrasonic wave depends on the size and distribution of crystal grains in materials. By applying this method, the thickness and width of the remelted zone of cast alloy AI-Cu-Si plates were evaluated nondestructively. Estimated thickness and width of the remelted zone were in good agreement with the results obtained by a destructive metallographic test: measurement of the remelted zone size on the cross-section of the specimen. KEYWORDS Nondestructive testing, ultrasonic testing, aluminum alloy castings, backscattering wave, remelted zone, thickness, grain size

INTRODUCTION Recently, some automobile engine parts made of A1 alloy castings are strengthened by a remelting treatment, and evaluation of a remelted zone size has been very important for insurance of reliability of these parts. Currently, only destructive metallographic examination provides a method to determine the remelted zone size. However, this method results in salable castings being sacrificed, offers no guarantee that all the other castings suppliedare of the required quality, and is time consuming. Hence, it is necessary to exploit a nondestructive method to evaluate the remelted zone size. The ultrasonic measurement is particularly useful and has been applied extensively in the evaluation of material structures and material properties, such as elastic modulus and density [1,2], grain size [3], porosity and texture [4], and plastic deformation [5]. Acoustic velocity, ultrasonic attenuation and waveform analysis were used as effective tools for the above purposes. On the other hand, the ultrasonic backscattering wave was also employed for evaluation of the microstructure and depth of the treatment layer [6,7]. In these studies, difference of ultrasonic properties between matrix and treatment layer caused by different phases and structures is employed. By using these characteristics, the depth of the treatment

186

G. Jiang et al.

layer was evaluated. However, it has never been clarified if ultrasonic characteristics are applicable for evaluation of the thickness of the remelted layer on materials in which the treatment layer has the same phases and structures as the matrix but in different grain sizes. The objective of this study is to investigate the feasibility of the ultrasonic measurement for evaluation of the thickness of a remelted layer in A1 alloy castings.

THEORETICAL BACKGROUND When the ultrasonic waves travel through a medium, the intensity P decays exponentially as: P(x) = P0 exp(- c~x)

(1)

where P0 is the intensity of the incident wave, x is a travelling distance of the wave, and a is the attenuation coefficient. The intensity of a backscattering wave (IB) is proportional to the rate of change in the wave intensity, and so to the wave intensity and the attenuation coefficient, such as: IB oc dP(x)/dx = A c~P(x)

(2)

On the other hand, ultrasonic attenuation is caused by various microstructural parameters but is, for most metals, primarily due to grain scattering. The relationship between attenuation and grain size depends on the ratio of the acoustic wavelength, ~, , to the average of some measure of grain size, D. Generally, three regimes are considered: Rayleigh regime (Z >>D) c~=KrD3t ~ (3) Stochastic regime (;t - D ) ~ =KsDfz (4) Diffusion regime (Z

"go 0

~:~:~-
]

~

sphere

E=15GPai v=0"24 i

~

//

U

::::::::::::::::::::::::::::::::::::::::::::::

~.........................................:. r.:.:.:.:.:.:.:.:.:.:.:.:.:.:.:.:.:. :.:, --3 ":':':':':':':':':" iiiiiiiiiiiii!i!!!!!iii!iiiiiiiiiiiii .d-.3..iiiiiifiiiii i :.:.:..:.:.1.- ::::::::::::::::::::::::::::::::::::::::::::::::::::::::::::::::: t....-.-.-.-.................. :-.'-.~:-'.'-.'-.~:~.:.~.~ ==================================

~~ . : . ' . ' . -~- . : 2 : 2 : . : . : .

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t.,,,-.....I.. . . . .

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::::::::::::::::::::::::::::::::::::::::::::::::::::: v =0.24~:i:i:i:i:i:!:i:!:i:i:i:i:!:i:i:i:i:i:i:i:i ~:.:.:.:.:.:.:.:.:.:.:.:.:.:.:.:.:.:.:.:.:.:.:.:.:.:.:. ~......................................................:.;.:.;.:.:.:.:.; i

======================================================================== i

(a)

(b)

Fig.7 Models for the finite element analysis. The unit is nm.(a) Low elasticity layer under the surface (b) Low elasticity layer with the radius of 3nm under the surface. 114

T.N

113

I

I

I

I

I

l

j

I

I

I

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I

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I

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I

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1.

-:

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1070

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,--112

8 110 r 109

I

homogeniously

i

radius of 3nm under the surface (b) low elastic layer i U2der the surface i

with the

.... ...

cantilever stiffness for caliculating

the resonancefrequency 50

100

i

-

150

Load F,nN Fig.8 The first resonance frequency calculated by the finite element method.

7'. Tsuji and K. Yamanaka

220 CONCLUSIONS

We applied UAFM to cleaved surface of HOPG, analyzed dislocations and obtained conclusions as described below. We were able to image subsurface dislocations by the amplitude distribution of the cantilever deflection vibration at the first resonance frequency measured at the position with no dislocation.

The resonance frequency measured at the position with no dislocation

monotonically increased as the contact load increased. theoretically calculated resonance frequency.

It was in good agreement with the

But the resonance frequency measured on the

dislocation showed a unique behavior and had a local minimum at a specific load and were able to be partly explained by the finite element analysis.

In future, we may obtain the

information of the dynamics and the depth of the dislocation by the detailed analysis of the load dependence of the resonance frequency measured on the dislocation.

We therefore

summarize that UAFM is particularly useful for analysis of subsurface defects on the nanoscale. AKNOWLEDGEMENT This work was supported by the Grant-in Aid for Science Research (No. 1045015), and by the Grant-in Aid for COE Research (No.lICE2003), The Ministry of Education, Science, Sports and Culture. REFERENCES 1.

Yamanaka, K., (1996) Thin Solid Films 273,116

2.

Yamanaka, K. and Nakano, S., (1996) Jpn. J. Appl. Phys. 35, 3737

3.

Yamanaka, K. Noguchi, A., Tsuji, T., Koike, T and Goto, T., (1999) Surf. Interface Anal.

27, 600 4.

Tsuji, T., and Yamanaka, K., submitted to (2000) Nano technology

MODELING

This Page Intentionally Left Blank

Nondestructive Characterizationof Materials X Green et al. (Eds) 9 2001 Elsevier Science Ltd. All rights reserved.

