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Nondestructive Characterization of Materials X Proceedings of the 10t" International Symposium on Nondestructive Characterization of Materials TM 26-30 June 2000, Karuizawa, Japan Editors Robert E. Green, Jr.

Center for Nondestructive Eva/uation, The Johns Hopkins University, Baltimore, USA

Teruo Kishi

National Institute for Advanced Interdisciplinary Research, Tsukuba, Ibraki, Japan

Tetsuya Saito

Nationa/ Research Institute for Meta/s, Tsukuba, Ibraki, Japan

Nobuo Takeda

Graduate Schoo/ of Frontier Sciences, The University of Tokyo, Bunkyo-ku, Tokyo, Japan

B. Boro Djordjevic

Center for Nondestructive Eva/uation, The Johns Hopkins University, Ba/timore, USA

2001.

ELSEVIER

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PREFACE The papers published in these proceedings represent the latest developments in nondestructive characterization of materials and were presented at the Tenth Intemational Symposium on Nondestructive Characterization of Materials held on June 26-30, 2000 in Karuizawa, Japan. The symposium was held concurrently with three other symposiums and one workshop in the Tenth Iketani Conference supported by Iketani Science and Technology Foundation, Japan. This symposium is the tenth in the series that began in 1983 in Hershey, Pennsylvania, USA and became an international meeting in 1986 with Montreal meeting. The symposium started with a Plenary Lecture entitled "Application of Non-Contact Ultrasonics to Nondestructive Characterization of Materials" by Professor R. E. Green, Jr., The Johns Hopkins University. Various characterization methods were presented at the symposium, including ultrasonic, X-ray, eddy currents, laser, thermal wave, acoustic emission, optical fibers, optics, magnetics and ultrasonic microscope.

The materials covered metals,

polymers, ceramics, composites and semiconductors. Thin films and coatings as well as smart materials were also emphasized in this symposium. The symposium organizers would like to thank the authors for their excellent presentations and manuscripts and the session chairpersons for keeping the session on time and stimulating lively discussions.

The symposium and these proceedings would have not been possible

without the enthusiasm and extremely hard work of Ms. M. Nakajima and Ms. A. Shirai, Secretaries, at the Research Center for Advanced Science and Technology (RCAST), the University of Tokyo, who answered correspondence and prepared material for all of the notices and publications. Symposium Co-Chairpersons Robert E. Green, Jr. The Johns Hopkins University Teruo Kishi National Institute for Advanced Interdisciplinary Research, MITI, Japan Tetsuya Saito National Research Institute for Metals, STA, Japan Nobuo Takeda The University of Tokyo B. Boro Djordjevic The Johns Hopkins University

This Page Intentionally Left Blank

vii Local Committee Members

H. Hirao, Osaka University H. Kato, Saitama University K. Yamanaka, Tohoku University T. Hattori, Hitachi, Ltd. M. Katoh, Kyushu Institute of Technology I. Kimpara, The University of Tokyo H. Kitagawa, Toyohashi University of Technology M. Kitahara, Tohoku University K. F. Kobayashi, Osaka University A. Mita, Keio University Y. Mori, Nihon University K. Nakamura, N.D.E. Forum M. Notake, Mitsubishi Research Instituete, Inc. Y. Obata, Nihon University Y. Ogura, Hitachi Construction Machinery Co., Ltd. M. Ohtsu, Kumamoto University N. Ooka, Japan Atomic Energy Research Institute M. Shiwa, Japan Power Engineering and Inspection Corporation M. Takagi, the Science University of Tokyo M. Takemoto, Aoyama Gakuin University S. Yuyama, Nippon Physical Acoustics, Ltd.

History of Previous Symposia

LOCATION: Hershey, Pennsylvania, USA TITLE: Nondestructive Methods for Material Property Determination EDITORS: C.O. Ruud and R. E. Green, Jr. PUBLISHER: Plenum Press, New York, USA (1984) LOCATION:

Montreal, Canada

TITLE: Nondestructive Characterization of Materials IT EDITORS: J.F. Bussiere, J.-P. Monchalin, C. O. Ruud and R. E. Green, Jr. PUBLISHER: Plenum Press, New York, USA (1986)

viii LOCATION:

Saarbriicken, Germany

TITLE: Nondestructive Characterization of Materials HI EDITORS: P. Holler, V. Hauk. G. Dobmann, C. O. Ruud and R. E. Green, Jr. PUBLISHER: LOCATION:

Springer-Verlag, Berlin, Germany (1989) Annapolis, Maryland, USA

TITLE: Nondestructive Characterization of Materials IV EDITORS: C.O. Ruud, J. F. Bussiere and R. E. Green, Jr. PUBLISHER: LOCATION:

Plenum Press, New York, USA (1991) Karuizawa,Japan

TITLE: Nondestructive Characterization of Materials V EDITORS: T. Kishi, T. Saito, C. O. Ruud and R. E. Green, Jr. PUBLISHER: LOCATION:

Iketani Science and Technology Foundation, Tokyo, Japan (1992) Oafu, Hawaii, USA

TITLE: Nondestructive Characterization of Materials VI EDITORS: R.E. Green, Jr., K. J. Kozaczek and C. O. Ruud PUBLISHER: LOCATION:

Plenum Press, New York, USA (1994) Prague, Czech republic

TITLE: Nondestructive Characterization of Materials VII EDITORS: A.L. Bartos, R. E. Green, Jr. and C. O. Ruud PUBLISHER: Transtec Publications, Zfirich-Uetikon, Switzerland (1996) LOCATION:

Boulder, Colorado, USA

TITLE: Nondestructive Characterization of Materials VIII EDITORS: R.E. Green, Jr., H. I. McHenry and C. O. Ruud PUBLISHER: LOCATION:

Plenum Press, New York, USA (1998) Sydney, Australia

TIILE: Nondestructive Characterization of Materials IX EDITORS: R.E. Green, Jr. PUBLISHER:

American Institute of Physics, Melville, New York, USA (1999)

CONTENTS PREFACE

INVITED PAPERS

Application of Non-Contact Ultrasonics to Nondestructive Characterization of Materials R. E. Green, Jr. Recent Advances in Structural Health Monitoring in Japanese Smart Material/Structure System Project N. Takeda

11

Sound Velocities and Microstructure H. Ledbetter

19

Recent Developments of the Laser Ultrasonic Technology and of its Applications J.-R Monchalin, A. Blouin, M. Choquet, D. Levesque and A. Moreau

27

COMPOSITES

Health Monitoring Technologies for Fiber Reinforced Plastics H. Aoyama, K. Tanaka and N. Takeda

45

Embedding Effect of SMA Wire/Optical Fiber on Damage Zone of CFRP by Acoustic Emission Signal Analysis J. H. Koo, B. K. Jang, N. Toyama and T. Kishi

53

Evaluation of Fracture Behavior in SiC/Ti-6AI-4V Composite by Acoustic Emission Technique A. Hirose and K. F. Kobayashi

61

Elastic-Stiffness Tensor of Metal-Matrix Composites Measured by Electromagnetic Acoustic Resonance H. Ogi, G. Shimoike, M. Hirao, K. Takashima and H. Ledbetter

69

Nondestructive Observation Technique for Alumina-Fiber Reinforced Plastics K. Tanaka, H. Aoyama and N. Takeda

75

ELECTROMAGNETICS AND RADIOGRAPHY

An Alternative to Film-Based Flash Radiography Using a High Speed Camera and Intensifying Screen D. Salafia, N. L. DeLeon and J. E Outlaw Investigation of Low Density Core Processes for Producing Ultra Lightweight Metals Using X-Ray Computed Tomography W. H. Green, J. M. Winter, Jr., and R. E. Green, Jr.

85

X-Ray Characterization of the Strain State in a Tensile Deformed Ti-Ni-Cu Shape Memory Alloy Y. Kishi, Z. Yajima, K. Shimizu and M. Asai

99

A New Theory of X-Ray Stress Measurement with Its Applications M. Kurita

105

Eddy Current Nondestructive Testing of Weld Zone Using Uniform Eddy Current Probe K. Koyama, H. Hoshikawa and N. Taniyama

113

A New Eddy Current Probe without Lift-off Noise H. Hoshikawa, K. Koyama and H. Karasawa

121

Microwave Measurement of Moisture in Encapsulant Resin of IC Packages Y. Ju, S. Sanpei, M. Saka and H. Ab~

129

ULTRASONICS

Precise Velocity Measurement of High Frequency Surface Acoustic Waves by Phase Velocity Scanning Method H. Cho and K. Yamanaka

135

NDE of Plates and Coatings Using Contact-Type Line-Focus Ultrasonic Probe H. Inoue, K. Kishimoto and T. Shibuya

143

Advanced On-Line Lamb Wave Inspection System Using Real-Time SSP Technique Y. Nagata, H. Yamada, Y. Konno and S. Naito

151

Reconstruction of Waveform from Photoelastic Ultrasonic Visualized Image T. Mihara, T. Musashi and K. Yamanaka

159

Nondestructive Test Field Survey for Assessing the Extent of Ettringite-Related Damage in Concrete Bridges R. A. Livingston and A. M. Amde

167

Characterization of Microstructure and Interfacial Properties of Advanced Ceramic Composites by Ultrasonics Scanning Acoustic Microscopy and Raman Spectroscopy M. H. Manghnani, M. Prasad, R V. Zinin, Y. Wang, J. Balogh, S. K. Deb, and R. Lemor

175

THIN FILMS AND COATINGS

Ultrasonic Evaluation of Remelted Zone Thickness in Aluminum Alloy Castings G. Jiang, H. Kato, Y.Yoshida and T. Komai

185

Acoustic Emission Measurement by Laser Interferometer in Ceramic Coatings M. Watanabe, M. Enoki and T. Kishi

193

Ultrasonic Evaluation of CoCrAIY Coating with Leaky Rayleigh Waves K. Kawashima, H. Fujita, T. Shima, I. Komura and T. Kubo

199

Quantitative Evaluation of SiC/NiP Composite Coating by Surface Acoustic Wave Spectroscopy and Nano-lndentation I. Ihara, T. Sawa and K. Tanaka

205

Detection of the Nano-Scale Subsurface Defects by Ultrasonic Atomic Force Microscopy T. Tsuji and K. Yamanaka

213

MODELING Modelling of Ultrasonic Attenuation in Unidirectional Fiber Reinforced Plastics S. Biwa, Y. Watanabe and N. Ohno

223

Attenuation Induced from Distributed Defects in an Elastic Solid M. Kitahara, D. Kishibe, T. Takahashi and N. Kishi

231

FDM Analysis of Ultrasonic Nonlinearity in Partially Degraded Material K. C. Kim, H. Yamawaki, T. Saito and K.Y. Jhang

239

Quantitative Evaluation of Interlaminar-Toughened CFRP Composites by Ultrasonic Micro-Spectrometer Y. Okabe and N. Takeda

247

CERAMICS AND CONCRETE Quantitative Damage Evaluation of Concrete Core Samples by Acoustic Emission M. Ohtsu and T. lida

257

Estimation of Damage Parameter and Crack Volume in Rock-Like Materials by SiGMA-AE Procedure M. Shigeishi and M. Ohtsu

265

Electromagnetic and Acoustic Emission Generated by the Cracks Creation in Solids J. Sikura, V. Lokajicek, B. Koktavy, V. Sedlakova, R Koktavy and J. Pavelka

273

Nondestructive Characterization of the Hydration Process in Tricalcium Silicate B. B. Djordjevic, L. L. Rouch, E. Shchukin, E. Amelina and I. Vidensky

279

LASER TECHNIQUES Visualization of Elastic Waves by Laser Ultrasonics J. Takatsubo, M. Imade and S. Yamamoto

289

Ultrasonic Detection Using Photorefractive Multi Quantum Well H. Hiratsuka, M. Yamamoto, H. Ikeda, S. Tsujikura, Y. Okamoto, Y. Matsuda, H. Nakano and S. Nagai

297

Compact and Robust Inspection System for Micro Cracking Detection Using LaserInduced Surface Waves M. Ochiai, N. Mukai, Y. Sano and H. Nakano

305

Detection of Defects in Micro-Machine Elements by Using Acoustic Waves Generated by Phase Velocity Scanning of Laser Interference Fringes H. Sato, S. Matsumoto, S. Nakano, H. Ogiso, H. Cho and K. Yamanaka

311

Low Frequency Laser Ultrasound under 100 KHz K. Yamanaka, H. Cho, T. Nakazawa and T. Watanabe

315

xii

MAGNETICS AND OPTICS

Quantitative Nondestructive Evaluation of a Surface Crack by a Remote Magneto-Optical Inspection System J. Lee, H. Kato, T. Shoji and D. Minkov

327

Magnetic and Optical Nondestructive Evaluations for Iron-Based Materials K. Yamada, K. Yamaguchi, S. Toyooka and Y. Isobe

333

Evaluation of Post Weld Heat Treatment Temperature of Low Alloy Steel Weld Joints Using AC Magnetic Method M. Shiwa, Y. Horii, R. Kume, H. Yoneyama and A. Yamaguchi

341

SQUID Nondestructive Damage Analysis for Austenitic Stainless Steel T. Suzuki, K. Hirano and K. W. Lee

349

Effect of Plastic Deformation on Stress Measurement Using Magnetostriction T. Yamasaki and M. Hirao

357

Visual Characterization of Wear in Large Caliber Weapons D. Salafia

365

Study of Visible and Infrared Methods of Buried Defects under Atmospheric Environment by Means of Thermal Image Techniques Y. Okamoto, A. Kamoi, K. Kondo and V. R Vavilov

371

MICROSTRUCTURE

Contactless Measurement of Induction-Hardening Depth by an Axial-Shear-Wave EMAT M. Hirao, H. Ogi and Y. Minami

379

Temper Embrittlement of Duplex Stainless Steel and its Non-Destructive Evaluation M. Katoh, K. Nishio, T. Yamaguchi, Y. Matsumoto and H. Ikeda

387

Measurement of Texture Coefficient W,oo of Aluminum Alloys by Resonance EMATS C.-S. Man, X. Fan, J. G. Morris and K. Kawashima

395

Change of Ultrasonic Attenuation and Microstructure Evolution in Crept Stainless Steels T. Ohtani, H. Ogi and M. Hirao

403

Nondestructive Evaluation of a Grey Cast Iron Using the Ultrasonic Measurement M. Mochizuki, Z. Yajima, Y. Kishi, T. Yoshida and K. Shimizu

411

AUTHOR INDEX

417

KEYWORD INDEX

421

INVITED PAPERS

This Page Intentionally Left Blank

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

APPLICATION OF NON-CONTACT ULTRASONICS TO NONDESTRUCTIVE CHARACTERIZATION OF MATERIALS ROBERT E. GREEN, JR. Center for Nondestructive Evaluation, The Johns Hopkins University Baltimore, Maryland 21218, USA

ABSTRACT Although piezoelectric ceramics have been predominantly used as ultrasonic transducer materials, a major problem associated with their use is the requirement that they need to be acoustically bonded to the test material with an acoustical impedance matching coupling medium such as water, oil, or grease. For a number of applications, a method of non-contact generation and detection of ultrasound is of great practical importance. A non-contact technique affords the opportunity to make truly non-contact ultrasonic measurements at elevated temperatures, in geometrically difficult to reach locations, in outer space, and to do this at relatively large distances from the test structure surface. KEYWORDS Ultrasonics, non-contact, lasers, interferometers, electromagnetic acoustic transducers (EMAT), air-coupled transducers, process control.

CONTACT ULTRASOUND Historically, piezoelectric crystals such as quartz were predominantly used as transducer materials. Currently, poled ferroelectric ceramics are most often used. A major problem associated with conventional ultrasonic techniques is the requirement that the piezoelectric transducers be acoustically bonded to the test material with an acoustical impedance matching coupling medium such as water, oil, or grease; or often more harmful, is the necessity of coupling the transducers to the test structure using water squirter systems. Although the couplant allows acoustical energy to propagate into the test material, it causes several problems in addition to potential harm to the material. For velocity measurements, which are necessary for material thickness measurements and to locate the depth of defects, the coupling medium can cause transit time errors on the order of one percent. Due to partial transmission and partial reflection of the ultrasonic energy in the couplant layer, there may be a change of shape of the waveform which can further affect velocity measurement accuracy. This can also lead to serious errors in absolute attenuation measurements of up to twenty percent of the measured values.

4

R.E. Green, Jr.

This latter fact is the reason that so few reliable absolute measurements of attenuation are reported in the scientific literature. It is also important to note that the character of the piezoelectric transducer itself exerts a major influence on the components of the ultrasonic signal, since conventional transducers do not respond as a simple vibrating piston and have their own frequency, amplitude, and directional response. In addition, they "ring" at their resonance frequency and it is extremely difficult to distinguish between the amplitude excursions caused by this "tinging" and the amplitude variations actually characteristic of the ultrasonic signal. They are also limited in their frequency response and, since they are in contact with the surface of the material to be tested, they can load this surface and thereby modify the ultrasonic wave itself.

NON-CONTACT ULTRASOUND A method of non-contact generation and detection of ultrasound is therefore of great practical importance. Several such techniques are presently available in various stages of development, namely capacitive pick-ups, electromagnetic acoustic transducers (EMATs), laser beam optical generators and detectors, and more recently air(gas)-coupled ultrasonic systems. However, as the name implies, capacitive pick-ups cannot be used as ultrasonic generators and, even when used as detectors, the air gap required between the pick-up and test structure surface is extremely small, which in essence causes the device to be very nearly a contact one. EMATs, on the other hand, have been successfully used for material defect characterization particularly in metal bars, tubes, pipes, and plates. One major problem with EMATs is that the efficiency of ultrasound generation and detection rapidly decreases with lift-off distance between the EMATs face and the surface of the test object, furthermore, they can only be used for examination of electrically conducting materials. Because of the physical processes involved, EMATs are much better detectors than generators of ultrasound. That is, the generator electromagnetic field needs to be very large (larger EMATs) in order to move atoms, while in detection the atoms are already moving (smaller EMATs). Laser beam ultrasound generation and detection overcomes all of these problems and affords the opportunity to make truly non-contact ultrasonic measurements in both electrically conducting and non-conducting materials, in materials at elevated temperatures, in corrosive and other hostile environments, in geometrically difficult to reach locations, and do all of this at relatively large distances, i.e. meters, from the test object surface. Furthermore, lasers are able to produce simultaneously both shear and longitudinal bulk wave modes as well as Rayleigh and plate modes. However, laser based interferometric detectors suffer from the fact that they can only be used to detect ultrasonic waves on surfaces which are relatively smooth and good reflectors of light at the same wavelength as the laser used in the optical interferometric detector. Air(gas)-coupled ultrasonic systems have been under development for some time and they are currently being optimized for practical non-contact ultrasonic applications. These systems are more similar to conventional contact ones and, when optimized, will play an important role in modem nondestructive evaluation. Non-contact ultrasonics systems, including hybrid ones, which use laser generation/EMAT detection or laser generation/air-coupled detection, have proven to be optimum for a number of materials characterization applications where the material to be investigated is harmed by water, grease, oil or other couplants, in hostile environments such as high temperatures or space, and

Non-contact Ultrasound Characterization

where the test objects are difficult to access with conventional transducers. Applications include C-scans of wooden panel paintings, lumber, composite prepregs, composite panels, graphite/epoxy tape placement, inspection of steel specimens in high temperature furnaces, inspection of aircraft panels, and inspection of railroad rails. It should be noted that an all aircoupled system (generation/detection) has not been considered due to the fact that air is the worst impedance matching material to couple ultrasound into a workpiece. A schematic of contact and non-contact ultrasonic techniques for materials inspection is shown in Fig. 1.

Fig. 1. Contact and non-contact ultrasonic techniques.

C-SCANS OF WOODEN PANEL PAINTINGS A variety of nondestructive methods have been used to detect voids, hidden cracks, and fine fractures in wooden panel paintings. These defects promote severe stress concentrations which are aggravated by the mechanical constraints imposed on the wood by the framing and mounting systems. The panel's structural condition needs to be mapped in a nondestructive fashion in order to predict how it will mechanically respond over time or in different environments. Decisions can then be made as to whether a panel can be safely removed from storage, if it can withstand conservation treatment, if it can be continuously displayed, or if it can be safely shipped to exhibitions. A through-transmission air-coupled ultrasonic system was used to perform C-scan images of a variety of wooden panel paintings. A unipolar, burst pulser (3 ohm impedance, 450 V) generated the signal which was then sent through an impedance-matching network (high-pass inverter). Both air transducers were 0.5 MHz with 1-in. diameter faceplates and 2 in. focal lengths and were positioned 1 in. away from the wood panel to be examined. One transducer sent the signal through the panel and the other received it after transversing the panel. The non-contact air-coupled system permitted defects to be seen that were not observed by x-radiography or other means.

6

R.E. Green, Jr.

C-SCANS OF COMPOSITE PRE-PREGS AND PANELS A similar air-coupled transmission C-scan system was used to examine the quality of as manufactured graphite/epoxy prepregs and multi-layer graphite/epoxy panels which contained controlled amounts of porosity.

Fig. 2. Air-coupled transmission ultrasonic C-scan of 16 ply multi-layer graphite/epoxy panel which contained controlled amounts of porosity.

COMPOSITE PROCESS CONTROL Composite materials are normally manufactured by one of two processes: (1) hand layup and autoclave cure, (2) filament or tow placement with a winding machine (sometimes multiple axis robotic controlled) and cured in place. The most important process requiting monitoring is the curing process for both thermoset and thermoplastic composites. The two parameters currently monitored are the temperature and the pressure applied to the filaments, tows, or laminates. Commonly this is done with conventional thermocouples and pressure gauges. The only other device used to monitor the cure process is a dielectric sensor mounted between two composite plys in autoclave curing which measures the ionic conductivity. Recently a non-contact hybrid ultrasound system using a pulsed laser as the source and an aircoupled transducer as a detector has been successfully used to monitor the mechanical tow placement system for graphite/epoxy composites. This system measures the velocity (elastic modulus and density) and the attenuation (resin viscosity) of the wave propagating through the composite during the cure process. The tow placement process is used in aerospace applications where large-scale non-autoclave processing of high-temperature thermoplastic composites is required. In order to improve the overall quality and performance of the final part, real-time process control is necessary. A system of laser generation of narrow-band surface waves and detection using air-coupled transduction is optimally suited for this application, since they are confined to the top few layers of the composite surface and, therefore, interact more strongly with the tow placement process than bulk longitudinal or shear waves. The incident laser pulse was delivered through a fiber-optic bundle to enhance safety and flexibility. The fiber-optic bundle was able to deliver Nd:YAG pulsed light at an average energy as high as 55 mJ/pulse for a pulse duration as short as 10ns, at a 20 Hz repetition rate, without signs of fiber damage. The narrow band generation of surface waves was accomplished

Non-contact Ultrasound Characterization

by spatially modulating the light using a periodic transmission mask. The center frequency of the generated toneburst signal was easily controlled by changing the spacing of the illumination sources, which permitted control of the surface wave penetration depth, since it is determined by the frequency of the wave. In addition, the spatial modulation method allows more light to illuminate the test surface while retaining the thermoelastic generation condition, which increases the signal detectability without material damage. Also, the signal narrow banding provides enhanced immunity to the broadband noise. A schematic diagram of the experimental system is shown in Fig. 3. Figure 4 shows a comparison between the ultrasonic signals detected in the cases of good and poor consolidation for a carbon/PEEK tow placed unidirectional cylinder. By monitoring the attenuation of the center peak in the wave frequency spectrum, a variety of other defects in this test piece were successfully detected, including grafoil inserts and delaminations. S ~ecimen k

Lens ~OpticalFiber

10ram \

.../'~

Mask ~ ~ !

Trigger

L =0.7mm

~

( ~ I Air-coupled Transducer

~

Data AcquisitiOnpc [~.[

.....

Scope

Fig 3. Schematic of hybrid laser generation/air-transducer detection ultrasonic system for control of thermoplastic composite tow placement.

40

50

60

Time (~s)

(a)

70

80

40

50

60

Time (~s)

70

80

(b)

Fig. 4. Non-contact ultrasonic waveforms as detected for carbon/PEEK composite: (a) good consolidation, (b) poor consolidation (two ply deep).

8

R.E. Green, Jr.

RAILROAD RAIL INSPECTION Maintenance of railroad rails is one of the greatest problems facing the transportation industry today. Despite the fact that a variety of inspection techniques have been used since the very early days of the introduction of railways, none of them is satisfactory for detection of many possible defects. Ranging from simple visual inspection, to magnetic induction, and water coupled ultrasonics, the techniques currently used are relatively low technology and far from optimized. All defects internal to the rail obviously cannot be detected by visual inspection and many defects on the surface of the rail are missed because of surface coverings of dirt, grease, or other foreign matter. Magnetic induction suffers from "lit~-off" problems (magnetic field strength decreases with distance from surface of rail) and can only be used for surface breaking cracks. Contact ultrasonics using piezoelectric transducers in rolling rubber wheels filled with water or oil can detect both surface and internal horizontal cracks (parallel to rail top surface), but cannot detect transverse cracking, vertical cracking, or any cracks inclined to the top surface. Research is currently underway in the Johns Hopkins CNDE to develop a non-contact ultrasound technique for railroad rail inspection. Preliminary results have shown that both laser generation (55 mJ pulsed laser) and detection and laser generation/air-coupled detection can operate effectively at distances of several inches from metallic surfaces which are sufficient to prevent contact with any rivets, tie-downs, or other obstacles attached to the rail. Equally important is the fact that pulsed laser systems generated surface waves, longitudinal, and shear waves in a single pulse, thus providing much more information in the same time frame.

Fig. 5. Laser based non-contact ultrasonic system for railroad rail inspection.

NON-CONTACT ACOUSTIC EMISSION DETECTION Although detection of the acoustic waves emitted from materials as a result of mechanical stressing of chemical attack (acoustic emission) has been a viable nondestructive testing technique for over 40 years, it has not shown it's full potential since the precise characteristics of the acoustic waves emitted from each type of source is not known. For this reason, a very quiet tensile test machine was used to deform a variety of metal specimens while the surface

Non-contact Ultrasound Characterization

displacement associated with internal metallurgical alterations were detected by means of a noncontact optical interferometer. In these tests, the dimensions of the gauge sections of tensile specimens were chosen to be very small so that any microstructural alteration would be visible on the gauge section surface and could be examined in detail using both optical and scanning electron microscopy. Because of the large fiat frequency bandwidth (0 - 60 MHz) of the optical probe, acoustic emission signals were obtained in a frequency regime not normally detected with conventional transducers. Tests run on a series of stainless steel specimens revealed a large number of acoustic emission signals at 9 - 10 MHz prior to fracture. Scanning electron microscopic examination of the fracture surface resulted in a one-to-one correspondence between these high frequency signals and the fracture of intermetallic particles in the steel.

Fig. 6. (a) Acoustic emission frequency spectrum from non-contact laser detection during tensile test of steel specimen, (b) optical image of slip lines on surface at yield, (c) optical image of voids on fracture surface, (d) electron microscope image of broken intermetallic particle causing high frequency acoustic emission.

10

R.E. Green, Jr.

CONCLUSIONS

New developments in laser ultrasonics, electromagnetic acoustic transducers (EMAT) and air(gas)-coupled ultrasound will permit design and construction of extremely efficient noncontact ultrasonic systems to detect fatigue damage and corrosion damage in aircraft and other structures. Not only will non-contact ultrasonic systems detect all defects currently detected by water, oil, or grease coupled ultrasonic ones but, perhaps more importantly because of the noncontact nature, permit ultrasonic attenuation detection of microcrack formation, incipient corrosion and other material changes prior to macrocrack and extended corrosion formation. The non-contact nature of laser acoustic detectors makes them optimum for making measurements in materials at elevated temperatures, in corrosive and other hostile environments, in geometrically difficult to reach locations, in outer space, and do all of this at relatively large distances, i.e. meters, from the test structure surface.

REFERENCES

10.

11.

Murphy, J., Glass, J.T., Majerowicz, S., and Green, Jr., R.E. (1990) Materials Evaluation 48, 714. Huber, R.D. and Green, Jr., R.E. (1991) Materials Evaluation 49, 613. Huber, R.D. and Green, Jr., R.E. (1991). In: Proceedings ofJANNAF NDE Subcommittee Meeting, Huntsville, AL, CPIA Publication 559, pp. 15-21. Murray, A., Green, R.E., Mecklenburg,M.F. and Fortunko, C.M. (1991). In: Nondestructive Characterization of Materials IV, pp. 73-80, Plenum Press, New York. Green, Jr., R.E. (1995). In: Proceedings of the Air Force 3rd Aging Aircraft Conference, Wright-Patterson Air Force Base, Dayton, OH, pp. 548-578. Murray, A., Mecklenburg, M.F., Fortunko, C.M., and Green, Jr., R.E. (1996) Journal of the American Institute of Conservation 35, 145. Green, Jr., R.E. (1999). In: Proceedings of the Second Joint NASA/FAA/DoD Conference on Aging Aircraft, NASA/CP-1000-20892, pp. 244-251. Lanza di Scalea, F. and Green, Jr., R.E. (1999) Ultrasonics 37, 179. Lanza di Scalea, F., Bern&, T.P., Spicer, J.B., and Djordjevic, B.B. (1999) IEEE Transactions on Ultrasonics, Ferroelectrics, and Frequency Control 46, 1551. Lanza di Scalea, F., Berndt, T.P., and Green, Jr., R.E. (1999). In: Nondestructive Characterization of Materials IX, pp. 616-621, Green, R.E. (Ed). American Institute of Physics, New York. Lanza di Scalea, F. and Green, Jr., R.E. (1999)ExperimentalMechanics 39, 329.

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

11

~ C E N T ADVANCES IN STRUCTURAL HEALTH MONITORING IN JAPANESE SMART MATERIAL/SRTUCTURE SYSTEM PROJECT Nobuo TAKEDA

Department of Advanced Energy, Graduate School of Frontier Sciences, and Department of Aeronautics and Astronautics, School of Engineering, The University of Tokyo, 7-3-1 Hongo, Bunkyo-ku, Tokyo 113-0033, Japan

ABSTRACT The present R&D status of our structural health-monitoring group (SHMG) is summarized. Activities of the SHMG are conducted as a university-industry collaboration program at the University of Tokyo along with 10 research organizations. The research themes include: (1) development of high-performance sensor system technology with newly-developed sensors, (2) development of a damage detection and self-diagnosis system for structural integrity based on micro-mechanical damage identification, and (3) development of application technology for model structures. The sensor technology includes: (1) development of small diameter optical fiber sensor, (2) damage suppression in composite laminate systems with embedded shapememory alloy films, and (3) development of maximum strain memory sensors with electrically conductive composite systems. The sensor output is correlated with the underlying damage evolution in structures such as aircrafts, satellites, high-speed trains and large-scale civil infrastructures. KEYWORDS Structural health monitoring, composites, microscopic damage, optical fiber sensor, shapememory alloy, electrically conductive polymers, aerospace structures, civil infrastructures.

INRODUCTION Light-weight composite material systems have been progressively used as structural members in various applications. However, more use has been proposed as primary structural members under severe operating conditions. In such applications, durability evaluation and health monitoring systems are two key technologies to be investigated.

12

N. Takeda

The author has been proposing a so-called "experimental micro-mechanics of composites" to bridge the gap between material fabrication and macroscopic mechanical properties. Based on in-situ observation using optical, scanning electron, and/or scanning acoustic microscopes with loading devices, microscopic deformation and damage has been quantified. Moreover, theoretical models have been established for damage evolution. These efforts can provide the methodology for the durability evaluation or the damage tolerance design of composites [ 1-9]. In real structural applications, however, since the strains applied to the composite structures are random and uncertain, the real-time strain monitoring is necessary to predict the present damage status in composites based on the above durability evaluation method. Moreover, if the damage can be detected by using sensors, more reliable estimation of the damage status or the residual life can be made. In Japan, a new "R&D for Smart Material/Structure System (SMSS)" project started in October 1998 as one of the Academic Institutions Centered Program supported by NEDO (New Energy and Industrial Technology Development Organization), Japan. This SMSS project includes four sub-themes: (1) structural health monitoring, (2) smart manufacturing, (3) active/adaptive structures, and (4) actuator materials development. The author acts as a group leader in the structural health monitoring group, which consists of 10 research organizations. Our structural health monitoring group (SHMG) is currently developing a health monitoring system, which conducts a real-time damage detection and self-diagnosis as well as damage control in light-weight composite structural systems. The research themes include: (1) development of high-performance sensor system technology with newly-developed sensors, (2) development of a damage detection and self-diagnosis system for structural integrity based on micro-mechanical damage identification, and (3) development of application technology for model structures. The sensor output is correlated with the underlying damage evolution in structures such as aircrafts, satellites, high-speed trains and large-scale civil infrastructures. The latest results in SHMG research efforts are reported for each research theme.

SUMMARY OF RECENT RESULTS OF SHMG Detection of Transverse Cracks in CFRP Laminates with FBG Sensors- Univ. Tokyo

Fiber Bragg Garting (FBG) optical fiber sensors have been developed to measure strain, temperature and so on, through the shift of the wavelength peak in the reflected light. A uniform strain within the gage section (typically 10 ram) is normally assumed. When an FBG sensor is embedded in 0-degree ply to detect transverse cracks in a 90-degree ply, a nonuniform strain distribution due to the initiation and evolution of transverse cracks causes the wavelength distribution in the reflected light (Fig. 1) [10]. Figure 2 shows the stress and crack

Structural Health Monitoring in Japan

13

density as a function of strain measured by a strain gage and a FBG sensor. A careful examination of the wavelength distribution can be used to detect the evolution of transverse cracks in composite laminates (Fig. 3). With increasing transverse crack density, the shape of the reflection spectrum was distorted; the intensity of the highest peak became small, some peaks appeared around it, and the spectrum became broad. As the crack density was close to saturation, the spectrum became narrow again and the highest peak recovered its height. These experimental observations can be well explained by the theoretical prediction [10].

Fig. 1. Cross-sections of CFRP cross-ply laminate with an embedded FBG sensor (a) normal, and (b) parallel to loading

Fig. 2. Stress and crack density versus strain measured by strain gage and FBG sensor

Fig. 3. Reflection spectra at various tensile strains corresponding to (A)-(F) in Fig. 2

Development of Small-Diameter Optical Fiber Sensors- Hitachi Cable The first notable accomplishment in our university-industry collaboration is the development of small-diameter optical fibers and the application to the FBG sensors. The developed optical fiber is with 40 gm in cladding diameter and 52 gm in polyimide coating diameter, which is easily embedded in one CFRP ply of 125 gm in thickness. Such optical fibers have mechanical properties similar to those of conventional optical fibers with 125 ~tm in cladding diameter and do not cause any reduction in strength of composites when embedded parallel to reinforcing fibers in laminas [11]. The developed FBG sensors also have optical sensitivities similar to

14

N. Takeda

those of conventional 125 ~tm-cladding diameter FBG sensors. The polyimide coating is highly compatible with epoxy or other high-temperature polymer matrix of CFRP composites under high-temperature exposure during fabrication and also in high-temperature use.

Fig. 4. Small-Diameter OFS

Fig. 5. Small-diameter optical fiber embedded within lamina

Identification of Impact Damage Parameters in Composites Using Embedded Optical Fiber Sensors- Kawasaki Heavy Industries Real-time detection of impact load on composite laminates was successfully made with embedded small-diameter optical fibers [12]. Figure 5 shows a small-diameter optical fiber embedded parallel to reinforcing carbon fibers in [02/902]s cross-ply laminates. Instrumented Charpy impact tests were made, and the optical loss during impact was recorded as a function of time simultaneously with the impact load and the surface strains measured by strain gages (Fig. 6). The bending loss can be observed only during impact loading. The magnitude of optical loss is found to be proportional to the magnitude of the impact load. 300

........ Load ..... ----9

Z "o

S"

200

"

Strain ( 0 deg.) Strain (90 deg.) Optical f i b e r 1 Optical fiber 2 .

. . . . . . Opucaz nDe~

I 1.5 ] ]

I

1

o ..J

-o.c:

~6 ~.

-~v) ,-

100

E 0

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

o.5~

i __ ZO

-5

0

5 Time(ms)

10

o

15

Fig. 6. Impact responses of normalized optical intensity, impact load and specimen surface strains in [02/902] s cross-ply laminates

Strain and Damage Monitoring of CFRP Space Satellite Structures Using Optical Fiber Sensors - Mitsubishi Electric

Structural Health Monitoring in Japan

15

Honeycomb sandwich panels with thin CFRP face sheets are normally used as space satellite structures, and are subjected to severe thermal and mechanical environmental conditions. Lowcost fabrication is highly required for future satellites and a structural health monitoring system is essential during fabrication as well as in practical use [13]. Figure 7 shows a specimen for simple compression test of a honeycomb sandwich panel where a small-diameter optical fiber is embedded within CFRP face sheets. The change in optical intensity can be an excellent indicator of compression failure in such structures as shown in Fig. 8.

Fig. 7. Compression test of sandwich panel

Fig. 8. Optical power loss by compression failure

Damage Suppression and Control in CFRP Laminates with Embedded SMA F i l m s - Fuji Heavy Industries >.,

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

<1 -0.002

7

0.70 Aa

-0.003

" """

0.6 - . .

i .... 0

i .... I .... 500 1000 t (min)

i. 1500

0.5

Fig. 5. Companion to Fig. 3. Recovery of the shearwave attenuation (4.7 MHz) and the low-frequency internal friction (0.2 Hz) after quenching. The solid line is from the Granato-Hikata-Lticke theory. The dashed line is from the Cottrell-Bilby-Harper model. -1 represents internal friction at t - 24 min. Qt=24

~

6

.>,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-

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Fig. 10. Companion to Fig. 9. Dependence of longitudinal sound velocity on frequency. Dispersion arises from aluminum layers separating B-Al-layer packets.

VOIDS" SHAPE CHANGE AND STIFFNESS DECREASE Figure 11 from Ledbetter et al. [8] shows the effects of porosity on Young moduli of sintered aggregates of titanium spheres. With increasing porosity, elastic stiffness decreases dramatically. And pore shape changes from near spherical toward oblate ellipsoidal (disc). The figure shows that the accompanying internal friction increases strongly, resembling Rayleigh scattering. We lack a good model for this void-interface contribution to internal friction.