223

M O D E L L I N G O F U L T R A S O N I C ATTENUATION

IN UNIDIRECTIONAL FIBER REINFORCED PLASTICS S. BIWA, Y. WATANABE and N. OHNO

Department of Micro System Engineering, Nagoya University, Nagoya 464-8603, Japan

ABSTRACT A theoretical model of ultrasonic attenuation is formulated for unidirectional fiber reinforced polymer-based composites. The model accounts for energy losses due to wave scattering by the fibers as well as viscoelastic absorption in the matrix. The modelling yields a simple formula for the overall attenuation coefficient of the composite in terms of the properties and volume fractions of the constituents. Numerical analysis is carried out for frequency-dependent attenuation behavior of a longitudinal wave propagating in a unidirectional carbon fiber reinforced epoxy composite. The analysis reveals that in a frequency range of practical NDE applications the viscoelastic absorption in the matrix is the major governing factor of the composite attenuation characteristics. The results of the analysis are also discussed in the light of the corresponding experimental results. KEYWORDS Ultrasonic attenuation, polymer-based composite, scattering, absorption, viscoelasticity

INTRODUCTION Attenuation measurements for fiber reinforced plastics have gained much interest from the viewpoint of nondestructive material characterization. Among others, Lhermitte et al. [1] measured the frequency-dependent attenuation of carbon fiber reinforced plastics (CFRP) for different propagation and polarization directions. In a series of detailed study Hosten and his associates have investigated characterization methods of anisotropic viscoelastic properties of CFRP by ultrasonics, e.g. [2]. Another issue of much concern has been the evaluation of matrix porosity, c.f. [3]. Scattering theories have been employed by many investigators to theoretically examine the wave attenuation in polycrystalline metals and elastic composites, e.g. [4]. When ultrasonic attenuation in polymer-based composites is to be analyzed, however, it is important to properly account for the viscoelastic nature of the matrix in addition to the scattering loss

224

S. Biwa, Y. Watanabe and N. Ohno

Fig.1 Modelling of wave propagation in a unidirectional fiber-reinforced composite.

due to microstructure. Although some of existing scattering theories are capable of treating viscoelastic composites, e.g. [5], apparently seldom of them actually discuss the correlation between the theoretical results and the corresponding measurements of attenuation. In the present study, a theoretical model for ultrasonic attenuation in unidirectional fiber reinforced viscoelastic-matrix composites is formulated based on the evaluation of the energy loss of the propagating wave due to the scattering by fibers as well as the viscous absorption in the matrix. The present model is an extension of the classical independent scattering theory [6] to incorporate the viscoelastic matrix-elastic fiber system. Based on the present modelling, numerical analysis is carried out for attenuation behavior of a longitudinal wave propagating in a unidirectional carbon fiber reinforced plastics. The frequency dependence of attenuation is also discussed in the light of experimental results.

MODELLING OF ULTRASONIC ATTENUATION IN FIBER REINFORCED COMPOSITE Consider a plane ultrasonic wave in a unidirectional composite which propagation direction lies in the cross-sectional plane normal to the fibers as in Fig.1. It is the objective of the present modelling to evaluate the energy loss when the plane wave propagates a unit distance that is assumed small on a macroscopic scale but sufficiently large compared to typical microstructural lengths such as fiber radius. When the composite consists of elastic fibers and viscoelastic matrix, both assumed in-plane isotropic, the ultrasonic attenuation is considered to be brought about due to the wave scattering by the fibers and the energy absorption in the matrix. In this situation, the spatial decay of the time-averaged energy flow density < e > of the plane wave as it propagates a unit distance in the xl-direction can be written as d ( e ) / d x I =--{(ISC")+(lmat)}, (1) where < Isca > and < l mat > are the scattering and the absorption loss rates, respectively, in a region V of unit scale and volume as depicted in Fig.1. Furthermore, it is assumed that the unidirectional fibers are arranged with volume fraction ~ in a random spatial configuration. In this circumstance, it is intractable to calculate < I sca > and < I " a ' > in an exact manner. These losses are instead estimated in the following way.

Ultrasonic Attenuation in Unidirectional FRP

225

To evaluate the scattering loss, a simplified problem is first considered for plane wave scattering by a single elastic fiber of radius a embedded in a viscoelastic infinite matrix. The where k I +ial incident plane wave is expressed by u~i"c - - R e [ ~ e x p { i ( k l x l - c o t ) } ] , denotes the complex wave number in the viscoelastic matrix and co the angular frequency. The frequency-dependent phase velocity and the attenuation coefficient of the matrix are denoted by c~(co) and al(co), respectively. Throughout the discussion, steady-state harmonic wave fields are considered and the time dependence of the form exp(-icot) is assumed. The time-averaged energy flow density of the plane wave decays along the propagation direction, presently chosen as the xs-direction having the unit vector denoted by fl,. As a reference, the energy flow density < e > is evaluated at the origin taken at the center of the fiber. Using the energy flux density vector, or Poynting vector, p/,c__ _cr,/.cti/.c of the incident wave, it is given by

=co/c1

T

(e)o=(e)lx~=O =lfpii"cfli]x~=odt.

(2)

o

Introducing the scattered wave u, sc~ and the refracted wave ul r'f , the wave fields in the case of the single fiber in the infinite matrix are written as u, -- u, i"c + u, ~" outside f2 and u, = u, "f inside f~. Here and hereafter, f2 denotes the region inside the fiber and F its boundary. The scattering energy loss rate is given by the time-averaged energy flow of the scattered wave integrated over F as T

(Isca) = l f f p i S C a n i d S d t = y S C a ( e ) o .

(3)

OF

In the above expression, psi. = _cr,/~a/i/~. is the energy flux density vector of the scattered wave and n, is the unit normal vector to the boundary, taken positive outwards from the fiber. This loss has been normalized by the reference incident energy flow density < e >0 and the scattering cross section ~lsca has been introduced. To proceed, the absorption loss as the wave propagates in the matrix is to be evaluated. In the present modelling, first the energy dissipation of the incident plane wave in the matrix region is considered where the unit-volume element V contains a single fiber. This is given by the time-average of the stress work rate integrated over the matrix region V-g2, as

(Imat>l -Zfo fV-f2 (Yij =

inc.inc

Eij a. . .v. .a [

_~

= 2a 1 (e)o - 2a-rall(2ala)(e)o.

1T

#ninc -Jl~iO v niV dS

+fee r incnidS

dt

(4)

To arrive at the above expression the divergence theorem has been used. The normal vector ni V is taken outwards at the boundary of V, 0V, while ni is taken similarly to that in eq.(3). Further, the integrations over 0V and F have been carried out with the normalization by < e >0. In the last expression in eq.(4), 11(') is the modified Bessel function of the first order. It is also noted that the attenuation of the incident wave for unit propagation distance has been assumed small and evaluated to a first order. The first term of the last expression in eq.(4), 2al < e >0, denotes the energy absorption when the plane wave propagates a unit distance in the viscoelastic matrix in the absence of the fiber. The second term represents the partial reduction of the absorption loss as a portion of the matrix is replaced by the fiber. An effect similar to this has been already incorporated by Brauner and Beltzer [7], who regarded the

226

S. B i w a , Y. W a t a n a b e a n d N. O h n o

energy flow into the fiber across the fiber boundary to attain a negative value and represent a supportive effect on the plane wave. However, when averaged for one period this quantity should vanish if the fiber is purely elastic. At a closer look, an erroneous manipulation is found in their subsequent derivation, and they end up with a similar result to eq.(4). The formulation of the matrix absorption presented here is thus considered to be an improvement of their theory. Based on the above consideration, now the energy loss rate of the wave in the composite containing n s = ~ / ( z a 2) fibers per unit volume is examined. In the present model, no interaction among the neighboring fibers is taken into account. Then the scattering loss due to ns fibers is given by < I sca >= ns < I ~c" > 1 ~- n s y sca < e >0, and the matrix absorption becomes < I mat >= 2~1 < e >0 - 2 a r a m l l ( 2 a l a ) < e >0. Identifying < e >0 with < e > and substituting these results into eq.(1) yield an ordinary differential equation for < e > (Xl), which is readily solved to give (e)(Xa) ~ e x p { - ( 2 a 1 - 2.amn,I 1(2ala) + n s y ,c" )x~}. (5) Since the energy density is a quadratic function of the displacement amplitude, the attenuation coefficient of the composite a is obtained as 1 a = a~ - zran~I~ (2ala) + ~ n , 7 "ca . (6) In ordinary situations, the argument 2ala of the modified Bessel function takes on a small value. Then Ix(2ala) can be well approximated by the leading term of its expansion. Therefore, a simple formula is obtained for the composite attenuation coefficient as 1 a = a 1 - ~ a 1 + - ~ n ~ ''ca . (7) The above expression implies that the presence of the fibers of volume fraction 4~ reduces the absorption effect of the matrix to the ratio (1-4~). The idea of viewing the matrix absorption to decrease in proportion to the volume fraction of the dispersed phase has been noted by Kinra et al. [8] in their study of ultrasonic attenuation in a particulate composite, without any theoretical reasoning though. This reduction effect has been naturally deduced on a theoretical ground in the present model. In the above discussion, the so-called independent scattering model has been employed implying that the scattering loss is determined based on the single fiber scattering in an additive manner. This is valid when the fiber fraction is not very large so that fibers are located sufficiently remote from each other, or when the magnitude of the scattered wave is sufficiently small compared to that of the incident wave. It is remarked here that the independent scattering assumption as in eq.(7) tends to overestimate the scattering loss.

Fig.2 Longitudinal wave propagating in a unidirectional CFRE

227

Ultrasonic Attenuation in Unidirectional FRP

ANALYSIS OF LONGITUDINAL WAVE ATTENUATION IN CFRP The formula derived above is subsequently applied to the attenuation of longitudinal ultrasonic wave in a unidirectional carbon fiber reinforced epoxy composite (hereafter, CFRP) propagating in the direction normal to the fibers, as shown in Fig.2. The analysis requires data of viscoelastic properties of the epoxy matrix as well as elastic properties of the carbon fiber. To this purpose, the phase velocities and attenuation coefficients of the epoxy resin, similar to that used in CFRP, are measured for longitudinal as well as transverse waves. From the velocities and attenuation coefficients, the complex moduli of the epoxy resin can be calculated. As a result of the measurements, it was found that within the range of frequency of the present interest, the phase velocities were well approximated as frequency-independent constants and the attenuation coefficients as first-order equations of the frequency, as shown in Fig.3. It is widely known that polymeric solids exhibit linearly frequency-dependent attenuation. Although the measured attenuation data are limited to a relatively low frequency range, in the following analysis these fitting results are extrapolated to higher frequencies. Carbon fibers (with diameter 7~tm) are assumed to exhibit transversely isotropic elasticity. Thus two in-plane elastic constants are chosen from the literature, e.g. [9]. The material properties of the epoxy resin and the carbon fiber used in the numerical analysis are summarized in Table 1. The single fiber scattering analysis was carried out by a classical technique introducing two-dimensional complex displacement potentials for the scattered and the refracted waves. These potentials were expanded into eigenfunctions and the expansion coefficients were determined from the continuity conditions at the fiber-matrix interface [10]. Once the scattered wave field is determined, the scattering cross section ),,ca is obtained from eq.(3), which yields the attenuation coefficient of the longitudinal wave in CFRP in eq.(7) by identifying the matrix attenuation as 5 1 ~ ~ L I " Table 1. Material parameters of matrix and fiber. Epoxy matrix Longitudinal velocity c~1 [m/s] Longitudinal attenuation coefficient

7

y6 ~5

2668

aLl(Og) ----aLl 0 + a L l 1 O)/(2.717) ;

all o [1/cm] all I [1/(cm"MHz)] Transverse velocity Crl [m/s] Transverse attenuation coefficient

"~ 4 o

~3

--0.607 0.495 1187

a T l ( t O ) ----aT10 + aT11 CO1 ( 2 ~ ) ; 0

1

~ ' - ""

2

~

3

~

~

'