A PARTICLE-REINFORCED COMPOSITE: SCATTERING The theoretical literature on acoustic-wave scattering is enormous, with many results transferring from quantum theory and electromagnetic-wave theory. In 1986, Ledbetter and Datta [9] gave a acoustical-wave-propagation theory based on the ensemble average of multiple scattering from ellipsoidal particles. This theory has many advantages: 1. In the long-wavelength limit, it gives the static elastic-stiffness coefficients, the Cij. 2. It gives the frequency-dependent Cij. 3. It considers any ellipsoidal particle shape [discs, spheres, rods (fibers)]. 4. It gives the imaginary parts of the Cij, which arise from scattering. 5. At least for the Cij, it makes correct predictions over the entire particle-concentration range. Although we applied this theory successfully to many and various composites to predict their elastic constants, we applied it only cursorily to scattering problems. Here, we summarize some

Sound Velocities and Microstructure

25

of the theory's predictions. (We made specific calculations for SiC particles in an aluminum matrix.) 1. Attenuation varies as f6; f = frequency. 2. Attenuation varies as (abc)2; a,b,c = ellipsoid axis lengths. 3. Transverse waves scatter more than longitudinal waves. 4. Maximum attenuation occurs near a volume fraction of 0.2. 5. For the same volume fraction, larger particles scatter more. 6. Spheres scatter most, then discs, then rods. 7. Lower matrix stiffness increases damping and moves the maximum toward lower volume fractions. 8. Higher particle stiffness produces the same effects noted in 7. Figure 12 shows the graphical results corresponding to 6 in this list.

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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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Fig. 3" Theoretical sensitivity versus ultrasonic frequency for several confocal Fabry-Perot configurations operating either in transmission or in reflection (cavity length = 50 cm).

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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Recent Developments in Laser Ultrasonics 2.2

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CHARACTERIZATION OF A PAPER WEB We have also demonstrated the application to the characterization of the mechanical properties of a paper web [28]. The on-line measurement of such properties is of great interest for the optimization of the paper making process. This application is based on the generation at one location of Lamb waves and on their detection at another location separated from the first one by a known distance. The measurement of the time of propagation provides a measurement of the velocities, which in turn allow the determination of several elastic moduli. These elastic properties are known to correlate empirically with paper strength. We have demonstrated unlike other reported works [29-31 ] that generation can be performed without any damage to the sheet by using a TEA CO2 laser. To be able to detect efficiently the So wave, which corresponds essentially to in-plane motion, the collection direction was tilted with respect to the normal to the sheet. In this case both So and A0 are actually detected. The velocity of So in a given direction is related to the in-plane Young modulus in this direction. In the paper industry it is customary to use instead the square of the in-plane extensional velocity (called the Tensile stiffness Index: TSI) and to plot the variation of the TSI with direction. The inplane extensional velocity is the velocity of the So mode at low frequencies. Fig. 13 shows a plot of the TSI for a copy paper sheet measured in static condition and moving at a velocity of 11 m/s. The two sets of data are in good agreement, which shows that the technique can be used on-line in a paper mill. The A0 mode is also detected at the same time. At low frequencies this mode becomes non-dispersive and merges into a membrane wave, the velocity of which is directly related to the applied tension. We have also demonstrated that web tension, which is also a parameter of great interest in the paper making process, could be measured as well [28].

40

J.-P. Monchalin et aL (Km/s)2

eo

1tO 110 (Km/s)2

r o Static RunnJngpaper~ paper

9

eo

Fig. 13: Plot of the TSI of static and running paper.

CONCLUSION We have presented several developments of the laser ultrasonic technology and its applications recently performed at the Industrial Materials Institute. The performances obtained with the two-wave mixing photorefractive demodulator have been compared with those of the more established confocal Fabry-Perot. The detection of small defects has been addressed by a numerical processing approach based on the Synthetic Aperture Focusing Technique. Laser-ultrasonics combined with SAFT can be used on non-planar specimens and can be applied to the detection of defects in welds and diffusion bonds. We have also reported applications to the inspection of polymer-matrix composites, the detection of corrosion in lap joints of aircraft structures, the characterization of the microstructure of steel and monitoring of its evolution with temperature and the characterization of a paper web. In conclusion, laserultrasonics offers a wide range of applications in many industrial fields ranging from the detection of defects, to measurement of elastic parameters and the characterization of microstructure. Efforts will continue not only to improve basic technology, but also to move applications from the laboratory to industrial and on-line demonstration (which is the case presently of several applications related to steel) and then ultimately to commercialization (which is only presently the case of the inspection of polymer-matrix composites).

Recent Developments in Laser Ultrasonics

41

REFERENCES

.

.

10. 11. 12. 13. 14. 15. 16. 17. 18. 19.

20. 21.

Scruby, C. B. and Drain, L. E. (1990). Laser-Ultrasonics: Techniques and Applications. Adam Hilger, Bristol, UK. Monchalin, J.-P. (1993). In: Review of Progress in Quantitative Nondestructive Evaluation vol.12, pp. 495-506, Thompson D.O. and Chimenti D.E. (Eds). Plenum Press, New-York. Monchalin, J.-P., N6ron, C., Bussi6re, J.F., Bouchard, P., Padioleau, C., H6on, R., Choquet, M., Aussel, J.-D., Camois, C., Roy, P., Durou, G., Nilson, J. A. (1998) Advanced Performance Materials 5, 7. Bussi6re, J.F., Dubois, M., Moreau, A. and Monchalin, J.-P. (2000). In: Nondestructive Characterization of Materials IX, Green Jr, R.E. et al, (Eds), Plenum Press, New-York, in press. Monchalin, J.-P. (1986) IEEE Trans. Ultrason. Ferroelectrics and Freq. Cont. UFFC33,485. Monchalin, J.-P., H6on R. (1986) Materials Evaluation 44,1231. Monchalin J.-P., H6on R., Bouchard P., Padioleau C. (1989) AppL Phys. Lett. 55, 1612. Blouin, A., Delaye, P., Drolet, D., de Montmorillon, L.-A., Launay, J.-C., Roosen, G., Monchalin, J.-P. (1998). In: Nondestructive Characterization of Materials VIII, pp.1319, Green Jr., R.E. and Ruud, C.O. (Eds). Plenum Press, New-York. Connes, P. (1958) J. Physique et le Radium 19, 262. Ing, R.K. and Monchalin, J.-P. (1991) Appl. Phys. Lett. 59, 3233. Blouin, A. and Monchalin, J.-P. (1994) AppL Phys. Lett. 65, 932. Delaye, P., Blouin, A., Drolet, D., de Montmorillon, L.-A., Roosen, G., Monchalin, J.P. (1997) J. Opt. Soc. Am. B 14, 1723. Delaye, P., Blouin, A., de Montmorillon, L.-A., Drolet, D., Monchalin, J.-P., Roosen, G. (1997) Proc. SPIE, 3137,171. Monchalin, J.-P. and H6on R. (1992) US patent # 5,080,491. Wang, X., Littman, M. G., McManus, J. B.,Tadi, M., Kim, Y. S., Askar, A., Rabitz, H. (1996) J. Appl. Phys. 80,4274. Noroy, M.-H., Royer, D., Fink, M. (1993) Appl. Phys. Lett. 63, 3276. Lorraine, P. W., Hewes, R. A., Drolet, D. (1997). In: Review of Progress in Quantitative Nondestructive Evaluation vol. 16, pp. 555-562, Thompson, D. O. and Chimenti, D. E. (Eds), Plenum Press, New-York. Blouin, A., L6vesque, D., N6ron, C., Drolet, D., Monchalin, J.-P. (1998) Optics Express 2, 531. L6vesque, D., Blouin, A., N6ron, C., Choquet, M., Monchalin, J.-P. (1999). In: Review of Progress in Quantitative Nondestructive Evaluation, vol. 18, pp.309-316, Thompson, D.O. and Chimenti D.E. (Eds) , Kluwer Academic/Plenum Publishers, New-York. Monchalin, J.-P., N6ron, C., Bouchard, P., Choquet, M., H6on, R., Padioleau, C. (1997) Canadian Aeronautics and Space Journal 43, 34-38. Fiedler, C. J., Ducharme, T., Kwan, J. (1997). In: Review of Progress in Quantitative Nondestructive Evaluation vol. 16, pp. 515-522, Thompson, D. O. and Chimenti, D. E. (Eds), Plenum Press, New-York.

42 22.

23. 24.

25.

26. 27.

28. 29. 30. 31.

J.-P. Monchalin et al. Nelson J.M., Nerenberg R.L., Rempt, R.D., Shepherd, W.B., Sorscher S.M., Woodmansee, W.E., Choquet M., Monchalin, J.-P. (1998), report to US Air Force D950-10322-1. Massabki, M., Choquet, M., L6vesque, D., Monchalin , J.-P. (2000) Canadian Aeronautics and Space Journal, in press. Monchalin, J.-P., Blouin, A., Drolet, D., Bouchard, P., H6on, R., Padioleau, C. (1997) In: Proceedings of the 39th Mechanical Working & Steel Processing Conference, vol. 35, pp. 927-931, Iron & Steel Society, Warrendale, PA. Dubois, M., Moreau, A., Dawson, A., Militzer, M., Bussi6re, J.F. (1998). In: Intelligent Processing of High Performance Materials, pp. 11-1 to 11-9, NATO RTO Workshop I. Dubois, M., Moreau, A., Militzer, M., Bussi6re, J.F. (1998) Scripta Materiala, 39, 735. L6vesque, D., Moreau, A., Lord, M., Dubois, M., Monchalin, J.-P., Padioleau, C. and Bussi6re, J.F. (1999). In:Advanced Sensors for Metals Processing, pp 53-66, Brusey, B., Bussi6re, J.F., M. Dubois, M. and Moreau, A. (Eds), Canadian Institute of Mining, Metallurgy and Petroleum, Montr6al, Canada. Blouin, A., Reid, B., Monchalin J.-P., (2001). In: Review of Progress in Quantitative Nondestructive Evaluation, vol. 20, Thompson, D.O. and Chimenti D.E. (Eds) , Kluwer Academic/Plenum Publishers, New-York, to be published. Brodeur, P.H., Johnson, M.A., Berthelot, Y.H., Gerhardstein, J.P. (1997), J. Pulp Paper Science 23, 238. Ridgway, P.L., Hunt, A.J., Quinby-Hunt, M., Russo, R.E. (1999) Ultrasonics, 37, 395. Walter, J.B. and Telshow, K.L. (2000). In: Review of Progress in Quantitative Nondestructive Evaluation, vol. 20, Thompson, D.O. and Chimenti D.E. (Eds) , Kluwer Academic/Plenum Publishers, New-York, in press.

COMPOSITES

This Page Intentionally Left Blank

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

45

HEALTH MONITORING TECHNOLOGIES FOR FIBER REINFORCED PLASTICS

H. AOYAMA and K. TANAKA

Mechanical Engineering Research Laboratory, Hitachi, Ltd. 502 Kandatsu, Tsuchiura, Ibaraki 300-0013, Japan N. TAKEDA

Center for Collaborative Research, The University of Tokyo 4-6-1 Komaba, Meguro-ku, Tokyo 153-8904, Japan ABSTRACT A new defect-monitoring technique using light transmitted through alumina/epoxy has been developed. Defects inside a semi-transparent alumina/epoxy laminate scatter the light and decrease its intensity. The spectra of the transmitted and reflected light were measured by using a spectrometer in order to identify the wavelengths most sensitive to defects in alumina/epoxy. As stiffness decreased, the intensity of the transmitted light decreased and that of the reflected light increased. These intensity changes were most obvious at wavelengths from 400 to 800 nm, so measuring the intensity of transmitted visible light between these wavelengths is a promising method for assessing the structural integrity of alumina-reinforced epoxy.

KEYWORDS Reinforced Plastics, Laminated Construction, Delamination, Ultrasonic Inspection, Composite Material, Transmitted light technique, Bending stiffness

INTRODUCTION In recent years, a lot of equipment that includes superconducting magnets has been developed. The superconducting coil, which is the main part of theses magnets, is kept in liquid helium. To connect the superconducting coil vessel (kept at liquid-helium temperature) with an outer vessel (kept at ambient temperature), load-support systems that have low thermal conductivity, as well as high stiffness and strength, are required. Fiber-reinforced-plastic (FRP), because it meets these requirements, has been used to make these systems. The most recent loadsupport systems are made of alumina fiber-reinforced epoxy (alumina/epoxy). Safe operation of the superconducting magnet will be jeopardized if the mechanical stiffness of the load-support systems begins to decrease. It is therefore necessary to utilize some form of continuous nondestructive-evaluation (NDE) monitoring in order to detect any changes in the mechanical "health" of the systems. If possible, it would be highly desirable to simultaneously assess the health of load-support systems without removing them from service. So in this study, a NDE technique, high-resolution ultrasonic scanning, was used to detect defects inside FRPs. And a fundamental experiment was carried out using light transmitted through FRP [1]-[5]. It was found that transmitted light technique produces similar images to those of ultrasonic scanning, but it can operate under more stringent constraints.

H. Aoyama, K. Tanaka and N. Takeda

46

APPLICATION OF TRANSMITTED LIGHT TECHNIQUE TO NDE Defects such as matrix cracking, delamination, and fiber breakage can occur in alumina/epoxy. Stiffness of a load-support system made of alumina/epoxy is an important criteria and must be kept constant to prevent irregular vibration of the superconducting coil. To understand what causes these defects and how they grow, a four-point-bending fatigue test on a flat alumina/epoxy specimen was conducted at room temperature. A schematic of the specimen is shown in Fig. 1. The test was carried out under a controlled compression load and the load amplitude was monitored. Bending stress calculated from the load and the bending stiffness K (where K=EI, where E is Young's modulus, I is moment of inertia of area.) of the specimen. The relation between the stress amplitude and the number of cycles to failure Nf is shown in Fig. 2. The bending stiffness K of the specimen was calculated from the measured load and the displacement amplitude. The relationship between cycles normalized by the number of cycles to failure Nf and the stiffness decrease ratio /1K/Ki (where Ki is initial stiffness) under several bending stress amplitudes /1 a is shown in Fig. 3. When /1K/Ki reaches 0.3, fiber breakage occurrs rapidly.

Fig. 1 Laminate specimen under four-point bending

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Fig. 3 Bending-stiffness decrease ratio of alumina/epoxy

The alumina/epoxy specimen used in this test is thin enough that light can be transmitted through it. If defects exist inside the FRP laminates, the light will scatter, and its intensity will decrease. To confirm this scattering phenomenon, we conducted a basic experiment, the setup for which is shown in Fig. 4. A test piece is mounted between a light source and a fiber scope. Transmitted light is detected by a CCD camera, and its signal is converted to a digital signal by a frame grabber board. A CCD image of the light transmitted through the specimen during the fatigue test is shown in Fig. 5. These images show that the shadow near the loaded points grew as stiffness decreased. It was confirmed after the test that this shadow was clearly caused by matrix cracking and delamination.

Fig. 4 Setup for transmitted light experiment

Ultrasonic methods are one of the most widely used methods for non-destructiveevaluation. The specimen was therefore tested using a 15-MHz ultrasonic transducer with a 50mm focal length; a pulse-echo mode was used and a C-scan image was taken. The C-scan image of the alumina/epoxy specimen is shown in Figs. 6. Comparing Fig. 5 and 6 clearly shows that the CCD and C-scan images are similar. The relationship between the damaged area ratio and the bending stiffness decrease ratio is shown in Fig. 7. The corresponding results of the ultrasonic scan are shown in the same figure. Both sets of results clearly fit the line.

48

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

Fig. 5 Transmitted light images of alumina/epoxy ( ,5 a =202 MPa)

Fig. 6 Ultrasonic scan images of alumina/epoxy ( A a =202 MPa)

Health Monitoring Technologiesfor Fiber Reinforced Plastics .o

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SPECTRA OF TRANSMITTED L I G H T A damage detection system for FRP structures in superconducting magnets must satisfy several constraints: 9No heat transfer between sensor and structure at low temperature 9Function in strong electromagnetic field 9Compact and low cost A transmitted light technique will satisfy all these constraints in detecting damage in FRPs. But using normal light causes heat generation of the materials because it includes infrared light. To determine which wavelength is sensitive to matrix cracking and delamination in alumina/epoxy, the spectra of the transmitted light of the alumina/epoxy were measured by a spectrometer (Shimazu AT-100HG) during the fatigue testing, and the area between the loaded points of the specimen was also measured. Figure 8 shows the relationship between wavelength and transmissivity. The transmissivity decreases as the stiffness decreases. In particular, the transmissivity in the visible region (400-800 nm) drops remarkably. Figure 9 shows the relationship between the wavelength and the total reflectivity: total reflectivity increases as stiffness decreases. In particular, the transmissivity in the visible region (400-800 nm) changes remarkably. 50

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

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The relationship between stiffness decrease ratio ,5 K/Kin and transmissivity change A T at a 600-nm wavelength is shown in Fig. 10. This relationship is clearly linear. So the decrease in mechanical stiffness can be predicted by measuring the decrease in transmissivity. Moreover, this optical wavelength (600 nm) does not produce heat in the test structure.

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OPTICAL SCAN SYSTEM To observe location and size of damages in alumina/epoxy, an optical scan system was made. Its setup is shown in Fig. 11. It consists of a halogen lamp unit, two spectrometers, and a X-Y stage. The specimen is held on the X-Y stage and moved in 0.1-mm steps. Maximum scan size is 100 X 100 mm. Light from a halogen lamp goes into the spectrometer and a selected

Health Monitoring Technologiesfor Fiber Reinforced Plastics

51

FRP and goes into another spectrometer. Optical power is measured in this spectrometer and the signal with information of the scanned location is fed into a PC. In the PC, amplitude of the signal is converted to a gray scale.

Fig. 11 Optical scan system

An alumina/epoxy specimen which includes delamination was scanned by this optical scan system. The image from the 650-nm light is shown in Fig. 12 and that from the 350-nm light is shown in Fig. 13. It is clear that the 650-nm wavelength can clearly reveal the delaminaton; however, the 350-nm wavelength only shows low-level contrast.

Fig. 12 Optical scan image ( 2 =650 nm)

52

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

Fig. 13 Optical scan image ( 2 =350 nm)

SUMMARY

Fundamental experiments for "health monitoring" of alumina-fiber-reinforced plastics conclude that the amplitude of a visible ray decreases remarkably as bending stiffness of FRP decreases. Therefore the transmitted-light technique offers a means of monitoring the stiffness decrease of a FRP structure such as a load-support system in a strong electromagnetic field. Accordingly, an optical scan system, which uses a selected-wavelength visible ray to detect delamination in almina/epoxy, was developed. This study was partially supported by a grant from the New Energy and Industrial Technology Development Organization (NEDO).

REFERENCES

Benaron, D. A. (1994) Laser Focus World, 79-82. Benarson, D. A., Lenox, M. A., and Stevenson, D. K. (1992) SPIE 1641, 35-45. Chance, B. (1992) SPIE 1641, 162-169. Benarson, D. A. and Stevenson, D. K. (1993) SCIENCE 259, 1463-1466. Wang, L., Ho, P. P., Liu, C., Zhang, G., and Alfano, R. R. (1991) SCIENCE 253,769-771.