I

4 5 6 7 Frequency [MHz]

I

8

Fig.3 Attenuation coefficients of epoxy resin (experimental results).

ano [1/cm] at,, [1/(cm"MHz)] Density Pl [g/cm3] Carbon fiber Lam6constant Z2 [GPa] Lam6constant /2 2 [GPa] Density P2 [g/cm3]

--0.554 2.34 1.23

9.97 5.02 1.67

228

S. Biwa, Y. Watanabe and N. Ohno

RESULTS AND DISCUSSION The attenuation coefficient of the CFRP has been computed as a function of the frequency in Fig.4 in the case of ~ = 0.2, for which the independent scattering may apply. The frequency f -o9/(2n') has been normalized by the matrix longitudinal wave speed cL1 and the fiber radius a, and the attenuation coefficient a by a. The illustrated range of the normalized frequency 0 < ( a / % ) f < 0.05 corresponds roughly to 0[MHz] < f < 38[MHz] in the present material system. In Fig.4, the attenuation coefficient of the matrix, obtained by linear fitting to the measurement, is delineated as a broken line. To indicate the relative contribution of the two loss mechanisms, i.e., scattering and absorption, the reduced matrix absorption (1-~)a~a is also depicted in Fig.4 as a chained line. One finds from Fig.4 that the scattering effect is relatively small in the low frequency range, while it grows substantially as the frequency increases. In the region where the normalized frequency (a/cL~)f is less than about 0.034, reinforcing the matrix with fibers results in reduction of the ultrasonic attenuation, i.e., a < al. This can be explained by reasoning that compared to the increase in the scattering effect by increasing ~, the decrease in the viscoelastic absorption due to the reduction of matrix volume fraction ( 1 - ~ ) makes a greater negative contribution to the composite attenuation. On the contrary, as the normalized frequency increases to a sufficient extent, the scattering tends to cause a relatively greater effect, and the composite attenuation may eventually exceed that of the matrix. The normalized frequency used in Fig.4 is equal to the ratio between the fiber radius a and the wavelength of the longitudinal wave in the matrix, A L l - ' C L 1 / f . For the composite system analyzed herein, the fiber radius is 3.5 ~tm and the so-normalized frequency (a/cL1)f is of the order of 10 -2 or smaller for the frequency range of practical interest. From this reason, the attenuation coefficient of the composite is expected to fall below that of the matrix, and this prediction can be supported by actual measurements as discussed below. However, if reinforcing phases were much greater, according to the theory the scattering contribution would become much more significant and the composite attenuation would exceed the matrix attenuation. Kinra et al. [8] in the aforementioned investigation measured the ultrasonic attenuation in a glass particle reinforced epoxy-matrix composite, with particle radius of approximately 150 lam, and reported composite attenuation coefficients greater than those of

0.008 o

' CFRP'(q~=0'.2,compuled) ' -'-'~~P~ 0.006

'

' /~ J']~~t

.,.

0.004 E Z

Scatter

0.002

Absorption

-

00

0.01

0.02

0.03

0.04

0.05

Normalizedfrequency(a/cl.a)f Fig.4 Variation of normalized attenuation coefficient of CFRP with normalized frequency.

Ultrasonic Attenuation in Unidirectional FRP u

,

|

,

,

,

,

A Epoxy(measured) - - - Epoxy (fitted) O CFRP (measured) CFRP (computed) -'~(1-~)al

--'3

|

229

y

1"


6 " ~ s.s ~ s

"'t

As-welded

..

~

~ 4.s 4

3.5 200

,

,

~rl

oo

i

400

,

.

,

i

600

,

,

i

I

m

800

|

,

!

.

1000

PWHT temperature / K

The Lissajous's curves in Figs. 5(a), (b) were obtained from the as-welded specimen and the PWHT specimen at 933K, respectively.

Fig. 4. Detected peak voltage vs. PWHT temperature of 2.25Cr- 1.0Mo weld metal.

344

M. Shiwa et al.

0

lO

................................................................................................

2. ).5Cr-l.0Mo /As-welded i

*.'

2.25Cr-1 0Mo / 933K

5

j ;}

5

0

,g

0

-5

i ii

-5

-10 ................................................................................................

-I0

(a) -12

-6

0

6

12

-12

-6

0

Exciting / V

6

12

Exciting / V

Fig 5 Lissajous's curves: (a) for the as-welded specimen; (b) for the PWHT at 933K specimen

12

n I . . . . . . . . . . . . . . . . . 2.2~i:r-i'.~iMo/)~s-ws

4 """

,::,,.,...,,.,. '"'"...."'""'"

,=

-4

:i::/

"";,

'

/.:: -

-12 0

10

5

15

20

>.

"

4

-4

Detected / V

-

25

30

35

-

40

-12 0

5

10

Time/gs

15

-

20 25 Time/gs

Detected / V 30

35

40

Fig. 6. The exciting and detected waveforms of Figs. 5 (a) for the as-welded specimen; (b) for the PWHT at 933K specimen.

10

/

-20

1

0

. . . .

/

i

. . . .

i

. . . .

i

. . . .

i

0

. . . .

2.25Cr-l.0Mo/As-welded 3rd harmonic component

"~ -30

, ....

, ....

, ....

2.25Cr-l.0Mo/933K L~ h a r m o n i c component

-30 =-40

"~ -40 t-S0

-50

C/I

-60

-60

-70

-70

-80

, ....

-20

/

m

....

-10

0

100

...... 200

300

Frequency / kHz

400

500

-80

0

100

200

300

400

500

Frequency / kHz

Fig. 7. The frequency spectrum of the detected waveforms shown in Figs. 6: (a) for the as-welded specimen; (b) for the PWHT at 933K specimen.