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

53

EMBEDDING E F F E C T OF SMA W I R E / OPTICAL FIBER ON DAMAGE ZONE OF CFRP BY ACOUSTIC EMISSION SIGNALANALYSIS

J.H. Koo, B.K. Jang, N. Toyama and T. Kishi

National Institute for Advanced Interdisciplinary Research, 1-1-4 Higashi, Tsukuba, [baraki 305-8562, Japan ABSTRACT We manufactured SMA wire and optical fiber embedded CFRP with [0/90/0/90]s in order to investigate the embedding effect on fracture process of the host material. In this work, we did compact tension test and detected acoustic emission signal and evaluated the damage zone of the specimens by anisotropic AE source location method. The used elastic wave for AE source location was extensional mode of Lamb wave. There are some reports in which flexural mode of Lamb wave was adopted to do AE source location in plate. They eliminated the extensional mode with wavelet transformation or cross correlation method. However, because the wave velocity with a frequency range similar to that of generated AE signal is required, flexural mode with frequency dependency is not proper to do source location. Therefore, in order to do accurate AE source location, we should use extensional mode that is not dependent on frequency instead of flexural mode. The AE signals detected in present work were mostly Lamb waves which had large extensional mode generally due to fiber breakage or bundle breakage that gave fatal effect to fracture of host material. In addition, from AE source location result we found that AE was not generated in vicinity of SMA wires or optical fibers. Then, it was found that the embedded SMA wires or optical fibers had not any bad effect on degradation of maximum load and damage zone of host materials. KEYWORDS CFRP, SMA, optical fiber, acoustic emission, source location, Lamb wave, extensional mode

INTRODUCTION Recently many researches about development of smart materials or smart structures which give self sensing and self recovery function to carbon fiber reinforced plastics (CFRP) that is used in aerospace and spatial fields have been reported. A number of attempts are prosperously performed, which give health monitoring function and self recovery function and vibration control function to host materials by embedding sensors (e.g. optical fiber or PZT patch) and actuator (e.g. SMA wire or PZT patch) [13]. However, because the embedded materials have thermal expansion coefficient different from that of host materials and bonding energy with host materials is weak, the interface is easy to become the origin of crack. In addition to, there are several fracture mechanisms like matrix crack, delamination, debonding, fiber breakage and bundle breakage in CFRP. Therefore it is required to investigate influence of crack generated in the interface on fracture of host materials. In present work, we manufactured SMA wire and/or optical fiber embedded CFRP and investigated effect of them on fracture process zone of host material. The used test was compact tension test. The

54

J.H. Koo et al.

generated AE signals during the test were detected and analyzed for source location, that is, evaluation of damage zone. The used velocity for source location is that of extensional mode Lamb wave. LAMB WAVE VELOCITY Because the thickness of the specimen is thin (about 1.5 ram), the elastic wave due to emission of energy due to crack generation propagates as Lamb wave. Lamb wave constitutes of two components. One is extensional mode which velocity is fast. The other one is flexural mode whose velocity is slow [4]. The typical wave form is shown in Fig. 1 [5]. In the case of composite plate, there are some report about AE source location using flexural mode [6,7]. Because wave velocity with the same frequency band as that of the generated AE signal is required for source location, however we should not use flexural mode of which velocity is variable according to its frequency but extensional mode of which velocity is not variable according to change of frequency. In addition to, in the case of analyzing AE due to fiber breakage or bundle breakage, which has generally large extensional mode in the front part of signal, velocity of extensional mode is useful. In Fig. 2, samples of AE signal detected from [0/90/0/90]s CFRP produced in this research are shown. Fig. 2 (a) shows AE signal that is generated by 0.5 mm pencil lead break on the surface and propagates 30 mm in fiber direction. Fig. 2 (b) shows AE signal detected in the same location condition as Fig.2 (a) during CT test. For the most AE signals detected during the test has the same pattern as Fig. 2 (b), we decided that they were due to fiber breakage or bundle breakage and used extensional mode for source location.

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I 300

I 350

T i m e , ~ts Fig. 1. AE signals generated by pencil lead break (a)on the surface of CFRP and (b)on the edge of CFRP (5).

-4 -5

|

0

i

i

I

20

.

,

.

i

40

,

,

,

i

60

,

,

,

i

,

,

80

T i m e , its Fig. 2. AE signals (a) generated by pencil lead break on the surface of CFRP and (b) detected during CT test for CFRP.

,

i

100

55

Damaged Zone of CFRP Studied by AE Analysis

"••

6500, o

A E sensor

o

E

Experimentall

Io

6000

0

.,., r

o ~ 5500 • E

m 0

*.-, 5000 o ,.~

. ,....

2 E I

I

I

[

[

[ ~ 5mm sharp pencil

!

1--, I-~i

i

,,~ ,..,i m.-l-~l |

,

0

,

,

I

30

,

,

45

Angle, degree

i

m.~, m.- I

10cm lcm Fig. 3. Schematic view of measurement of Lamb wave velocities.

I

15

Fig. 4. Velocities of extensional mode of Lamb wave.

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

1- c~ 80)(V0 - V45)+ V45 10

0< 0 <2~r

(1)

Here, V 0 is velocity in 0 direction, V 0 is for 0 degree direction, V45 indicates for 45 degree direction. ANISOTROPIC AE SOURCE LOCATION The denotations for numerical analysis are as follows [8]. P (x,y) 9approximate coordinates of AE source Pi(xi,Yi) 9location of i-th sensor Di: distance between P and Pi r" radius of sensor Vi 9velocity of P, Pi direction t" approximate occurring time of AE T i 9empirical arrival time in i-th sensor t ~" approximate arrival time of AE in i-th sensor If x, y and t are selected as initial values of coordinates of AE source and occurring time and size of sensor is considered, t i is expressed by the following equation. t~ = fi(t,x,y) = t + (D i - r) / V i

(2)

J.H. Koo et al.

56

If the corrected value of (t,x,y) is (t+At,x+Ax,y+Ay), the corrected value of ti becomes ti*.

ti

- fi

(t + ht, x + Ax, y +

-.s (t#,y) +

ofi at

At+

hy)

ofi ox

~+

ofi -~y

ay

(3)

= t~ + aiAt + b i A x +ciAY

Here,

afi

a i = ~

at'

bi =

afi

afi

--~x'

ci = ~ . ay

When we let the difference of T i and t i• ei, Y. ~

becomes as follows.

i=l

i= 1

i= 1

= ~ (T i - ti) 2 - 2(T i - ti) ( a i At + b i Ax + C i Ay )

(4)

i=1

+ ( ai At + bi Ax + ci AY )2 The necessary condition that Y. ~i2

has minimum is as follows.

i=l

i=1 OAt

i=1

OAX

i=1

aAy

(5)

(i=~laiai)At+(i=~laibi)Ax+(~laici)AY=~lai(Ti-ti) (i=~l biai)At+(i=~ 1 bibi)Ax+(.~=l biCi)AY= .~=lbi(Ti-ti)

(6)

(i=~lciai)i~tq'(i=~lCibi)Axq'(~__lCiCi)Ay=~=lCi(Zi-ti)

EXPERIMENTAL METHOD For manufacturing the host material, prepreg ( t : 200 #m, carbon/epoxy, Hexel Co. ) was stacked in [0/90/0/90]s and held in 180 *C and 0.3 MPa during 2 hours. 4 types of specimens are produced : 1) CFRP only, 2) SMA wire ( 0 : 200 #m, TiNi system, Daidou Special Steel Co., Japan ) embedded CFRP, 3) optical fiber ( r : 250/zm, 3M Co. ) embedded CFRP and 4) both embedded CFRP. Each wire and fiber are embedded in the interval of 5 mm. Especially in the case of 4th type specimen wire and fiber are embedded in turns. Fig. 5 shows stacking configuration and Fig. 6 shows dimensions of CT specimen. In Fig. 7 experimental method is shown. Cross head speed is 0.5 mm/min. AE signals were detected by four AE sensors ( r : 4 mm, model : 304, resonance frequency : 180 kHz, Fuji Ceramics Co. ) which have internal head amp ( 24 dB ) and amplified with -13 dB amplifier. Sampling rate is 20 MHz. The

Damaged Zone of CFRP Studied by AE Analysis

Fig. 5. Geometry of stacking of embedded CFRP specimen.

Fig. 6. Geometry of specimen fq tection test (unit :mm).

Fig. 7. Schematic diagram of CT test and AE signal detection. numerical analysis for anisotropic AE source location is achieved in personal computeI RESULTS AND DISCUSSIONS

Behavior of load and AE Fig. 8 shows the relation of time, No. of event, load and amplitude of first peak. In ea shows linear behavior up to about 1.5 kN but shows nonlinear behavior after that. Amp] AE is below 1.8 V. AE with larger amplitude is, however, observed when load decrease: From maximum load to final fracture there is several step that shows catastrophical dec Maximum load is higher in the case of embedded specimens but it seems to be included i experimental error.

Observation of fractured specimens Photographs of fractured specimens are in Fig. 9. In all types, fracture surface is rough al tip bundle fractures are observed. In the case of type 1, the bundle with 3 mm width is ol

58

J.H. Koo et al.

Fig. 8. Photographs of fractured specimens : (a) CFRP, (b) CFRP + SMA wires, (c) CFRP + optical fibers and (d) CFRP + SMA wires + optical fibers.

Fig.9. Relation of Time-Load-No. of Events-Amplitude. interface. In the case of type 3, 1st and 4th fiber from notch tip are fractured but 2nd and 3rd fiber are pulled out. It seems to be due to low tensile strength and its brittleness. For the case type 4, any wire and fiber is not fractured. Damage zone evaluation by A E source location

The AE source location results are shown in Fig. 10. In all cases AE signals were detected only until maximum load because sensors fell apart by shock due to catastrophical decrease of load. Generally in the vicinity of SMA wires and optical fibers the probability of occurrence of AE due to delamination

Damaged Zone of CFRP Studied by AE Analysis 6O

40

I

(b)

(a)

..~

59

4O .!

~o

~i~ .:~." .:.

20

,o

0

0

10

20

30

40

50

0

60

. . . .

0

I

. . . .

!0

I

. . . .

20

I

.

,

,

30

i

,I.

,

,

I

40

I

1 1

9 I . .h.~ ..'.1, I ....................................... " ~ i : 1. !". .

i,,... I ........................................ ~ l ' r ' r

.7.

~'," [

i

:

1

I. ! I

0

'

1!

I 0

. . . .

0

I

I0

. . . .

I

20

. . . .

I

30

.

.

!

I

60

'(d)

(c)

9

. . . .

50

i .

I I . . . .

40

I

50

. . . .

I

60

0

. . . .

0

I

I0

. . . .

I

20

. . . .

I

30

. . . .

I I

40

.

~

i

i

I

50

. . . .

,

60

Fig. 10. AE source location results ; (a) CFRP, (b) CFRP + SMA wires, (c) CFRP + optical fibers and (d) CFRP + SMA wires + optical fibers (unit" mm).

or debonding is high. However, all wave forms are like that of Fig. 2 (b), that is, have large extensional mode. The Lamb wave with large extensional mode is generally due to matrix cracking, fiber breakage and bundle breakage. But the emission energy of matrix cracking is very small as compared with fiber breakage and bundle breakage. Therefore the source of all AE detected during the test seems to be fiber breakage and bundle breakage which have a fatal effect on fracture. From the results we found that all AE were concentrated in notch tip. In Fig. 10 (b),(c) and (d), we compared the AE location with that of SMA wire and optical fiber. In order to facilitate comparison, only two vertical lines that indicated embedded wire or fiber were drawn in figures. In Fig. 10 (b), concentration of AE is observed in the vicinity of the first line but there is > 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

(a)

Composite ecimen

i

J ' i

i

_i

_i~l

1~ 20-'-1b':

J-

I

J

J

..._--

_l~i_ 20 80

(b)

t -~

This section cut away alter gr ippi ng fixed by epoxy I cement i Si C | her

r-~-~t Grip

-'-lb'- 20~-]Co~v area er skin

11111

Fig. 1. Shape of tensile test specimens: (a) for composite (t=-0.35mm for 1-ply and 1.50mm for 10-ply); (b) for SiC fiber.

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.

(a)

(b)

1500 ~-

I

A A.

30

:E 1000v

|

o

o

100

_= m ii0

- 20~

3

O" 0

o c o

500-

J 0o

,t- 50

-10 o uJ <

/ I

1.0

0.5 Strain

o

0

0

(%)

~

2.0 4.0 6.0 8.0 10.0 AE amplitude (V)

Fig. 6. AE behavior in tensile test (one-ply composite annealed at 1123K for 230.4ks): (a) AE activity during tensile test; (b) frequency distribution of AE amplitude. 1500

AE a m p l i t u d e Below 4V

I~

~:~Over 4V

o

m

/

9

"

r

20 "i 4,J

~ 1000o im __r'7

500

-

0

22

10 .--.~

0.5

I

1.0

.~

o o w

0

Strain ( % )

Fig. 7. AE activity during tensile loading (one-ply composite annealed at 1123K for 230.4ks).

A. Hirose and K.F. Kobayashi

66

Figure 6 (a) shows a stress-strain curve and AE behavior in the tensile test for an annealed oneply composite specimen. The AE signals are classified into two groups by their energy; i.e., below 4V and over 4V as shown in Fig. 6 (b). Figure 7 shows AE event counts classified by the AE energy in the tensile test for an annealed one-ply composite specimen. In this case possible sources of the AE events are considered to be as follows: i) fracture of the fibers ii) cracking of the reaction zone iii) debonding at the interface of fiber/reaction zone. In the tensile test the load was released when stress reached 90% of the average tensile strength of the composite. After the tensile test, observation of the fibers extracted from the composite revealed that no fibers fractured. It is assumed from the result that the sources of the AE events are cracking of the reaction zone and debonding at the interface of fiber/reaction zone. The former may cause the AE events having lower energy below 4V and the latter may cause the AE events having higher energy over 4V. The results means that debonding at the interface of fiber/reaction zone can occur before the cracks in the reaction zone reach the fibers, causing fracture of the fibers. AE measurements were also performed during tensile tests of annealed ten-ply composite specimens. Figure 8 shows a typical example of the results. The distribution of the AE amplitude is similar to that of the one-ply composite shown in Fig. 6 (b), consists of two distinct groups having lower AE energy below 4V and higher one over 4V, respectively. Since the AE behavior during the tensile test shown in Fig. 8 (b) is also similar to Fig. 7, the ten-ply composite is known to exhibit the same fracture behavior as the one-ply composite. From these results, it is considered that reaction zone is debonded from fibers before the cracks in the reaction zone reached the fibers, and so the factor (2) is not a reason of the strength degradation. The average tensile strength of as-received fibers and extracted fibers from the composites are shown in Table 1. The tensile strength of the extracted fibers from the as-fabricated composite was almost the same as that of the as-received fibers. The tensile strength of the extracted fibers from the annealed composite decreased. Thus the tensile strength of reinforcing fibers in the composite was degraded by the annealing. The tensile strength degradation should be due to

(a)

(~

I AS an~nude / J"l BelOW 4V

~9

s w

2000

IOO0

2OO

!

r g

100 I I

m 0

2.0 4.0 6.0 8,0 AEImldlWdl(V)

10.0

o

0

0.5

Strain

(%)

1.o

0

Fig. 8. AE behavior in tensiletest (ten-ply composite annealed at 1123K for 57.6ks): (a) AE activity during tensiletest; (b) frequency distribution of AE amplitude.

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

67

increasing defects of the fibers. Figure 9 shows appearances of an as-received fiber and extracted fibers from the annealed composites. While the as-received fiber has smooth surface, the extracted fibers have rough surfaces. The progress of surface damages due to the annealing was observed. When bare as-received fibers were annealed in a vacuum chamber, the surface of the fibers was not damaged and the tensile strength was not degraded. Therefore the flaws on the extracted fiber from the annealed composites were caused by the heterogeneous growth of the reaction zone as shown in Fig. 10.

Table 1. Tensile test results of SiC fibers.

Number of samples Average tensile strength (MPa)

Extracted fiber from composite

As-received fiber

As-fabricated

1123Kx32.4ks annealed

1123Kx129.6ks annealed

44

36

26

35

4964

4849

4350

2266

Fig. 9. SEM images showing surfaces of SiC fibers: (a) as-received fiber; (b) extraced fiber from as-fabricated composite; (c) extracted fiber from annealed composite at 1123K for 32.4ks; (d) extracted fiber from annealed composite at 1223K for 129.6ks.

Fig. 10. SEM image of F/M reaction zone seen in composite annealed at 1223K for 230.4ks.

A. Hirose and K.F. Kobayashi

68

Tabel 2. Ratio of measured composite strength and estimated one by the rule of mixture.

As-fabricated

1123Kx32.4ks annealed

1223Kx129.6ks Annealed

0.903

0.954

0.996

Measured strength R.O.M. strength

Table 2 shows the ratio of the measured strength of the composite/the calculated strength by the rule of mixture (R.O.M.) using the strength of the extracted fiber. The R.O.M. strength was defined by the following equation: o u = ofVf + Om(1-Vf) (1) where Ou : composite strength, of: fiber strength, Om: strength of matrix, Vf: fiber volume fraction. The ratio of the measured strength/the R.O.M strength is more than 90%. This suggests that the degradation of the strength of the composite is caused by the degradation of fibers themselves. Therefore, in the present case the previously mentioned factor (3) is valid for the mechanism causing the degradation of the composite by the annealing.

CONCLUSION The strength of the SiC/Ti-6A1-4V composites significantly decreases when the reaction zone thickness exceeds 1~tm. It is assumed from fractography analysis and AE analysis that a fibermatrix 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.

REFERENCES 1. 2. 3. 4. 5. 6. 7.

Rhodes, C.G. and Spurling, R.A. (1985) Recent Advances in Composites in the United States and Japan, ASTM STP 864, 585. Ochiai, S., Osamura, K. and Murakami, Y. (1983) Z Metallkd., 74, 68. Ochiai, S., and Murakami, Y. (1982) Z Metallkd., 73, 229. Ochiai, S., and Murakami, Y. (1979) J. Mater. Sci., 14, 831. Ochiai, S. and Osamura, K. (1983) Metallurgical Transactions., 18A, 673. Smith, P. R. and Froes, F. H. (1984) J. Metals, March, 19. Soumelidis, S, Quenisset, J. M. and Naslain, R. (1986) J. Mater. Sci., 21,895.