Evaluation of PWHT Using AC Magnetic Method The material is 2.25Cr-l.0Mo weld metal. These figures showed that the profile of Lissajous's curve turned to be constricted in the middle in the case of applying PWHT at 933K. Figure 6 shows the exciting and detected waveform of Figs. 5. It can be seen that the profile of the detected waveform corresponding to the PWHT specimen at 933K has higher harmonic components wave compared with the as-welded specimen. Figure 7 shows the frequency spectrum of the detected waveforms shown in Figs. 6. The intensity of the 3rd harmonics component of the detected waveform from the PWHT specimen was about 20dB higher than that of the as-welded specimen. It is concluded that the intensity of the 3rd harmonics component of the detected waveform depends on the PWHT temperature.

|

10

In order to investigate how the harmonic component intensity of the detected waveform depends on the exciting voltage, several kind of various exciting voltages were applied for the 2.25Cr-l.0Mo steel specimen and the detected voltages were compared with each other. Figure 8 shows Lissajous's curves of various exciting voltages. It can be seen that the Lissajous's curves strongly depend on the exciting voltages. As the hysteresis curve obtained by exciting voltage of 20V was bigger than that of the exciting voltage of 5V, the exciting voltage of 20V was used in this investigation. This result suggests that profile of Lissajous's curve as just over the Rayleigh loop region of small hysteresis, which was observed under the exciting voltage of 10V, in the ferromagnetic material can be useful to the weld metal.

9 .

|

9 9 .

o

|

....

,

9 .

|

.

,

.

i

2.25Cr-l.0Mti X XXx x X X .. X~ <X>~ X xO" ~ X

5

0

-5

Xx

x x x x

.

-10

.

i

.

,

.

-8

-12

|

.

o o

Exciting Exciting

voltage voltage

of 5V of 10V

o

Exciting

voltage

of 15V

x

Exciting

voltage

,

.

i

-4

.

.

,

|

0

.

,

of 20V

.

I -...

4

8

12

Exciting voltage / V

Fig. 8. Lissajous's curves of various exciting voltages

=~

P WHT temperature evaluation by the normalized 3rd harmonic component intensity

.

345

-5

~ -lo

i

,w

.-~ .,~ ~ -15 ,~ 2-20 m "~ "~ - 2 5 2 ;~ -30

J 99 ."

~X

0

/ f /

---o- 1.25Cr-0.5Mo/

, 9 f %, a s

~

CP

- 125Cr-0.5Mo/MP

- 0-- 2.25Cr-l.0Mo/CP

/

300

Q

....'"'J

S:~ / /

i

~

- -x-- 2.25Cr-1.0/MP ,-, ~-,

i

400

,

l

500

~

i

i

600

PWHT

t

|

700

temperature

i

800

i

l

900

,

i

1000

/ K

Fig. 9. Relationship between the normalized 3rd harmonic component intensity (Fh3) and the PWHT temperature, corresponding to the 1.25Cr-0.5Mo steel and 2.25Cr-1.0Mo steel specimens.

The curves in Fig. 9 show the relationship between the normalized 3rd harmonic component intensity (Fh3) and the PWHT temperature, corresponding to the 1.25Cr-0.5Mo steel and 2.25Cr1.0Mo steel specimens. Since a value of the Fhs was more stable compared with the detected peak voltage, the normalized 3rd harmonic component intensity of the detected signals, which is defied as the ratio of the 3rd harmonic component intensity to intensity of the harmonic component of fundamental exciting frequency, was used in this investigation. It is shown that the Fh3 increases with the PWHT temperature in the all cases.

M. Shiwa et al.

346 DISCUSSIONS

Accuracy of correlation curves To evaluate the PWHT temperature accurately, it is important to correlate the PWHT temperature with measurement data accurately. Here, the correlation curves are made as a set of regression curves based on the measurement results as the master evaluation curves. The regression curves of MP, F h ~ e, and CP, Fh3cP, of 1.25Cr-0.5Mo steel specimens shown in Fig. 9 can be estimated as following equations,

F h ~ P = -30.14+0.043T+2.350 * 104 T 2

(1)

Fh3c1~= -18.24+0.015T+2.367 "104 T 2

(2)

where T is PWHT temperature. The coefficient of correlation of the MP (RMP) in Eq. (1) is 0.98 and the CP (R cP) in Eq. (2) is 0.98. The regression curve of MP, F h f ~, and CP, Fh~c1~, of 2.25Cr-l.0Mo specimens shown in Fig. 9 can be estimated as following equations,

F h ~ P = -44.81+0.067T-3.510"104 T 2

(3)

F h f P= -33.51+ 0.017T-6.234" 10"6 T 2

(4)

The coefficient of correlation of the MP (_RuP) in Eq. (3) is 0.99 and CP (R cP) in Eq. (4) is 0.99. These high values of coefficient of correlation mean that regression curves model the measurement data very well and can be used as the master curves for evaluating the PWHT temperature.

Effect of the surface treatments Fig. 10 is the enlargement of a part of Fig. 9 where the PWHT temperature ranges between 843K and 1013K. It can be seen that the Fh3 corresponding to the CP increases with the temperature monotonously, but the Fh3 corresponding to the MP does not. The reason is thought to be as follows. Penetration depth of magnetic flux, 6, was evaluated as the following equation, "~ ---o- 1 . 2 5 C r - 0 . 5 M o / C P - *- - 2 . 2 5 C r - l . 0 M o / C R -8

. . . .

-9

8-- 1/(~f ~LO)1/2

(S)

. . . .

i

. . . .

i

. . . .

i

. . . .

- 1.25Cr-0.5Mo/MP-x--2.25Cr-l.0/MP

t~ .~

where f is the exciting frequency, ~t is magnetic permeability and o is electric conductivity. When the exciting frequency is 60kHz, the magnetic permeability is 750 and the electric conductivity is 60 x 10s D, the penetration depth is about 100 ~tm. It is known that hardening depth caused by mechanical polishing is about 50 to 100 gm. Signals obtained by the 60kHz exciting frequency may be influenced form the hardening layer. Therefore, when the high frequency exciting probe is used to evaluate PWHT temperature, the CP is preferable for the measurement.

~

i

i

-I0

/

-ll

'~ -12

~

~f;~

.... x,,.,,,,\'-'

: ....

-

x''/'

-14

"'x

@

Z

-15

800

. . . . . . . . . . . . . . . . . . .

850

900

950

'

....

1000

1050

P W H T temperature / K

Fig. 10. The enlargement of a part of Fig. 9 where the PWHT temperature ranges between 843K and 1013K.