Nondestructive Characterizationof MaterialsX Green et al. (Eds) Published by Elsevier Science Ltd., 2001

69

ELASTIC-STIFFNESS TENSOR OF METAL-MATRIX COMPOSITES MEASURED BY ELECTROMAGNETIC ACOUSTIC RESONANCE

H. OGI*, G. SHIMOIKE*, M. HIRAO*, K. TAKASHIMA**, and H. LEDBETTER***

*Graduate School of Engineering Science, Osaka University Toyonaka, Osaka 560-8531, Japan **Precision and Intelligence Laboratory, Tokyo Institute of Technology, Yokohama, Kanagawa 226-8503, Japan ***Materials Science and Engineering Laboratory, National Institute of Standards and Technology, Boulder, CO 80303, USA ABSTRACT We developed a contactless acoustic-resonance technique to measure all independent elastic stiffnesses of solids. We call this method electromagnetic acoustic resonance (EMAR). The measurement setup includes a solenoid coil, into which a rectangular-parallelepiped specimen is inserted, and permanent magnets to apply uniform static magnetic field. Using the Lorentz-force coupling, we measured resonance frequencies of free vibrations without coupling agents. Controlling the Lorentz-force direction by changing the field direction, we independently excited and detected one vibration group among eight groups of free vibration of a rectangular parallelepiped. This is a significant advantage in identifying the vibration modes of the observed resonance frequencies and yielding the complete set of elastic stiffnesses by solving an inverse calculation. We applied this method to silicon-carbide fiber-reinforced Ti-alloy-matrix composites and determined all elastic-stiffness components. KEYWORDS Elastic Stiffnesses, Electromagnetic Acoustic Transducer, Metal-Matrix Composite, Non-contact Measurement, Resonance

INTRODUCTION Fiber-reinforced metal-matrix composites are promising materials for advanced applications because they retain high strength and toughness at high temperatures. Their elastic properties, especially elastic stiffnesses, are indispensable to designing such an engineering structure. However, two principal problems prevent us from measuring exactly the complete set of elastic stiffnesses: (1) such a fiber-reinforced composite exhibits low elastic symmetry and possesses many independent stiffnesses, typically five to nine. (2) The conventional pulse-echo method, which is useful to determine a particular stiffness, is inapplicable to some stiffnesses because of wave scattering and mode conversions at the interface between the fiber and matrix. In this study, we present a contactless acoustic-resonance method for measuring all independent elastic stiffnesses using a single small specimen. We call this method electromagnetic acoustic resonance (EMAR). The interaction between the electromagnetic field and mechanical vibration allows us to make the contactless coupling. With this method, we determined all components of

70

H. Ogi et al. X2

/,,

a

b

X2

Xl

X3

(a) Material U

(b) Material X

Fig. 1 Coordinate systems of SiCf/Ti composites. the elastic-stiffness tensor of SiC-fiber-reinforced Ti-alloy-matrix composites. The measurements agreed favorably with results by a pulse-echo method. MATERIAL We used two composites made of polycrystalline Ti-6A1-4V matrix reinforced by continuous SCS-6 SiC fibers. One is reinforced unidirectionally and the other alternately as shown in Fig. 1. We call the former Material U and the latter Material X. These are candidate materials for components of jet-engine or aerospace structures because of their high-temperature strength, stiffness, and toughness [ 1]. They were fabricated by the foil-fiber-foil technique (8 plies) at 1173 K with 65 MPa compression. The fiber diameter was 140 gm and the fiber volume fraction was 0.35. The average grain size of the matrix was 10 gm. Figure 2 shows the microstructure of Material X. For both composites, we prepared rectangular-parallelepiped specimens with dimensions given in Table 1. We measured mass density by the Archimedes method, which also appears in Table 1. Material U exhibits orthorhombic symmetry with nine independent elastic stiffnesses Cij: -Cll

C12

C~3

0

0

0

C 22

C23

0

0

0

C33

0

0

0

C44

0

0

C55

(1)

0 C66

For Material X, we expect tetragonal symmetry because of the crossply structure and the nine Cij in Eq. (1) are reduced to six because Cl1=C22, C44=C55, and C13=C23.

Table 1. Specimen dimensions and density. Material U Material X

a (mm) 3.986 4.977

b (mm) 1.815 4.015

c (mm) 3.023 1.905

p(kg/m 3) 3886 3930

MEASUREMENT Figure 3 shows a typical EMAR-measurement setup. We insert the rectangular-parallelepiped specimen into the solenoid coil, which is located between the permanent magnets. Driving the

Electromagnetic Acoustic Resonance of MMC

71

Fig.2 Microstructure of Material X.

compeer ]

permanentmagnet ~ J ~ sohmoid coil

~,--~.i=====~-[ 9

superheterodyne speedometer

solenoidcoil

X3

/

Fig. 3 Typical EMAR measurement setup.

Fig. 4 EMAR measurement setup for B2ggroup.

solenoid coil by tone bursts, eddy currents arise on the specimen surfaces and interact with the static magnetic field from the magnets to produce the Lorentz forces. The Lorentz forces cause free vibrations of the specimen. The same coil detects integrated deformations over the surfaces through the reverse-Lorentz-force mechanism. Sweeping the frequency of the driving bursts and measuring the deformation amplitude, we obtain the resonance spectrum. The free vibrations of a rectangular parallelepiped with orthorhombic symmetry fall into eight groups depending on the deformation symmetry [2, 3]'. Because the Lorentz-force direction is determined by the vector product of the magnetic field and eddy currents, we can control the deformation symmetry and select a vibration group by changing the geometrical configuration between the static field and the solenoid coil. For example, in Fig. 4, the magnetic field is applied along the x3 axis and the Lorentz forces occur on the Xl-X2 faces along the Xl axis. The Lorentz forces cause the Xl-component displacement ul. This displacement can be detected most effectively by the same coil through the reverse-Lorentz-force mechanism when Ul is an odd function about x3 and an even function about Xl and x2. Among the eight groups of free vibration, only the B2g vibration group (torsional vibration about x2 axis) satisfies this condition [2]. Thus, only B2g vibration modes are generated and detected with this configuration. In this study, we independently excited and detected Big, B3g (torsional vibration about xl axis), and Ag (breathing vibration) groups for Material U; and B2g, B3g, and Ag groups for Material X. The EMAR configurations for other groups appear in Refs. 4 and 5. Resonance frequencies of free vibration can be calculated using the specimen dimensions, mass

H. Ogi et al.

72 |

!

i

I

Ag-9 Ag-12Ag-14

i LA-lal

Ag-3

<

I

I

I

I

B2g group

'

I

I

I

]

B2g-3t B2g-5

B3g-2 B3g-1

B3g-5

B3g-3

t Big-6 JLJ .

0.5

I

B2g-8

B2g-1

9

I

B2g-4

B2g-2

<

I

,

I

i

,

,

.

.

.

J

.

I

,

,

1.5

1

Frequency (MHz) Fig. 5 Resonance spectra measured by the EMAR method for Material 22. Ag denotes the breathing group, and Bzg and B3g denote torsional vibration groups about x2 and x3 axes, respectively [2, 3]. density, and all independent elastic stiffnesses [2]. Because the first two are easily obtainable, an iterative approach allows us to find a complete set of Cij that leads to the closest resonance frequencies to the measurements. The key to deriving the correct Cij with such an inverse calculation is to find an exact correspondence between the calculated and measured resonance frequencies (mode identification) [6]. If we make an incorrect mode identification, the inverse calculation fails to converge or converges to a physically meaningless answer. Conventional resonance spectroscopy using piezoelectric transducers [2, 7] usually excites all the eight vibration groups simultaneously. This results in a large number of resonance peaks and makes mode identification very difficult. However, because the mode-selective principle with the EMAR method described above can measure each vibration group separately, it simplifies mode identification and provides a reliable set of Cij. For comparison, we applied the pulse-echo method to some of the Cij. We used a longitudinal wave and two shear waves. All ultrasonic waves were propagated in the thickness direction. They provided C22, C66, and C44 for Material U; and C33 and C44 for Material X. The pulse-echo method applied to other propagation directions was unsuccessful because the surface areas were small. RESULTS AND DISCUSSION Figure 5 shows the measured resonance spectra for Material X. Different EMAR configurations led to a different spectrum, as expected. According to the mode-selection principle, we identified

73

Electromagnetic Acoustic Resonance of MMC Table 2 Measured (fEMAR) and calculated (/c,lc) resonance frequencies for Material U. Mode notation follows Mochizuki [3 ]. The average difference offEMAR-fc,lc was 0.4 %.

fEM~(MHz)

fo.,o (MHz)

diff. (%)

Big-1 Big-2 Big-3 Big--4 Big-5 Big-6 Bl~7 B1~8

0.633727 0.761856 1.112402 1.189886 1.286432 1.427158

0.639325 0.767144 1.11038 1.179541 1.282397 1.429456 1.604195 1.64696

0.88 0.69 -0.18 -0.87 -0.31 0.16

Ag-1 A~-2 Ag-3 Ag-4 Ag-5 Ag-6 Ag-7 Ag-8

0.674023

0.671491 0.798739 0.954882 1.244947 1.486785 1.60392 1.66483 1.672069 1.67851 1.828107 1.864525 1.944108 2.010206

-0.38

Mode

1.638866

0.958798 1.490086 1.594972 1.673748

A~-9 Ag-10 Ag-ll Ag-12 Ag-13

1.83982 1.85286 1.938973 2.007576

0.49

-0.41 -0.22 0.56 -0.1 -0.64 0.63 0.26 0.13 -0.11 0.23 0.46 -0.07 -0.58 -0.36

1.772487

0.551056 0.918502 1.057635 1.263364 1.309443 1.393513 1.552009 1.650926 1.781171

1.996798

1.999062

0.11

B3g-1 B3g-2 B3g-3 B3g-4 B3g-5 BN -6 B~-7 B3g-8 B~-9

0.551647 0.916413 1.052841 1.264239 1.317111 1.398568

B3~-10

0.49

Table 3 Elastic stiffnesses (GPa) of SiCf/Ti-6A1-4V composites measured by EMAR and pulse-echo (PE) methods. B denotes bulk modulus and E Young's modulus. Cll

C22

C33

C44

C55

C66

C12

C13

C23

E1

E2

E3

B

Material Ua EMAR 190.2 191.0 249.8 55.93 54.09 54.49 75.58 70.29 67.57 151.6 153.8 213.7 117.6 PE

~

192

--

57.3

~

54.2

.

.

.

.

.

.

.

MaterialXb EMAR 224.7 224.7 187.8 54.59 54.59 51.77 79.32 76.90 76.90 181.2 181.2 147.5 122.5 PE

m

m

188

53.7

.

.

.

.

.

.

.

.

.

aFiber-longitudinal direction is along the X 3 axis and the ply-layer interfaces are normal to the X2 axis. b The ply-layer interfaces are normal to the x3 axis. Fibers lie along the Xl and x2 axes.

the resonance modes and made the inverse calculation developed in the previous study [8] to derive all independent elastic stiffnesses. Table 2 compares the measured and calculated resonance frequencies after convergence of the inverse calculation for Material U. For all modes, they agreed within 1%. The typical rms error was 0.4 %. Table 3 presents the elastic stiffnesses determined by the EMAR and pulse-echo methods. The EMAR Cij and pulse-echo Cij agreed within a few percent. Concerning Material U, the

74

H. Ogi et aL

measurement results Cl1~C22, C44~C55, and C13~C23 suggest that the composite approximately shows a tetragonal symmetry, which is expected from the manufacturing procedure. For transverse isotropy, the additional requirement is C66=(Cll-C12)/2. This was observed within 5 %. The simple linear rule-of-mixture enables us to estimate Young's modulus along fiber (E3) of Material U using previously reported Young's moduli of isotropic Ti-6A1-4V matrix (E~=I 11 GPa [9] ) and the SCS-6 fiber (Ec= 400 GPa [ 10]): E3 = cmEm+cfEf.

(2)

Here, Cmand cf denote matrix and fiber volume fractions, respectively. Equation (2) leads to E33= 212 GPa, which favorably agrees with the present result in Table 3. (Note that Eq. (2) is valid only for E33 and for no other elastic stiffnesses.) From the fiber alignments, we expected that C22 for Material U equals C33 of Material X. This was correct within 1.5 %. CONCLUSIONS The present study demonstrated the EMAR capability of measuring the complete set of elastic stiffnesses of fiber-reinforced composites. The mode-selective principle worked well just by changing the geometrical configuration between the magnets and the solenoid coil, which can be a significant advantage for making an exact mode identification. The determined Cij agreed well with those determined by the pulse-echo method. ACKNOWLEDGMENT S. Kim (NIST) provided the pulse-echo results and much other valuable help. Professor IF'.Bowen (University of Birmingham) provided the materials. REFERENCES

.

6. 7. 8. 9. 10.

Wadsworth, J. and Froes, F. (1989) JOM 41, 12. Ohno, I. (1976)J. Phys. Earth 24, 355. Mochizuki, E. (1987) J. Phys. Earth 35, 159. Ogi, H., Takashima, K., Ledbetter, H., Dunn, M., Shimoike, G., Hirao, M., and Bowen, P. (1999) Acta Mater. 47, 2787. Ogi, H., Dunn, M. L., Takashima, K., Ledbetter, H. (2000) J. Appl. Phys. 87, 2769. Ogi, H., Ledbetter, H., Kim, S., and Hirao, M. (1999) J. Acoust. Soc. Amer. 106, 660. Migliori, A. and Sarrao, J. (1997) Resonant Ultrasound Spectroscopy, Wiley, New York. Ledbetter, H., Fortunko, C., and Heyliger, P. (1995) J. Appl. Phys. 78, 1542. Naimon, E, Weston, W., and Ledbetter, H. (1974) Cryogenics 14, 246. Ward, G., Herrmann, D., and Hillberry, B. (1995) J. Comp. Tech. Res. 17, 205.

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

75

NONDESTRUCTIVE OBSERVATION TECHNNIQUE FOR ALUMINA-FIBER REINFORCED PLASTICS

K. TANAKA*l, H. AOYAMA*I, and N. TAKEDA*2 91Hitachi, Ltd., Mechanical Engineering Research Laboratory, 502 Kandatsu, Tsuchiura, Ibaraki 300-0013 Japan 92 Graduate School of Frontier Sciences, the University of Tokyo 4-6-1 Komaba, Meguro-ku, Tokyo 153-8904, Japan

ABSTRACT To detect the damage of almina-fiber-reinforced plastics (ALFRP), we studied the transmitted-light technique, FBG sensors, and acoustic emission. They can detect the internal damage in ALFRP structures. It was found that the transmitted-light technique was the best method for detecting the damage of the ALFRP: it can measure the exact size of damage and observe the internal damage easily and directly.

KEYWORDS ALFRP, internal damage, transmitted-light technique, FBG sensor, acoustic emission

INTRODUCTION There are several possible techniques to detect damage in fiber-reinforced plastics (FRP). Some examples are edge-replica, X-ray imaging, scanning electric microscope, scanning acoustic microscope, thermography, holography, strain gauges, and acoustic emission. Moreover, fiber bragg grating sensors (FBG) and light-transmitted imaging have been developed recently. The superconducting magnet is used in extremely low temperature, and the load support system of the superconducting coil is usually made of almina-fiberreinforced plastics (ALFRP). It is covered with a coil case, so that the size of the damage detectors has to be small, and the damage should be detected from the outside of the case. A scanning acoustic microscope or X-ray devices are too large to fit in the coil case. The damage

76

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

detector is used under a high magnetic field, and then an electric strain gauge is highly influenced by the surroundings. The edge-replica method or the scanning electric microscope is inadequate for detecting internal damage of materials because these methods observe only the surface of materials. Table 1 summaries the advantages and disadvantages of the various damage investigation methods. We studied the transmitted-light technique, FBG sensors, and acoustic emission, because they can detect the internal damage in ALFRP structures.

Table 1 Comparison of various damage investigation methods Method

Size

Mounting ease

Visibility of

Anti-magnetic field

internal damage Edge replica

Small

Bad

Bad

X-rays imaging

Large

Bad

Excellent

Excellent N/A

Scanning electric microscope

Large

Bad

Bad

N/A

Scanning acoustic microscope

Large

Bad

Excellent

N/A

Thermography

Medium

Bad

Excellent

N/A N/A

Holography

Medium

Bad

Good

Strain gauge

Small

Excellent

Bad

Bad

Acoustic emission

Medium

Good

Excellent

General

Fiber bragg grating sensor

Small

Excellent

Excellent

Excellent

Transmitted light technique

Small

Excellent

Excellent

Excellent

N/A: not applicable

Fig. 1

Setup of transmitted-light system

NDE Techniquefor ALFRP

77

TRANSMITTED-LIGHT TECHNIQUE ALFRP has a milky-white color and is semi-transparent. Its degree of internal damage can be determined from the reduction in transmitted-light intensity. A transmittedlight system is composed of a CCD camera unit and a light source (Fig.l). Figure 2 shows a transmitted- light image and explains the procedure for mapping the cracks. First, cracks are distinguished using the brightness threshold of the transmitted-light. Then, every crack position is measured closely and each crack length is calculated. And any measurements of intersecting cracks are removed. Next, the crack density, which is the number of cracks per unit length, is calculated. Finally, all cracks are mapped.

Thickness: 1 mm Epoxy resin, r m almina-fiber

[0~ 150 80

15

~10

,

I,,,,,j,,,~li,,,,,Ihl,ll,I,l~,l,;I,,l,,,,,,l,,I,,,.... ,,1,, t

~: Loading direction

20

40

60

80

Distribution of cracks (mm)

Observation area

(a) e = 0.47% .~::iiii!ii!~i.

[

.

.

.

~."~

(1) Transmitted-Light image

[

"

l '

~ II, ;al[~,,,fl,l,:~l,ltlhl,~l[l l~I~,J:l,t,l,a'Jl,,I],,l,t?,,l

~.

0

20

40

Distribution of cracks (mm)

I Imageprocess I 9Crack distinction from a brightness threshold 9Measurement of crack position 9Calculation of crack length 9Calculation of crack density 9Mapping of all cracks

(b) e = 0.60% ~0

~

O0

I1,'

I i 20

40

_ 60

80

Distribution of cracks (mm) (c) e = 0.72% "- I (2) Sample image of crack mapping I

Fig. 2

Transmitted-light image and crack mapping

78

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

Figure 3 shows the relationship between crack density and tensile stress during a loading test. Until strain reaches the point marked b, no cracks were observed. Then, cracks initiated and rapidly grew. The internal cracks

were thus observed clearly and

nondestructively by means of the transmitted-light technique.

500[

/10

400~"

e

8

Stress

b

300

r.,r

200 I-

,~, .~ o

6 l

/

/

c

d

~

0.0 0.2 0.4 0.6 0.8 1.0 Strain e Fig. 3

4

"" -~=~

o

O

7Crack density I

"

(%)

Change of crack density in [002/9004/002]

FBG SENSOR Figure 4 shows the principle of the FBG sensor. A diffraction grating with a specified wavelength is set in the optical fiber. The sensor can measure the change in distance between the gratings. Initiation of cracks can thus be indirectly detected from the change in the grating, which generates a change in wavelength. Figure 5 shows the change of half width of power spectrum, A, during loading test. It was found that the sensor can detect the initial cracks in wavelength range b-c, but the wavelength in range c-d does not change. The FBG sensors can detect the occurrence of some cracks from the change of the spectrum width, but they are not suitable for precisely measuring multiple cracks. If the target is on initial cracks, the FBG sensor is appropriate, because it is not influenced by a high magnetic field. It can be mounted in the FRP structure easily and can be checked from the outside. For example, the FBG sensor is suitable for monitoring the plumbing of a power-generation system, in which initial cracks are serious damage. On the other hand, the load support system of the superconducting magnet is generally designed to be damage-tolerant. Initiation of a lot of cracks, development of a main crack and delamination constitute a threat to the safety of the load support system. Thus there needs a further improvement for the proper application of FBG sensors to the superconducting magnet.

NDE Techniquefor ALFRP dx

79

Core

Incidence light ... ~ ~

:,:90--

A,o: initial wavelength A,i: wavelength under loading ( do: 0.5/.tm)

-"

Reflection light I

fl

TM

10mm

[!

Incidence _ .