Evaluation of PWHT Using AC Magnetic Method

347

CONCLUSIONS The AC magnetic testing method of high exciting frequency (60kHz) by using a small coaxial probe is proposed to evaluate the PWHT temperature, nondestructively. Low alloy steel welded joints specimens (1.25Cr-0.5Mo and 2.25Cr- 1.0Mo) under several PWHT temperature conditions were investigated. Also influence of the surface condition on the testing result was investigated. Two types of surface treatment, a chemical polishing (CP) and a mechanical polishing (MP), were used in the test. 1) The normalized 3rd harmonic component intensity (Fh3) of the detected signals, which is defined as the ratio of the 3rd harmonic component intensity to the fundamental harmonic component intensity, changes considerably with the PWHT temperature. 2) The regression curves based on the measurement value of Fh3 were formulated to evaluate the PWHT temperature as the master curves. The evaluated PWHT temperature based on the master curves agrees with the actual PWHT temperature very well. 3) Measurement after the CP surface processing gives more reliable results.

REFERENCES

Yamaguchi, A. and Shiwa, M. (1996) 14th International Conference on NDE in the Nuclear and Pressure Vessel Industries, ASM, Stockholm, 177. Kwun, H. and Bukhardt, G. (1987) J. of Phys. 61, 1576-1579. Jiles, D., Thoelke, J., Clark, W., Iyer, J. and DeNal, R. (1991) Review in Progress in QNDE 10B, pp2015-2020, Plenum Press, New York. Chen, Z., Govindaraju, M., Jiles, D. and Biner, S. (1994) 1EEE Transaction on Magnetic 30, 4596-4598.

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Nondestructive Characterizationof Materials X Green et al. (Eds) 9 2001 ElsevierScience Ltd. All rights reserved.

349

SQUID NONDESTRUCTIVE DAMAGE ANALYSIS F O R A U S T E N I T I C STAINLESS STEEL T. SUZUKI 1, K. HIRANO 1, and K.W.LEE 2 1. Mechanical Engineering Laboratory, AIST, MITI Namiki 1-2, Tsukuba-shi, Ibaraki-ken, 305-8564, Japan 2. Korean Advanced Institute of Science and Technology

ABSTRACT The superconducting quantum interference device(SQUID) is an ultrahigh-sensitive magnetic sensor and began to be applied to nondestructive damage analysis of structural materials. In this study fatigue damage distribution for austenitic stainless steel SUS316L and damage formation under cyclic and static loading for austenitic stainless steel SUS304 was studied using SQUID. To measure fatigue damage distribution, SQUID output data were analyzed using the reference data subtraction (RDS) method, and it was found that fatigue damage was correctly detected by this method. To measure damage formation under cyclic and static loading, in-situ SQUID damage monitoring tests were conducted, and it was found that the start of damage formation was correctly detected by the change in SQUID output.

KEYWORDS Nondestructive damage analysis, SQUID, austenitic stainless steel, fatigue crack, ct'- martensite, in-situ damage monitoring.

INTRODUCTION The superconducting quantum interference device(SQUID) is an ultrahigh-sensitive magnetic sensor and began to be applied to nondestructive damage analysis[I-4]. We[2,3] conducted nondestructive damage evaluation tests and proposed SQUID output data analysis methods of the adjacent data difference(ADD) and the reference data subtraction(RDS). And it was found that at a cryogenic temperature, the location of fatigue crack tip and the amount of fatigue damage for austenitic stainless steel SUS316L were correctly determined by detecting ot'-martensite using these methods. In this paper, two kinds of SQUID damage evaluation tests were conducted at room temperature. Fatigue damage distribution for austenitic stainless steel SUS316L was studied by a nondestructive SQUID damage analysis system. Damage formation under cyclic and static loading for austenitic stainless steel SUS304 was studied by an in-situ SQUID damage monitoring system. SPECIMENS AND EXPERIMENTAL PROCEDURE Materials and ,specimens Materials used in this study were austenitic stainless steel SUS316L (C0.015, Si0.59, Mn0.98,

350

T. Suzuki, K. Hirano and K. W. Lee

P0.026, S0.001, Ni12.12, Cr17.43, Mo2.10%) and SUS304 (C0.04, Si0.46, Mn0.86, P0.028, S0.003, Ni8.17, Crl 8.17%). The shape and dimensions of the specimen is shown in Fig. 1. Experimental procedure Fatigue tests were conducted by the nonmagnetic electrohydraulic materials testing system at a stress ratio of 0.05 in a constant load amplitude and at a frequency of 10 Hz. Fatigue crack length was monitored by a traveling microscope and a scanning laser microscope. The SQUID output distribution around the fatigue crack tip was measured by a SQUID nondestructive damage analysis system. The schematic representation of the system is shown in Fig. 2. The system consisted of a magnetically shielded room, a SQUID sensor, a SQUID controller, a cryostat, a X-Y stage, a gantry, and a computer. The magnetic shielding ratio of the static magnetic field in the magnetically shielded room exceeded 1/100. A magnetometer DC" SQUID sensor was used to improve spatial resolution. The gantry was made of FRP and aluminum alloy. SQUID output measurement, X-Y stage control, and data analysis were conducted by the computer located outside the magnetically shielded room. The change in SQUID output under cyclic and static loading for austenitic stainless steel SUS304 was measured by an in-situ SQUID damage monitoring system. The schematic representation of the system is shown in Fig. 3. The system consisted of a SQUID sensor, a SQUID controller, a cryostat, a computer, and an electrohydraulic nonmagnetic materials testing machine. The cryostat was placed above the specimen so that the specimen and the bottom of the cryostat were 1 mm apart. The DC. SQUID sensor was used with a 1-dimensional differential pickup coil to remove noise around the system. In the electrohydraulic nonmagnetic materials testing system, nonmagnetic structural materials were used within 1 m of the SQUID sensor. EXPERIMENTAL RESULTS AND DISCUSSION SQUID output distribution around fatigue crack tip In the previous papers[2,3], we proposed the reference data subtraction(RDS) method. This method used initial SQUID output Bo as reference data to calculate the difference between SQUID output after mechanical test B and reference data Bo. SQUID output distribution around the fatigue crack tip analyzed by the RDS method is shown in Fig. 4. The reference data was SQUID output distribution before the fatigue test. By the RDS method symmetrical SQUID output was obtained. The scanning laser micrograph around the fatigue crack tip for austenitic stainless steel SUS316L is also shown in Fig. 4. In the scanning laser micrograph, plastic deformation region was shown as the darker region. It was found that SQUID output distribution around fatigue crack tip corresponded to the scanning laser micrograph. In our previous researches[5,6] with SUS316L using a magnetic force microscope or a X-ray diffraction device, it was found that SUS316L had metastable austenite. At cryogenic temperatures, &-martensite, which was a high-magnetic phase, was easily formed around the fatigue crack tip. However, at room temperature, almost no or only slight amounts of ot'-martensite was formed. Then, the change in SQUID output around the fatigue crack tip at room temperature was induced not only by (~'-martensite around the fatigue crack tip but by the difference in magnetic characteristics between plastic deformation region and the original region.