A,2 .~

im

Transmitted

----I~ .,..~

o

III

o

;to

22

z~

/13

Wavelength Fig. 4

FBG sensor

1.5

1 c

"~

0

-1

Fig. 5

9

0

9

.

i

,

I

,

2 4 6 8 Crack density D (count/cm)

Relation between crack density and wavelength

ACOUSTIC EMISSION Figure 6 shows the AE measurements of FRP during the loading. Marks b, c, and d in Fig.6 (a) correspond to those in Fig.3. In range b-d, some cracks are observed using the transmitted-light technique, but high AE energy signals cannot be detected in range b-d in Fig.6 (a). Initial cracks have a low AE energy, and so acoustic emission cannot detect the exact position and the number of matrix cracks precisely. Only at the moment of fiber fracture, the AE system can detect the signal well. On the other hand, the transmitted-light technique can detect the internal cracks in detail, can identify at least five cracks per millimeter clearly, and can be used under a high magnetic field.

K. Tanaka, H. Aoyama and N. Takeda

80

c~

b

r~

1000

5000

800

4000

600

Stress

">

3000

AE energy

9

400

2000

200

1000

r~

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

82

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

This Page Intentionally Left Blank

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

85

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

86

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.

An Alternative to Film Based Flash X-Ray

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

90

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

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.

NondestructiveCharacterizationof MaterialsX Green et al. (Eds) Published by ElsevierScienceLtd., 2001

91

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

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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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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

Investigation of Metal Foams Using X-Ray CT

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

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.

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

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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 -

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t~ t--

=-9

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E

o~

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t"q O

o~

'

-

E o

n,

UIO .

.

20

.

.

.,

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~ 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 -

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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.

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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

'

'

I

i

S/N=10.29

0.02

0.02-

'

'

flaw signal

I

S/N=12"20

0) @

O

E @

E ........................

o

!d..no!se

cy

.....

flaw signal

-0.02 I

-0.02

~

,

O o

o'

0 . . . . . . .weld ...............................

-0.02

I

I

0 0.02 In-phase component

-0.02

,

,

I

0 0.02 In-phase component

(a) Differential surface tangential probe (b) Uniform eddy current probe Type 1 Figure 7 Patterns of the flaw signal and the weld noise when probes scan over the flaw parallel to the weld line i

0.02

'

,

i

I

S/N=7.00

0.02 -

_

'

flaw signali

I

'

9

I

S/N = 11.00

flaw signali O

0

E

E

0 o

O

.....

w;idnois

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

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

"x3

o'

-0.02 i

-0.02

I

I

i

id ol,;-

-0.02 [

I

i

I

0 0.02 -0.02 0 0.02 In-phase component In-phase component (b) Uniform eddy current probe Type 2 (a) Differential surface tangential probe Figure 8 Patterns of the flaw signal and the weld noise when probes scan over the flaw perpendicular to the weld line

-

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. '

I

'

I

'

--o-- Differential Pancake Coil

----o- Differential Pancake Coil

z

Z

0

@

0.8

0.8 j=... r

~ 0.6

~- 0.6

~,,.~

~0.4

0.4

g

0.2ff

0.2

06

,

I

,

I

,

,

I

,

I

,

1 2 3 lift-off [mm] litt-off [ram] (a) Flaw parallel to the weld line (b) Flaw perpendicular to the weld line Figure 9 Change of normalized flaw signal amplitude for vs. lift-off

1

2

3

0

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.

Figure 1 A new eddy current probe "~

Detecting coil

!

/ Eddy current ~

Figure 2

~'

Eddy current induced in the material by a circular exciting coil

124

11. Hoshikawa, K. Koyama and H. Karasawa

,

,,W

..-IP

~--

'11--

Detectingcoil

'~ 41--,

,

Detecting coil

..4P' "&

Flaw~..~,o,~lT.

41T. . . . . ~

'~

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,r

Flaw...~:...~.

_, Eddy current ~

q...

4"

1r

Eddy current ~g~ q...

(a) Plus signal Figure 3

4... ~ff~

(b) No signal

r = ......

-,

Eddy curren//t ~K. 41-

Aft'

(c) Minus signal

Eddy current generated by a flaw crossing the exiting coil circle

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

r Scanning area Detecting c&]"................ 7

Exciting coil

Brass plate

Exciting coil ; outer diameter 7 mm, cross section of windings lxlmm 2 Detecting coil ; length 5mm, height 3mm, cross section of windings lxlmm 2 Brass plate ; 160x160xl.5mm 3 Test Frequency ; 32kHz Figure 4

Experimental setup

Test frequency" 32kHz Flaw depth D, length 15mm, width 0.5mm, Lift-off: L I

c~

I

'

'

.5

I

O

lift-off noise

O O

'

4

O O O

O O

\ flaw signal L=0.162mm k , D=80% L - 0.082mm " ~ D - 60% 0

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

~_~_-_-~:0

"0 _

,

,

,

,

,

,

,

,

,

,

,

,

lift-off noise L = 0.082+0.589mm

o

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,

C

D -0.5

~0%

" ~ "D =600/0 D = 80%

i i

CJ

I

-4

,

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-2 In-phase

,

i

,

0 component

I

2

-0.5

0

In-phase

component

0.5

(b) the new probe (a) pancake coil probe Signals by surface flaws of different depths and lift-off noise Figure 5

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

Test frequency" 32kHz, Flaw length 15mm, width 0.5mm _1

'

I i t

'

'

I_

I

'

'

I

'

t

o 0.2

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-1

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.

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.

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.

.

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.

.

.

.

.

.

.

.

.

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-0.2 0 0.2 In-phase component

(a) front surface flaws

(b) back surface flaws

Figure 6

.

Signal flaw patterns for differem flaw depths

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.

Test frequency" 32kHz, Flaw depth 80%, width 0.5mm 1_1

'

I t r

'

I_

t t t

0 O

~

0 flaw length 5mm lOmm 15mm _

-1

-1 Figure 7

0 1 In-phase component Signal pattems by flaws with different lengths

A New Eddy Current Probe without Lift-off Noise

127

Test frequency : 32kHz, Flaw : depth 80%, length 15mm I

'

~

I

1 o //'~

t~

0

o 0 ./" flaw width - - - 0.2mm 0.5mm . . . . . 0.8mm I

-1 Figure 8

i

i

I

0 1 In-phase component

Signal patterns of flaws with different widths

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

O

o

I

'

I

'

60% 100%

--L~.

0

~ -0.5

~~ I ~\~\ \ \ \

UJ',,' ','-.3

-1 I

-1

Figure 9

'

0.5

9"

O'

I

,~. , " , - . / q


o

'

front surface flaw depth 4O%

/\ I 80% 40~ " \ " 60% , , back, ;urface flaw de~th

-0.5 0 0.5 1 In-phase component

Signal pattems by flaws of different depths with normalized amplitude

128

H. Hoshikawa, K. Koyama and H. Karasawa

Test frequency" 32kHz, Flaw" length 15mm, width 0.5mm 100

I

'

I

'

80

I

'

-

'

/ ./ / /

"~ 60

O" / ,/ /

4o

cy Back surface flaw

Front surface flaw

20 I

-150

J

I

~

I

~

I

-100 -50 0 Flaw signal p h a s e , deg

,

50

Figure 10 Flaw depth versus flaw signal phase

CONCLUSION The authors have devised a new eddy current probe that generates minimal lift-off noise. The experimental results have indicated that the new probe provides phase information on the depth of surface flaws. Thus the new probe can make it possible for eddy current testing to evaluate flaw depth more quantitative than the conventional pancake coil probe. Further research has to be conducted to confirm the basic characteristics of the probe under various test conditions.

REFERENCES 1.

E Mcintire, "Nondestructive Testing Handbook," Vol. 4, p58, ASNT (1986)

2.

C.V. Dodd, W. E. Deed, "Analytical Solution to Eddy Current Probe-Coil Problems,"

3.

M. Onoe, "An Analysis of a Finite Solenoid Coil Near a Conductor," (in Japanese)

Journal of Applied Physics, Vol.39, No.6, pp2829-2838 (1968) Journal of IEE of Japan, Vol.88-10, No.961, pp1894-1902 (1968) 4.

H. Hoshikawa, K. Koyama, "A New Eddy Current Probe Using Uniform Rotating Eddy Current," Materials Evaluation, Vol.56, No.l, pp85-89 (1998)

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

129

MICROWAVE MEASUREMENT OF MOISTURE IN ENCAPSULANT RESIN OF IC PACKAGES Y. JU 1, S. SANPEP, M. SAKA 1 and H. ABE 2 1 Department of Mechanical Engineering, Tohoku University, Aoba 01, Aramaki, Aoba-ku, Sendai 980-8579, Japan 2 Tohoku University, 2-1-1 Katahira, Aoba-ku, Sendai 980-8577, Japan

ABSTRACT A microwave method to measure moisture in encapsulant resin of IC packages was studied. A network analyzer was used to generate a continuous microwave signal fed to an open-ended coaxial line sensor and to measure the amplitude of effective reflection coefficient at the sensor aperture. The coaxial line sensor acted both as a source and receiver of microwave signal, which was transmitted into the encapsulant resin and reflected at the surface of a metallic plate placed behind the resin. To evaluate the moisture of encapsulant resin, fundamental experiment was carried out for epoxy resin samples with different moisture contents introduced in various environmental conditions. The relationship between the amplitude of effective reflection coefficient and the moisture content of the encapsulant resin was found, and the evaluation equation was built. The present technique indicates a possibility to determine the moisture in IC packages directly without drying and weighing the package. KEYWORDS Moisture, encapsulant resin, IC package, microwave, coaxial line sensor

INTRODUCTION With the development of IC technology, the size of IC chip has been enlarged to raise the level of integration and, simultaneously, the package has been made thinner and smaller to increase the density of surface mount. Therefore, during the soldering process, as the package is exposed to the solder melting temperature, the stress that arises from the thermal expansion mismatch and the pressure induced by the evaporation of the moisture absorbed in the package sometimes cause it to crack, as shown in Fig. 1. It has been reported that the package cracking initiates from the delamination, which mostly takes place at the interface between the chip pad and the epoxy resin [ 1]. Also, moisture diffusion in IC packages will lead to their failure due to corrosion. Therefore, it is important to understand the contribution of moisture to the delamination in IC packages and the diffusion behavior of moisture in the encapsulant resin. Concerning the studies of moisture contained in the encapsulant resin that affects the reliability of IC packages, the basic issue is the measurement of the moisture content in the encapsulant

130

Y. Ju et al.

resin. In the present paper, a new microwave method to measure the moisture absorbed in the encapsulant resin is studied. Most standard methods of determining moisture content require weighing a sample, in some cases drying it for several days (up to several weeks), and reweighing. The primary advantage of microwave method over the standard weighing method is that the moisture content can be determined directly from a wet material without drying and weighing it [2]. Fig. 1 Actual delamination and crack observed Microwave has been used to determine the by microscope moisture content of grains [3]. However, conventional microwave moisture measurement cannot be used for IC packages due to lower spatial resolution. Recently, the authors have developed a microwave imaging technique by using an open-ended coaxial line sensor to detect the delamination in IC packages [4]. The sensor same as the above is used in the present experiment. It can effectively transmit and receive microwave with a relatively high spatial resolution.

BASIC DEFINITION AND PRINCIPLE The moisture content of the encapsulant resin, M, expressed in percentage, is defined as M =

mw x 100 m w + md

(1)

that is, as the ratio of the weight of the contained water per unit volume, m w, to the total weight of wet material, where mj is the weight of dry material per unit volume. The principle of the technique described here is based on the interaction of microwave signal with the encapsulant resin. An open-ended coaxial line sensor is used as a source and receiver of the microwave signal that is transmitted into the resin, and reflected at the surface of the metallic placed behind the resin. The amplitude of effective reflection coefficient in decibel, A, is associated with the permittivity, the thickness of the resin and the operating frequency. The permittivity of wet resin can be considered as a resultant of those of water and dry resin. Since the permittivity of water is significantly larger than that of dry resin, the change in moisture absorbed in the resin will lead to the change in permittivity of wet resin. Therefore, when the thickness of the resin and the operating frequency are fixed, the amplitude of effective reflection coefficient is only a function of the moisture of the encapsulant resin.

EXPERIMENTAL PROCEDURE The configuration of the microwave measurement system is shown in Fig. 2. A network analyzer (HP8510) was used to generate a continuous wave signal fed to the coaxial line sensor and to measure the amplitude of effective reflection coefficient at the sensor aperture. As

131

Moisture Measurement by Microwaves

Network Analyzer

~

2c

I<

[ Computer

.......

~,

I Sensorfixture

Z Stage

I ST~;aWle I~,/ Metallic . ~ .... ~ ~//plate

VIIIIA

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

7////AT/A7/III

X-Y Stage

l~lange Fig. 2 Configuration of the microwave measurement system

tra tluoro ethylene

Center conduetor

Fig. 3 Configuration of the open-ended coaxial line sensor

shown in Fig. 3, the sensor has an inner and outer radii, a = 0 . 4 6 and b = l . 5 0 mm, respectively. It was terminated into a fiat metallic flange with radius c= 14.50 mm. A computer was employed to record the data output from the network analyzer and synchronize the stage translation in the x-, y- and z-directions. Encapsulant resin samples were formed by filling silica powder in 83 wt%. The dimensions of the sample are shown in Fig. 4. The moisture was introduced with the environmental conditions of 358K, 85RH% for 48 to 168 hours. The moisture contents were measured using the standard weighing method in advance. Dry sample was treated in 375K for 24 126.55 hours. For eliminating the thermal L .I influence induced by the process of moisture absorption, microwave iiii~!~iiiiiiiiiiiiiiiiiiiiiii!i~!i~i~i!!i~i!i!i~i~!!iii~ii~ii!iii~ii~iii'~iiiiii~iiiiii!/!iiii~iiii!!'ii!ili 12.90 measurement was carried out after the sample had been exposed to the ambient for 90 minutes. In the experiment, the coaxial line sensor was contacted with the surface of the sample. To increase measurement sensitivity, the frequency was swept from 45 MHz to 20 GHz for finding the resonance frequency.

i!:!:!i~i~:i:/:~i~i~i~i~!~i~!~'ii~:~i~i~i:::::::::::::::::::::::::::::ii~!::~i~i~i!~i~i~:~!::! Dimensions in mm

|........................,..,.,.,.,

~7~,.~...............................................

........

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

~ 1.00 u..~@,.

|

t Fig. 4 Descriptive geometry of the encapsulant resin sample

RESULTS Figure 5 shows the amplitude difference of the effective reflection coefficient, measured between the wet and dry samples, versus the frequency around the resonance frequency. It is shown that at the frequency of 12.5 GHz, larger changes in amplitude were obtained for different moisture contents. In other words, higher sensitivity can be obtained. Figure 6 shows the relationship between the amplitude of effective reflection coefficient and the moisture content measured for different wet samples at the frequency of 12.5 GHz. The amplitude of effective reflection coefficient, A, is found to be linear with the moisture content. The evaluation equation is determined as M = -0.2506A- 1.229

(2)

132

Y. Ju et al.

1.0

-4.8 .................................................. -5.0

0.5

--..!1---

0.17%

cD

=

_.-

0.22%

o.o

~ -5.4

-.~.

-0.5 :

<E

i ,

-1.0

i _ _ ~

GHz ]

m -5.2

i

~

~

j

i

12.00 12.25 12.50 12.75 13.00 13.25 Frequency (GHz) Fig. 5 Amplitude difference between the wet and dry samples versus frequency

< -5.6 -5.8 -6.0 ' ' 0.00 0.05 0.10 0.15 0.20 0.25 Moisture content (%) Fig. 6 Relationship between the amplitude of reflection coefficient and the moisture

By using Eq. (2), once the amplitude of effective reflection coefficient is measured, the moisture contained in the encapsulant resin can be evaluated. It is noted that, for other encapsulant resins, once an evaluation equation is made as demonstrated in the present paper, it is also possible to determine the moisture directly by using microwaves.

CONCLUSIONS The method to determine the moisture in encapsulant resin by using microwaves was demonstrated. The relationship between the amplitude of effective reflection coefficient and the content of moisture contained in the encapsulant resin was found to be linear. By using the established evaluation equation, it is possible to determine the moisture content directly without drying and weighing the sample. Therefore, the technique may be used in an on-line environment.

ACKNOWLEDGEMENTS The authors would like to thank Mr. K. Oota of Sumitomo Bakelite Co., Ltd. for preparing the samples. This work was partly supported by: The Ministry of Education, Science, Sports and Culture under Grant-in-Aid for Encouragement of Young Scientists (A) 12750065 and Scientific Research (B)(2) 11555025; COE Research 11CE2003.

REFERENCES 1. 2. 3. 4.

Lee, H. and Earmme, Y. Y. (1996) IEEE Transactions on Components, Packaging, and Manufacturing Technology-Part A 19, 168. Ju, Y., Saka, M. and Ab6, H. (1999) Proc. the 1st International Workshop on Electronics Materials and Packaging, 182, Singapore. King, R. J., King, K. V. and Woo, K. (1992) IEEE Transactions on Instrumentation and Measurement 41, 111. Ju Y., Saka, M. and Ab6, H. (1999) Advances in Electronic Packaging 1999 26-1, 847, ASME, Hawaii.

ULTRASONICS

This Page Intentionally Left Blank

Nondestructive Characterization of Materials X Green et al. (Eds)

135

9 2001 Elsevier Science Ltd. All rights reserved.

PRECISE VELOCITY MEASUREMENT OF HIGH FREQUENCY SURFACE ACOUSTIC WAVES BY PHASE VELOCITY SCANNING METHOD Hideo CHO and Kazushi YAMANAKA Department of Materials Processing, Graduate School of Engineering, Tohoku University. Aoba-yama 02, Sendai 980-8579, JAPAN. E-mail: [email protected]

ABSTRACT We achieved precise anisotropy measurements of surface acoustic waves ( SAW ) by the phase velocity scanning ( PVS ) method on single crystal silicon and synthesized 33 degree st-cut quartz with aluminum thin film, a material frequently used for electrodes or reflection gratings. We succeeded to generate and detect high frequency SAW up to 320 MHz by improving the scanning interference fringes approach of the PVS method. We measured SAW velocity anisotropy of single crystal silicon with a relative error within 0.05 %. We found that the SAW velocity of quartz with 100 nm A1 film was lower than that with 20 nm A1 film in propagating around 90 degree from x-axis. The maximum difference was around 4 m/s at 90 degree from x-axis and the measured velocity agreed well with calculated one. Velocity difference due to A1 film thickness difference in SAW propagating around x-axis was smaller than that in SAW propagating around 90 degree from x-axis. KEYWORDS

PVS method, precise measurement, surface acoustic waves, velocity anisotropy, quartz

INTRODUCTION Surface acoustic waves (SAW) devices have been widely used in various field. In particular, quartz has been most popular as SAW device material before because it's physical, chemical and frequency stability is good. Moreover, quartz SAW resonator devices with Q-value of over 20000 was developed and fabricated and these have important roles as frequency control devices in a range of VHF to UHF band. However, SAW devices always need electrodes and/or reflection gratings for generating and detecting SAW waves. These electrodes and gratings were thin films whose thickness are sub-micro order and deposited onto a surface of base materials. In the design of SAW devices, it is essential to measure and control the film thickness because the velocity of SAW changes significantly depending on these thickness. Recently, the rapid growth of mobile telecommunication technology has led a range of frequency-band from MHz to GHz and excitation power from mW order up to W order[l]. Therefore, the large stress induced by SAW caused stress-migration of electrodes and/or gratings. Necessity for estimation of electrodes and/or gratings thickness is more increasing. To characterize thin films, it is essential to generate and detect high frequency SAW and to measure the SAW velocity with high precision. Scanning acoustic microscope (SAM) has been employed for precise velocity measurement of high fi-equency SAW. SAM always requires liquid cuplant to transmit and receive the SAW. The metal electrodes may be damaged by liquid cuplant. On the other hand, laser ultrasonics is a potentially useful tool for surface characterization without liquid cuplant. In the conventional laser ultrasonics, the laser with short

136

H. Cho a n d K. Y a m a n a k a

pulse width allows to generate high frequency SAW. However, the high peak power may lead to ablation damages on the surface. There are a few reports for high frequency SAW generation without ablation damages. For example, Harata et. al.[2] suggested the laserstimulated scattering microscopy. In this method, stationary interference fringes on the surface produced by two laser beams with the same frequency and pulse width of 84 ps can generate the narrow-band high frequency SAW. Using this method, they obtained the images of distribution of SAW velocity and thermal diffusivity on Ar-implanted single crystal silicon. Another method was reported by Rogers[3]. Simple pattern mask to diffract a laser beam, named 'phase mask ', was specially designed to produce a variable separation of light fringes instead of utilizing the interference. This method can generate narrow-band 500 MHz SAW and then measure the velocity dispersion of Pt-coated glass fiber[4]. Benet et. a/.[5] succeeded to estimate the Ta and Cu films thickness of 20 - 200 nm on single crystal silicon using the method proposed by Rogers. In this method, the 'phase mask' has to be replaced for changing the SAW frequencyl The light fringes produced by these two method are stationary and always generate not only SAW but also bulk waves into the specimen. The bulk waves may have significant influence on precision of SAW velocity measurement with thin substrates. We have proposed and developed a novel laser ultrasonic method, phase velocity scanning (PVS) method, utilizing interference of two laser beams with different frequency[6]. The PVS method enables us to obtain the SAW velocity dispersion in the frequency range of 20 to 100 MHz within 1% error and estimate both the elastic properties and thickness of various thin films simultaneously by an inverse analysis[7]. In this study, we improved the apparatus to generate and detect higher frequency SAW up to 320 MHz. Also, we measured the velocity anisotropy of SAW on single crystal silicon and synthesized quartz with aluminum thin film, a material frequently used for electrodes or reflection gratings. PRINCIPLE AND EXPERIMENTAL SETUP FOR SAW GENERATION

Figure 1 shows principle of scanning interference fringes (SIF) approach of PVS method. Two laser beams with different frequencies irradiate the sample surface with specific laser incident angles (01,02) and form interference fringes scanned at phase velocity of SAW in one direction. Therefore, this method has some advantages, such as 1) enhanced amplitude of SAW, 2) high selectivity of SAW and bulk waves, and 3) excellent directivity of the generated ultrasonic waves. Since the wavenumber of SAW is equal to that of SIF, Fig.1 Principle of scanning interference generated SAW velocity V s A w is expressed by fringes approach of the phase velocity scanning method. VSAW = ~lk s (1) where, co is angular frequency of generated SAW. Wavenumber of SIF kf is determined from the wavelength and incident angles of the laser beams as kf -- /~1 sin 01 ~- k 2 Sin 02

(2)

Precise High Frequency SA W Measurement

137

However, inherent wavenumber of second harmonic YAG laser beam without frequency shift is much larger than the amount of change of wavenumber due to frequency shift. For example, if the amount of frequency shift is 320 MHz, change in wavenumber due to YAG laser frequency shift is as small as 6.0 x 10-r. Therefore, we assumed that kl almost equals k2 and rewrite eq(2) as k / = K(sin 01 + sin 02) (3) where, K ( = k 1 ~ k2) is wave number of laser beams. Equations (1) and (3) imply that frequency of generated SAW depends on the laser incident angles and measurement accuracy of SAW velocity is governed by the accuracy to set the laser incident angles. The modified system schematically shown in Fig. 2 is almost the same as reDigitizing ported before[7]. However there are two Braggcell higher frequency SAW. First, for generak~_ ~ [Nd:YAG I tion, we installed three Bragg cells on the w~ ~ mTrigger paths of two YAG laser beams to provide ~ I , ~ Ap~ a large frequency difference. Two Bragg leO2 l cells are fixed frequency shifters of 110 MHz BraggC ~ ( ~ r 0 ~ ~ and another is a variable frequency shifter (6~110MHz) / ~ in the frequency range of 60 - 110 MHz. ~~ _ _ - - - J ~ Secondly, for detecting by the optical knife v

edge method, we inserted a beam expander of the Ar laser beam to obtain a laser beam spot smaller than the wavelength of SAW, around 10 pm. The diameter of YAG laser beam spot was about 4mm and separation from generating to detecting point of SAW was 10- 20 mm.

I n+~l lArgo

I 11

APDI i.,~Knifeedge i~ ..~);.. /~j /~

Specimen

~ ~ ] Beamexpansionandcollimationsystem Fig.2 Experimental setup.

EXPERIMENTALRESULTS Measurements on velocity anisotropy of single crystal silicon Figure 3 shows a typical SAW waveform and its power spectrum on Si(001) surface along [110] direction. We observed an ideal sinusoidal signal after 20 times averaging. The peak of its power spectrum indicates the frequency of 320MHz. We carried out the measurement of velocity anisotropy on silicon (001) surface by in-plane rotation of silicon wafer using the computer controlled stage and plotted in Fig.4. The rotation step was 0.5 degree and the laser incident angles were kept constant. The symbol O indicates the measured velocity at around 320MHz and solid lines indicate calculated velocity with the properties listed in table 1. In this experiment, since wavenumber of generated SAW ,which equals that of SIF, is constant, the frequency of detected SAW depends on the SAW phase velocity of sample. Therefore, we can obtain the SAW phase velocity of sample by eq.(1). However, it is difficult to measure the incident angles of laser beams directly because these angles are very small. Therefore, we used the calculated velocity ( 5083 m/s) along [110] direction to calibrate the measured velocity.

H. Cho and K. Yamanaka

138

!

~176176 V

~o ---0.00

O

'1"

I

'1

I

I

I

I

!

'l

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O MeasuredData (320 MHz) ..... Calculated Data

510s

0.05 Time,

t'

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E

490 200

i

l

M H z

,

,

25

I

i

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50

.|,

75

Rotated angle, deg

400

300

Frequency,

i

Fig.4 Measured and calculated velocity anisotropy of single crystal Si.

Fig.3 Detected SAW waveform and it's power spectrum.

Measured velocity shows the periodic anisotropy of 90 degree and agreed sharply with the calculated one along all propagation direction. The relative error was within 0.05 % and twice higher accuracy was obtained than that in the measurement using around 110 MHz[8]. We suppose that this accuracy was caused by higher directivity of SAW at the higher frequency.

Table 1 Elastic properties and densities of silicon, quartz and aluminum used for velocity anisotropy calculation.

Silicon

C13 12 (GPa) ........{Gea) 63.9

i,(Gpa) 116

C14 C33 C44 C66 Density (Gea) (Gpa) (G.Pa)..... (GPa) (kg/m3) 79.6 233O ,~

Quartz Sato et. al. .

.

.

.

.

.

.

85.6

Bechmann

86.74

Aluminum

111.3

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5.49

5.93

98.3

58.2

39.85

2650

6.99

11.91 -17.91 107.2

57.94

39.88

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59.1

.

.

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.

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.

.

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26.1

2699

Measurements on velocity anisotropy of single crystal quartz with aluminum films We prepared two 33 degree ST-cut quartz 3 inch wafers onto which an aluminum film of 20 and 100 nm thickness was deposited by sputtering. We measured velocity anisotropy of each sample using around 210 MHz SAW in the propagation direction range of-20 to 40 degree and 65 to 110 degree from x-axis. Measured SAW frequency in the propagation range of 65 to 110 degree from x-axis is plotted after moving average in Fig.5. The measured anisotropy curves are almost symmetrical with respect to the maximum value at 90 degree from x-axis. The SAW frequency of quartz with 100 nm A1 film was lower than that with 20 nm A1 film along all direction. To estimate the SAW velocity from measured SAW frequency, we need a reference SAW velocity along a specific propagation direction. We calculated the reference velocity using

139

Precise High Frequency SA W Measurement

elastic constants and density of quartz reported by Bechmann et.al.[9] and Sato et.al.[lO] listed in table 1. Elastic constants and density of A1 film used for the calculation are also listed in table 1. The reported values by Sato et.al have been estimated from SAW velocity anisotropy of the same sample using frequency of 80 MHz by an inverse analysis. On the other hand, Bechmann et. al. determined the elastic constants by measuring resonance frequencies of quartz plates. Figure 6 shows calibrated SAW velocities of the quartz with a 20 nm thickness A1 film. We used velocities calculated by Bechmann's constants ( 3515.6212 m/s ) and Sato's constants ( 3511.0774 m/s ) along 90 degree from x-axis to calibrate the measured data. Both calibrated velocities agreed with the calculated ones around 90 degree. However, agreement between calibrated and calculated velocity by Bechmann's constants became worse than that by Sato's constants as separating away from 90 degree. Therefore, we adopted Sato's as a reference.

216

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.

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i

,

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Fig.6 Comparison with the reported constants of Bechmann and Sato.

Figure 7 shows measured and calculated SAW velocity anisotropy curves. Symbols O, 9 represent the A1 film thickness of 20nm and 100nm respectively and dotted line indicate calculated velocity anisotropy without A1 film. SAW velocity of quartz with 100 nm A1 film was calibrated by the same calibration constant in the case of 20 nm A1 film quartz specimen. Measured velocities agreed sharply with calculated ones within 0.07 % error. We clearly observed that the SAW velocity of quartz with 100 nm A1 film was lower than that with 20 nm A1 film along all direction. The maximum difference was around 4 m/s at 90 degree from x-axis. The difference was caused by the lower SAW velocity of A1 ( 2906 m/s) than that of quartz. We also recognized that calculated velocity of quartz without A1 film was different from that with 20 nm A1 film. Figure 8 shows measured and calculated SAW velocity anisotropy of quartz with 20nm and 100nm A1 films propagating around x-axis. In this case, the calculated velocity along x-axis was used to calibrate the measured one. Two measured velocities showed almost the same value in all propagation direction and agreed with calculated ones within 0.18

H. Cho and K. Yamanaka

140

% error. The minimum calculated velocity difference between 20nm and 100 nm A1 film sample is 1 m/s along x-axis. We thus found that SAW velocity is less influenced by the thickness of A1 film in propagating direction around x-axis than around 90 degree from x-axis. ,

i

,

,

,

,

]

.-

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E

-~ 350( o

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0

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<

k

"~

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~-100ore 9 20 nm o 0 nm 20

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(Quartz)

80

1 O0

Propagation angle, degree Fig.7 Measured and calculated SAW velocity anisotropy of single crystal quartz with aluminum films propagating around 90 deg from x-axis.

-20

.

. 13 .

.

--

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

Propagation angle, degree Fig.8 Measured and calculated SAW velocity anisotropy of single crystal quartz with aluminum films propagating around x-axis.

CONCLUSION

We improved our system to generate and detect higher frequency SAW and measured the SAW velocity anisotropy of single crystal silicon and quartz with A1 films. The results obtained are summarized as follows" 1)

We succeeded to generate and detect higher frequency SAW up to 320 MH using three frequency shifters of laser beams.

2)

We measured SAW velocity anisotropy of single crystal silicon and achieved a relative error within 0.05 % , twice higher accuracy than that in the measurement using SAW around 110 MHz.

3)

In quartz specimens, agreement between measured and calculated velocities is also good and a relative error is less than 0.18 %. Velocity difference due to A1 film thickness difference in SAW propagating around x-axis was smaller than that in SAW propagating around 90 degree from x-axis.

4)

Our improved system enables us to estimate the quality of devices by precise velocity measurement of high frequency SAW. We can evaluate the influence of the film thickness on the SAW velocity depending on the propagation direction.

ACKNOWLEDGMENT

This work was supported by the Proposal-Based New Industry Creative Type Technology R&D Promotion Program from the New Energy and Industrial Technology Development Organization (NEDO) Japan.

Precise High Frequency SA W Measurement

141

References [1] N. Matsukura, T. Morisaki, E. Ootsuka, Y. Yamamoto and A. Kamijo, Surface acoustic wave filter with highly textured A1 film, 'Recent Research for Elastic waves devices' edited by the 150th committee on elastic wave technology in Japan society for the promotion of Science (2000) pp.413-417 ( in Japanese ) [2] A. Harata and T. Sawada, Evaluation and Imaging of Materials Using Picosecond Laser-Induced Ultrasonics, Jpn. J. Appl. Phys. Vol.32, No. 5B (1993) pp. 2188-2191 [3] J A.Rogers, Complex acoustic waveforms excited with multiple picosecond transient gratings formed using specially designed phase-only beam-shaping optics, J. Acoust. Soc. Am, Vol.104, No.5 (1998) pp. 2807-2813 [4] J A. Rogers, Coherent wavelength-tunable ultrasonic modes stimulated in optical fiber, micorcapillaries, and planar microfluidic networks using crossed-picosecond laser pulses, J. Appl. Phys., Vol.86, No.6 (1999) pp. 2959-2966 [5] M. J. Banet, M.Fuchs, J A. Rogers, J H. Reinold. Jr, J M. Knecht, M. Rothschild, R. Logan, A. A. Maznev and K A. Nelson, High-precision film thickness determination using a laser-based ultrasonic technique, Appl. Phys. Lett., Vol.73, No.2 (1998) pp.169-171 [6] K. Yamanaka, O. Kolosov, H. Nishino, Y. Tsukahara, Y. Nagata and T. Koda, Analysis of excitation and coherent amplification of surface acoustic waves by the phaser velocity scanning method, J. Appl. Phys., Vol.74 (1993) pp.6511-6522 [7] H. Cho, H. Sato, M. Takemoto, A. Sato, K. Yamanaka, Surface acoustic wave velocity and attenuation dispersion measurement by phase velocity scanning of laser interference fringes, Jpn. J. Appl. Phys., Vol.35 (1996) pp.3062-3065 [8] H. Nishino, Ph.D. Thesis, Tohoku University (1995) p.121 [9] R. Bechmann, A. D. Ballato and T. J. Lukaszek , Higher-order temperature coefficients of the elastic stiffness and compliances of Alpha-Quartz, Proc. IRE Aug. (1962) pp.1812-1822 [10] H. Sato, H. Nishino, H. Cho, H. Ogiso and K. Yamanaka, Estimation of elastic constants from surface acoustic wave velocity by inverse analysis using the downhill simplex method, Jpn. J. Appl. Phys., Vol.37 (1998) pp.3116-3119

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

143

N D E OF P L A T E S A N D C O A T I N G S U S I N G CONTACT-TYPE LINE-FOCUS ULTRASONIC PROBE

HIROTSUGU INOUE, KIKUO KISHIMOTO and TOSHIKAZU SHIBUYA

Department of Mechanical and Control Engineering, Tokyo Institute of Technology 2-12-10-okayama, Meguro, Tokyo 152-8552, Japan

ABSTRACT Lamb waves and Rayleigh-like waves are often used for ultrasonic NDE of plates or coatings. The dispersive nature of these waves causes difficulties in excitation and detection of ultrasonic waves as well as in the analysis of detected signals. In this study, a contact-type line-focused transducer made of a quarter-cylindrical PMMA wedge with a P V D F film adhered to the wedge is developed. It is shown that dispersive waves in plates or coatings are successfully excited and detected by this transducer. Two techniques utilizing the wavelet transform are employed to determine the dispersion relation of Lamb or Rayleigh-like waves from detected signals. In addition, the elastic properties of plates or coatings are estimated from the dispersion relation by applying an optimization technique. KEYWORDS Ultrasonics, Lamb waves, Rayleigh-like waves, Plates, Coatings, Wave dispersion, Wavelet transform, Optimization.

INTRODUCTION Lamb waves and Rayleigh-like waves are often used for ultrasonic NDE of plates or coatings. One of the difficulties in using these waves arises from the dispersive nature of these waves. If a conventional angle beam transducer having a specific angle of incidence is used, ultrasonic waves can be excited and detected only at some specific frequencies determined by Snell's law. Various techniques have been developed for exciting and detecting Lamb waves in plates or Rayleigh-like waves in coatings. One of the most popular techniques is the use of immersion-type focused transducers including acoustic lenses of the acoustic microscope. Electro-magnetic acoustic transducers are also known well as a non-contact technique. Excitation with a pulsed laser beam and detection by a laser interferometer are being studied actively in recent years (e.g. Achenbach [1]). Other interesting techniques include

144

H. Inoue, K. Kishimoto and T. Shibuya

interdigital transducers [2], electrostatic transducers [3], stacked wedge transducers [4], and phase velocity scanning of single laser beam [5] or laser interference fringes [6]. While the above techniques can be successfully used to excite and detect Lamb or Rayleighlike waves, they require more or less special equipment in practical applications. It is desired to use a conventional ultrasonic pulser/receiver for exciting and detecting Lamb or Rayleigh-like waves. In this study, a new contact-type line-focused transducer is developed for exciting and detecting Lamb or Rayleigh-like waves by using a conventional ultrasonic pulser/receiver. The developed transducer is applied for exciting and detecting Lamb waves in aluminum plates and Rayleigh-like waves in titanium coatings on steel substrate. The dispersion relation of ultrasonic waves is determined from detected signals by two techniques utilizing the wavelet transform. In addition, the elastic constants of plates and coatings are estimated from the dispersion relation by an optimization technique.

CONTACT-TYPE LINE-FOCUSED TRANSDUCER The conventional angle beam transducer having a specific angle of incidence can excite ultrasonic waves only at some specific frequencies determined by Snell's law. If the angle of incidence covers from 0 to 90 degrees, all possible waves would be excited simultaneously. The detection of ultrasonic waves also becomes possible by the same principle. Based on this simple idea, a new contact-type line-focused transducer was developed as shown in Fig. 1. The transducer is made of a quarter-cylindrical PMMA wedge with a PVDF film and electrodes adhered to its cylindrical surface. As an electric pulse is supplied to the electrodes, a cylindrical wave is emitted from the PVDF film and focuses into a line at the bottom of the wedge in contact with the specimen. Although the angle of incidence of this transducer covers from 0 to 90 degrees, the range of the wave velocity (i.e. the frequency band of the wave) to be excited and detected by this transducer would be limited to some extent mainly due to the critical angle of incidence from the PMMA wedge to the specimen. Nevertheless, this transducer is expected to show performance at least better than conventional angle beam transducers.

Fig. 1. Contact-type line-focused transducer developed in this study.

NDE Using Contact-Type Line-Focus Ultrasonic Probe

145

Fig. 2. Experimental setup for ultrasonic testing of plates and coatings. Table 1. Mechanical properties of materials tested. Young's modulus [GPa] 66.1 ~ 71.2 ~ 112 b 211 d

Poisson's ratio

A1 plate (1-mm thick) 0.31 ~ A1 plate (0.3-mm thick) 0.32 ~ 0.32 c Ti coating Steel substrate 0.28 d Quasi-static tensile test b Quasi-static bending test c Typical value d Ultrasonic test using bulk longitudinal and shear waves

Mass density [kg/m a] 2662 2692 4391 7808

ULTRASONIC TESTING OF PLATES AND COATINGS Ultrasonic tests of aluminum plates (1-mm and 0.3-mm thick) and titanium coatings (3mm, 1-mm and 0.5-mm thick) on steel substrate were conducted as shown in Fig. 2. A pair of transducers were placed at a distance d facing each other. A conventional pulser/receiver (Krautkrs USIP 12) was used to drive the transducers. Received signals were recorded by a digital oscilloscope (Hewlett Packard, 54522C) at a sampling interval of 10 ns. The aluminum plates tested are commercially available ones. On the other titanium coating on steel substrate was manufactured by explosive welding. thickness of the coating was originally 6 ram, it was reduced to specified values the surface. Mechanical properties of tested materials are summarized in Table

hand, the Since the by milling 1.