SQUID Nondestructive Damage Analysis

351

Changes m SO_.UID output under cyclic and static loading Calibration of nonmagnetic materials" testing ,system. Because of the ultrahigh sensitivity of the SQUID sensor, it was possible that nonmagnetic structural materials in the testing system affected SQUID output measurement. Then, SQUID output was measured without specimen by moving the actuator for 2 mm. The calibration results are shown in Fig. 5. It was found that SQUID output was changed only by moving the actuator, then all output data were compensated for by the results in Fig. 5.

Changes in SQUID output during cyclic loading. The relationship between SQUID output and the displacement, and that between the load and the displacement at the crack length of 5.9ram are shown in Fig. 6. In each stage of fatigue crack growth, the load was related linearly to the displacement, as was SQUID output. SQUID output differences between the maximum and minimum load under cyclic loading are shown in Fig. 7. The differences in SQUID output remained almost constant during fatigue crack growth. Then it was found that fatigue damage formed in each cycle could not be measured by in-situ SQUID damage monitoring. Changes in SQUID output under static loading. The specimen was statically loaded up to 9.8 kN over the maximum value of fatigue test at the crack length of 6.0 mm. The relationship between SQUID output and the displacement, and that between the load and the displacement are shown in Fig. 8. The load was initially related linearly to the displacement. However, as the load increased over the maximum value of the fatigue test, the relationship between the load and the displacement became nonlinear. It was also found that SQUID output was initially related linearly to the displacement, then the relationship between SQUID output and displacement became nonlinear. The load at which the load-displacement relationship became nonlinear coincided with the starting load of the nonlinearity of the SQUID output-displacement relationship. This was because large plastic damage formation began at this load. Then it was found that by in-situ SQUID damage monitoring, the start of a large plastic damage formation could be detected. CONCLUSIONS (1)

SQUID output distribution around the fatigue crack tip for austenitic stainless steel SUS316L was studied at room temperature. By the RDS method the change in magnetic field induced by plastic deformation and ot'-martensitic formation around the fatigue crack tip was correctly detected.

(2)

SQUID output change during cyclic and static loading was studied by an in-situ SQUID damage monitoring system. In static loading SQUID output was changed consistently with large plastic damage formation. It was concluded that the start of the large plastic damage formation could be detected by in-situ SQUID damage monitoring.

REFERENCES Suzuki, T. and Hirano, K. (1994) Proc. 72ndJSME Fall Annu. Meet., 581 Suzuki, T. and Hirano, K. (1996) Proc. 73rd JSME Spring Annu. Meet., Vol. 1I, 272. Suzuki, T. and Hirano, K. (1999) Progress' m Experimental and Computational Mechanics in Engineering and Material Behavior, Northwestern Polytechnical University Press, 432. Kasai, N. Applied Physics (1998) Vol. 67-4, 417. Suzuki, T. and Hirano, K.(1998) Proc. 12th Bienni. C'onf on Fracture, Vol. 1, 97. Suzuki, T. and Hirano, K. (1999) Proc. 7th Int. Fatigue Cong., Vol. 1, 463

352

T. Suzuki, K. Hirano and K. W. Lee Position of SQUID Sensor

//

01

t=2

'9

~,,.~ . . . .

;4 2;.0.3 I

c) or)

140 Unit :mm

Fig. 1.

Fig. 2.

Shape and dimensions of specimen.

Schematic representation of SQUID non-destructive damage analysis system.

SQUID Nondestructive Damage Analysis

HydraulicPomp

i;~~/~ I

Fig. 3.

353

LoadCell HydraulicGrip SQUID&~Cryostat//// SQUIDController

I[~ ~

Specimen

Schematic representation of in-situ SQUID damage monitoring system.

Fig. 4. SQUID output distribution by RDS method and scanning laser micrograph around fatigue crack tip for SUS316L: (a)SQUID output distribution by RDS method ; (b)scanning laser micrograph.

354

T. Suzuki, K. Hirano and K. W. Lee x 10-3 15

4.a

+a

o

10

s O CO 4.o ro

.m

}.

j

E3

J

~

/"

0

I

0.5

0.0

I

I

1.0

1.5

Half Crack Length

2.0

x 10 -3

m

Fig. 5. SQUID output calibration for nonmagnetic materials testing system. 10.0

I

--

8.0

20.0 x 10-3 oad i SQUID Output 15.0 >

Z _~ 6.0

Q,. "-,z

10.0 O C3

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0.1

I

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0.2

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Displacement (P . . . .

0.0

OJ x l O -3

m

7.84 k N, a=5.9mm)

Fig. 6. Relationship between SQUID output and displacement, and relationship between load and displacement under cyclic loading for SUS304.

SQUID Nondestructive Damage Analysis

355

x 10 -3 15 > ,,l.a

10

0

O0

0

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0

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0

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I

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4.0

4.5

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Fig. 7 Differences in SQUID output between maximum and minimum load during fatigue crack growth for SUS304.

'--'L;au"DOu'0u

10.0

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Z v'

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Fig.8 Relationship between SQUID output and displacement, and relationship between load and displacement under static loading for SUS304

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Nondestructive Characterizationof Materials X Green et al. (Eds) 9 2001 Elsevier Science Ltd. All rights reserved.

357

EFFECT OF PLASTIC DEFORMATION ON STRESS MEASUREMENT USING MAGNETOSTRICTION Tomohiro YAMASAKI

Faculty of Engineering, Osaka City University, Sumiyoshi, Osaka 558-8585, Japan Masahiko HIRAO

Graduate School of Engineering Science, Osaka University, Toyonaka, Osaka 560-8531, Japan

ABSTRACT Effect of plastic deformation on magnetostriction of steel was investigated. We have already proved that the maximum value of the magnetostriction curve can be used in the nondestructive residual stress measurement in elastically deformed steel. However, it is known that the plastic deformation changes the elastic and magnetic properties, which results in the significant error in the nondestructive stress evaluation by ultrasonics and magnetics. In this study, to reveal the applicability of the maximum magnetostriction method to plastically deformed samples, we measured the magnetostriction curves of plastically elongated specimens under various levels of applied stress. It was predicted that the residual compressive stress was introduced in the deformation direction at the surface of the specimen. The evaluated stresses were compared with those by some other methods, showing that the magnetic anisotropy supports the results by the maximum magnetostriction method. KEYWORDS Stress measurement, magnetostriction, steel, plastic deformation, residual stress

INTRODUCTION Magnetic properties of ferromagnetic steel depend on magnetic domain structure, which is varied by either magnetic field or stress. Nondestructive stress measurement is then possible by detecting the stress induced change in the magnetic properties [1]. We have proposed a new method using the maximum value of the magnetostriction [2]. The magnetostriction of steel shows the maximum value during magnetization process, which can be easily measured with sufficient accuracy, while the measurement of other magnetic properties are greatly affected by the measuring conditions. Comparing the maximum magnetostriction with the master curve, obtained by the loading test prior to the stress evaluation, the stress can be predicted. However, the magnetic properties may also be influenced by the plastic deformation, which causes error in the stress measurement. In this study, we investigated the effect of the plastic elongation on master curves of maximum magnetostriction method.

T. Yamasaki and M. Hirao

358

STRESS DEPENDENCE OF MAGNETOSTRICTION Individual grain of polycrystalline steel is divided into many magnetic domains, each of which is magnetically saturated. In the demagnetized state, the domain is magnetized parallel to one of the crystallographic axes and six types of the domains are distributed randomly in each grain. When an external magnetic field is applied, domain walls move and the volume of the domains magnetized parallel to the field increases at the sacrifice of the other domains. As a result of the domain realignment, the magnetization proceeds and the positive strain, called the magnetostriction, appears in the magnetization direction, because the domain is slightly elongated in its magnetization direction. At the same time, to keep the total volume unchanged, the negative strain appears in the plane normal to the field. When the domain realignment is almost completed, rotation of domain magnetization starts to occur and the magnetostriction starts to decrease as shown in Fig. 1. The stress also moves the domain walls, so that the domains either parallel to the tensile stress or perpendicular to the compressive stress expand. While this domain restructuring results in no magnetization, the magnetostriction appears in addition to the elastic strain. Since the maximum magnetostriction is decided by the volume of the domains normal to the field, the magnetostriction thus shows the stress dependence. Then the maximum magnetostriction is larger for the magnetization either normal to the tensile stress or parallel to the compressive stress as shown in Fig. 2. By comparing the maximum magnetostriction with master curves in Fig. 2, the direction and magnitude of the stress can be evaluated nondestructively.

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Fig. 1. Magnetostriction curves for magnetization parallel to tension.

0 -11i0-100-50 i

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i

I

i

I

0

i

I

oi

I

i

i

50 100 150200

Applied stress , MPa Fig. 2. Master curves for stress evaluation.

EFFECT OF PLASTIC DEFORMATION ON MASTER CURVES Experimental Procedure We prepared tensile specimens from JIS-SM490A low carbon steel plate. The dimensions are 350mm x 39mm x 11mm. Chemical compositions and mechanical properties are listed in Table 1. The specimens being loaded parallel to the rolling direction are called specimens R, and those being normal are called specimens T. We also prepared the full annealed specimens from the same

Magnetostriction of Plastically Deformed Steel

359

plate, called specimens HR. To investigate the effect of the plastic deformation, we measured the master curves for elongated specimens. Prior to the measurement, to avoid the effect of magnetic hysteresis, the specimen was Table 1. Chemical compositions and mechanical properties. Chemical compositions (mass%)

Yield

Tensile

stress

strength

C

Si

Mn

P

S

(MPa)

(MPa)

0.16

0.35

1.43

0.018

0.003

418

548

Fig. 3. Maximum magnetostriction of elongated specimen.

Fig. 4. Shift of master curves.

360

T. Yamasaki and M. Hirao

demagnetized. Then, the magnetostriction was measured in the directions both parallel and normal to the field simultaneously using biaxial semiconductor strain gauges, while increasing the magnetic field stepwise. The measurement was done repeatedly under various levels of uniaxial tensile stress. Results and Discussion Figure 3 represents the master curves of specimen R-2 at several levels of elongation. Because the master curves are greatly shifted by the elongation, the stress measurement seems to be impossible after plastic deformation. However, if the curves are translated to the left, they overlap with each other as shown in Fig. 4. Therefore we supposed that the shift of the master curves is due to the residual compressive stress.

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(a) Maximum magnetostriction of is 9 state. (b) Stress shift. Fig. 6. Magnetostrictive properties after elongation.

Magnetostriction of Plastically Deformed Steel

361

Figure 5 shows the magnetostriction curves of specimen R-2 after 0.3% elongation. Monotonical increase of the transverse magnetostriction for magnetization normal to the elongation predicts the existence of the compressive stress, as explained later. Figure 6 indicates the magnetostriction and the stress at the crossing of the master curves. Since the magnetostriction at the node is almost constant even after plastic deformation, it can be regarded as the isotropic state. Thus the stress shift is considered to indicate the magnitude of the residual compressive stress. The master curves were again measured after stress relief annealing and full annealing. Results are shown in Figs. 7 and 8. After heat treatment, the master curves almost coincide with those of specimen HR-1 before deformation, which also supports our guess.

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