Figure 3(a) shows received signals for 0.3-ram thick aluminum plate for various values of d. The time origin is arbitrary but conmaon for all distances. The received pulse becomes broader in time as the distance increases, which is a typical characteristic of dispersive waves. It can be said that Lamb waves were successfully excited and detected by the

H. Inoue, K. Kishimoto and T. Shibuya

146

1

I

'

I

'

'

1

I

'

I

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0

-1 2 1 0 -1

~,2

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d = 30 mm

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0 -1 2 ]

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(a) Lamb waves for 0.3-mm thick aluminum plate

30

0

,

I

10

, Time [l~s]

I

20

,

SO

(b) Rayleigh-like waves for 1-mm thick titanium coating on steel substrate

Fig. 3. Signals received by the developed transducer. developed transducer. The dominant component detected in this case is the a0 mode while other modes are also detected simultaneously. The characteristics of the developed transducer is not yet investigated very well. It might be possible to excite and detect a specific mode more efficiently if the design of the transducer is improved. Received signals for 1-mm thick titanium coating on steel substrate are shown in Fig. 3(b). Although these signals are relatively noisier than those in the previous case, the developed transducer is also applicable to excite and detect Rayleigh-like waves in coatings.

DETERMINATION OF DISPERSION RELATION BY WAVELET ANALYSIS The frequency dependence of the wave velocity, i.e. the dispersion relation is often a key for evaluating elastic constants, thickness, or subsurface defects in ultrasonic NDE of plates or coatings. If an ultrasonic pulse is detected at two locations along the direction of wave propagation, the wave velocity can be determined from the difference between arrival times at these locations. For dispersive waves, however, it is difficult to determine the arrival times directly from the detected signals. Two techniques utilizing the wavelet transform were employed in this study.

NDE Using Contact-Type Line-Focus Ultrasonic Probe The wavelet transform of a time signal

147

f(t) is defined by

l/Yf(a,b)=--~

~f(t)r

T

dt,

(1)

where a > 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

[ (Wo/7)2t2]exp(iwot) -

2

(2)

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

148 ~3200~

....

j

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(.9 290~';. ....

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Frequency [MHz]

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3200

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._~ 3000

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o 24001-, 0.5

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Frequency [MHz]

10

(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

n = ~[~,(~.; E..) - q(~)]~.

(3)

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

NDE Using Contact-Type Line-Focus Ultrasonic Probe 3200 --,

. . . . .

3200 . . . . .

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~ ~o~i 5 . . . .

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~ 2~% ....

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|

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~' 3200 L 31 00]--~ o 3000 ~

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,

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/

Frequency [MHz]

'

A

.=. 31 oo L-

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,

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lO

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.--.3200L 3100 ~

"-" ]_ ~ o 3000 ~ - - ~ ~ @~

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'

,

,

j

~'32~

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,

= 10

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9

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,

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,

I 5

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....

3100 ~ ]-

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~

~ 3000 ~ - - - - ~ " X

ooop

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,

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,

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(c) 0.5-ram thick titanium coating (d, = 15 mm, d2 = 40 ram) Fig. 5. Dispersion relations of the Rayleigh-like wave in titanium coatings on steel substrate obtained by Method A (left) and Method B (right). was accomplished by simply applying a commercial software for nonlinear least squares method (Optimization Toolbox of Matlab, the MathWorks). No constraint was considered because no effective constraint was available at the cost of increasing computational task. The results of estimation are listed in Table 2 in which relative errors with respect to the values in Table 1 are also indicated. In addition, the dispersion relations derived from these estimates are indicated in Figs. 4 and 5 with broken curves. Young's modulus is accurately estimated in all cases except for the case of 0.5-mm thick titanium coating. As mentioned above, this is mainly due to the nonuniformity of the coating thickness. On the other hand, Poisson's ratio is estimated less accurately in almost all cases. This is not due to the optimization procedure because the dispersion relations derived from these estimates fit the measured results well as shown in Figs. 4 and 5. Since Poisson's ratio is very sensitive to the group velocity, further improvement of the experimental procedure will be necessary.

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

152

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

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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

.~ <

4.0 "~

\

i

50o

6oo

Fig. 6. Waveforms after full-wave rectification and after SSP in the case of the artificial defect, at the position 40 mm from the strip edge.

Y. Nagata et al.

156

SNR calculations based on Fig. 6 showed that the SNR = 3.24 of the raw waveforms increased to SNR = 7.07 after SSP. Fig. 7 shows the SNRs of raw signals and processed signals in the case of the artificial defects 50, 40 and 30 mm from the strip edge. The improvement of SNR by using SSP was clearly observed.

SN 25 20 15 10 5 0

Raw ~ignal --ft..Processed signal 50

40

30

Distance from the edge (mm) Fig. 7. SNR values of raw signals and processed signals in the case of artificial defects at the positions 50, 40 and 30 mm from the strip edge.

DEVELOPED ON-LINE SYSTEM As the effectiveness of SSP was confirmed by the results of SSP application experiments described so far in the preceding sections, the realization of an SSP system which can be used on-line was investigated. As obvious from Fig. 1 and Eq. (5), SSP is a combination of linear filtering processing and nonlinear filtering processing. Therefore, it has been difficult to realize real-time SSP, and no report has ever been made on SSP incorporated in an on-line inspection system. There is a report on simple linear filtering performed by software for a Lamb wave inspection system [16], and another report on SSP calculations by hardware [15]. In our research the realization of real-time SSP by software was investigated to enhance the flexibility of the setting of the SSP parameters and the realization of an on-line commercial-level system. In order to realize real-time SSP, the necessary number of SSP calculations and the data processing ability required for the computer should be considered. If the number of filter coefficients used for filtering in SSP is m, m times of multiplication and (m-l) times of addition totaling (2m-l) times of calculation are necessary for calculating each filtering processing shown by Eq. (5). Therefore, if the number of data sampled is N, the total number of calculations A is expressed as follows: A = nN(2m-1).

(6)

If the strip width is W, and if the phase velocity of the Lamb wave mode used during inspection is V, the time T required for an ultrasonic reflected echo to reach the probe from the strip edge is: T=2W/V. Therefore, in the case of 20 MHz sampling, the number of data N necessary for sampling is expressed as follows: N = 20 x 106 x 2W/V.

(7)

From Eqs. (6)-(7), A = 40 X 106 • n(2m-1) W/V.

(8)

If, for example, W=l.5 m, V=3000.0 m/s, m=500 and n=10, A -- 2.0• Therefore, if the repetitive frequency of inspection is 500 Hz, the necessary calculation speed B = 500A - 99.9 (GOPS). In this case, the calculations of the nonlinear processing in SSP shown by Eq. (3) are

Real-Time SSP

157

not included. Data preprocessing before SSP calculations, as shown in Fig. 8, is highly effective for drastically reducing such an enormous number of calculations. As previous description, the Lamb wave ultrasonic inspection system detects defects using narrow-band toneburst waves (at a frequency of about 2.25 MHz in the case of this system). Therefore, SSP calculations are performed after passing A/D-converted data through band-pass filters which include the bands of toneburst signals, processing the data to select and leave one out of 10 data and reducing the number of data N and the number of filter coefficients m to 1/10 as shown in Fig. 8. According to the sampling theorem, even after such data processing, the data reduced to 1/10 still preserve the original information contained in the bands of the burst wave signals. In other words, if k is an integer, the frequency is f, the sampling frequency is fs and the signal spectrum is X(f), there is the property that X(f)=X(f+f~k) in the case where there is no aliasing [17]. This corresponds to the case where k = 2, ~ = 1 MHz and 0.0
A'= N(2m-1)+nN(m/5-1)/lO

(9)

Regarding the data processing to reduce the number of samples to 1/10, the calculations are not included in the number of the above consideration because the processing time in the computer can be neglected. For the further improvement of the processing ability, it is also very effective to perform SSP calculations, image processing, etc., all in parallel by using a high-speed personal computer having two latest-type CPUs. In this way the real-time SSP was realized at a repetitive frequency of 500 Hz by the drastic reduction of the number of calculations and parallel data processing. Fig. 9 shows a block diagram of the developed system. This on-line system makes 12-bit A/D conversion of ultrasonic signals from the existing Lamb wave inspection system at a maximum sampling frequency of 20 MHz, transfers the digital data to the personal computer and performs realtime SSP calculations (all calculation in the computer: 16-bit) and two-dimensional display of the results. To facilitate the setting and change of such SSP parameters as the number of filters n, each filter band b, center frequency 1"1 and center frequency interval DI, the system is provided with a screen for setting the SSP parameters. The system also can make real-time display of two-dimensional waveform amplitude values by following up the passing of the strip being inspected. Furthermore, the system is capable of real-time recording of the waveforms (raw and after-SSP waveforms) of strip parts judged as defects and redisplaying the waveforms later. At present this system is used connected to the on-line Lamb wave inspection system for pickled strip.

CONCLUSIONS It was confirmed that the defect detection performance for sample strips with artificial defects was improved by optimum selection of SSP parameters. Furthermore, we succeeded in developing a system which completely synchronizes with the on-line Lamb wave inspection system for detecting defects at a repetitive frequency of 500 Hz and performs real-time SSP.

158

Y. Nagata et al.

[ Signalcollection

I

Existing equipment

I A/D conversion(20 MHz sampling)

lamb wave inspection I [ system

[ Bandpass filter (band" 2.0-2.5MHz) [

II

I Reducin~the numberof samolesto 1/10

I SS ca' u'a,io. I Fig. 8. Flow of data processing before SSP calculation for the realization of real-time calculation.

Developed system

~signalf

PR1] s i g n a l PG sign~ si~)

<~gl ~

Status ~ignal

~robe/~

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"~

t Functi~ ,..

I

Generated/received Lomb ".v~,:c

P

\

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#

Fig. 9. Block diagram of developed system.

REFERENCES

. .

6. 7. .

9. 10. 11. 12. 13. 14. 15. 16.

17.

Krautkramer, J. and Krautkramer, H. (1983). In: Part C: General Testing Technique, Ultrasonic Testing of Materials: Third Edition, Springer-verlag, New York. Alleyne, D.N. and Cawley, P. (1992)IEEE Trans, Ultrason. Ferroelectr. Freq. Control UFFC-39(3), 381. Datta, S.K., A1-Nassar, Y. and Shah, A.H. (1991). In: Review of Progress in Quantitative Nondestructive Evaluation, Vol.10A, pp.97-104, Thompson, D.O. and Chimenti, D.E. (Eds). Plenum Press, New York. Newhouse, V.L., Bilgutay, N.M., Saniie, J. and Furgason, E.S. (1982) Ultrasonics 20, 59. Karpur, P., Shankar, P.M., Rose J.L. and Newhouse, V.L. (1987) Ultrasonics 25, 204. Amir, I. (1986). PhD Thesis, Drexel University, Philadelphia, PA, U.S.A. Benharit, U., Kaufinan, J.L., Bilgutay, N.M. and Saniie, J. (1986) Proc. 1986 IEEE Ultrason. Symp., 1021. Bilgutay, N.M., Bencharit, U., Murthy, R. and Saniie, J. (1990) Ultrasonics 28, 90. Rose, J.L., Karpur, P. and Newhouse, V.L. (1988) Mater Eval. 46, 114. Xin, J., Donohue, K.D., Bilgutay, N.M. and Li, X. (1991) Mater Eval. Aug., 987. Gustafsson, M.G. and Sepinski, T. (1997) Ultrasonics 35, 31. Gustafsson, M.G. and Sepinski, T. (1993)IEEE Trans, Ultrason. Ferroelectr. Freq. Control UFFC-40(6), 659. Kraus, S. and Goebbels, K. (1980)NBS Pub. 596, 551. Bilgutay, N.M., Murthy, R., Bencharit, U. and Saniie, J. (1990) J. Acoust. Soc. Am. $7(2), 728. Aussel, J.D. (1990) Ultrasonics 28, 229. Nakamura, M., Yokoymaka, H., Yamano, M. and Murayama, R. (1998) In: Nondestructive Characterization of Material 1/711, pp.235-238, Green Jr., R.E. (Ed). Plenum Press, New York. Newland, D.E. (1984) In: An Introduction to Random vibrations and Spectrum Analysis 9Second Edition, pp.118-120, Longman Scientific & Technical, England.

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

159

RECONSTRUCTION OF WAVEFORM FROM PHOTOELASTIC ULTRASONIC VISUALIZED IMAGE T. MIHARA*, T. MUSASHI** and K. YAMANAKA* *Dept. of Materials Processing, Graduate School of Engineering, TOHOKU University Aobayama 02, Sendai, 980-8579, JAPAN

e-mail, [email protected]

**Tohoku University graduate school

ABSTRACT Since ultrasonic waveform by piezoelectric probe depends on each probe, reconstruction procedures are required for accurate waveform analysis. Photoelastic visualization method can estimate the quantitative ultrasonic traveling behavior in the glass specimen as the model of the industrial inspection in steel structures. However since ultrasonic wave transmit during flashing of stroboscope and intensity of visualized image is second power of sound pressure, reconstructions of visualized image for each factor must be requires for accurate waveform analysis.

In this study,

new process for quantitative waveform analysis was proposed combining the deconvolution using Jacobi method with the polarization recovering.

To confirm the availability of the process, surface

displacement on the specimen was measured by a laser interferometer and compared with the reconstructed waveform.

As the results, surface displacement agreed well with the reconstructed

waveform from visualized image and the availability of the process was confirmed.

KEY WORDS Photoelasticity, Ultrasonic Visualization, Waveform Analysis, Deconvolution,

INTRODUCTION An ultrasonic waveform generated by a piezoelectric probe depends strongly on the character of each probe.

Even in many nondestructive evaluations, an experimental waveform obtained on a

standard testing block by a piezoelectric probe is still utilized as a standard waveform. Then, reconstruction process eliminating the probe dependence is required for an accurate waveform analysis. These procedures are well known to be useful because slight features of the waveform in time domain or frequency domain have been utilized for a flaw characterization or a materials evaluation.

In order to investigate the availability of these analyses, a laser interferometer and a

capacitance transducer which measure the time dependence of surface displacement are developed. For more accurate and convenient analysis, a new technique that can observe the transmitting waveform in a specimen is also required. The photoelastic methods for visualization and evaluation of ultrasonic sound pressure have been

160

T. Mihara, I". Musashi and K. Yamanaka

studied [1-7].

We have also studied the ultrasonic visualization method using a photoelastic

procedure [8-10] to estimate an ultrasonic traveling behavior in a glass specimen as a model of an industrial inspection of a steel structure.

Since the intensity of the visualized image is proportional

to the square of the sound pressure, the quantitative behavior of ultrasonic transmission, reflection and directivity could have been investigated using this technique.

However in actual experiment,

the visualized image becomes diffuse because of the ultrasonics traveling during the flashing of the stroboscope.

In addition to this, the visualized image does not have a polarity because the

intensity of visualized image is proportional to square of the sound pressure.

Then reconstruction

of the visualized image eliminating these effects is required for estimation of measurement error. Though Date et al. pointed out this problem [5,6], an accurate reconstruction of waveform using a photoelastic visualized image remains unsolved. In this paper we developed the deconvolution analysis procedure to eliminate the stroboscope dependence and the recovery algorithm for polarity using an experimental loading method. Furthermore, to confirm the availability of this process, the surface displacement at the back surface was measured using a laser interferometer and compared with the reconstructed waveform.

As a

result, the waveform reconstructed with the above process agreed well with the waveform of a surface displacement.

ULTRASONIC VISUALIZATION AND STROBOSCOPE DEPENDENCE The experimental setup for a photoelastic ultrasonic visualization system is shown in Fig. 1.

Fig. 1 Photoelastic ultrasonic visualization system The system consists of a linear polariscope, a commercial stroboscopic light source (flashing time is about lOOns), an 8 bit CCD camera and a 16 bit digital image-processing system.

The trigger

pulse drives the stroboscope with some delay time and an ultrasonic transmission can be observed as a "frozen" image by varying the delay time.

The Pyrex glass of 20 mm in thickness was used as

a model specimen for steel because the velocities are similar to those of steel.

Figure 2 shows an

example of a visualized image and the intensity profile along the ultrasonic transmission direction

Photoelastic Visualization of Ultrasonics

161

of the solid line in Fig.2.

Example of visualized image

Distribution d intensity along a and b

Fig.2 Example of visualized image and intensity distribution In this experiment the principal angle of the polarizer and the analyzer was set perpendicular as a dark-field condition.

Thus the intensity of image I 1 in linearly polarized light is given hy

11 = a2sin2 20 sin2{Cd(crl - a2)}

(1),

where d is the thickness of specimen, C is the photoelastic constant, a is the constant, 0 is the angle between principal stress direction and principal axes of polarizer and 01 principal stresses.

and 02 is the

If we rotate the polarizer and analyzer together to change the principal axes by

45 degrees, keeping the dark-field condition, the intensity becomes 12 =a2cos 2 20 sin2{Cd(o~- crz)}

(2).

Since the intensity 11 and 12 depend on 0, a quantitative estimation is difficult because the intensity is depend on the ultrasonic transmission direction.

Then we used the image processing

technique adding the image 11 and 12 which depend only on the principal stress difference as shown by I = 11 + 12 = a z sinZ{Cd(crl-02) }

(3).

This equation can be approximated by the following because Cd(cr~ - t y 2 ) is much less than unity I = a 2 C2d2(c~1 - o"2

(4).

The ultrasonic sound pressure expressed as the difference in principal stresses t ~ - o 2 corresponds to cr~ for

a

longitudinal

wave

((Y2 = 0 )

and

to

2cT1 for

a

shear

wave

((Y2 =-'~I)-

RECONSTRUCTION OF THE WAVEFORM USING VISUALIZED IMAGE Elimination of the stroboscope dependence Since an ultrasonic wave travels a certain distance in a model specimen during the flashing of a stroboscope, a visualized image becomes diffuse as shown in Fig.3. We consider that the intensity distribution of a visualized image D(t) in the direction parallel to the ultrasonic transmission is a convolution of the original waveform Y(t) by the flashing waveform of the stroboscope H(t) as shown in a following equation D(t) = Y(t) * H(t) where the symbol * denotes the convolution.

(5),

T. Mihara, 7'. Musashi and K. Yamanaka

162

t=t 1

,.

( Start )

v

I

Y0(t)=D(t) (Measured Wave)

I

I

[ Yn(t)=Yn_l(t)+k (D(t)-Dn_l'(t))["~

v I 2 ~ ' ~

t=t 2

H(t)

Y(t) t1

t 2 t3

t

1' Convolution Cal. Dn'(t)=Yn(t),H(t ) [

v t=t 3

Waveform of stroboscope

Image of each(tl,t2,t3) Profile of intensity - - - - - - - ~ .... i.:~~'"...... di stributi on

( End )

Visualized image D(t) Fig. 3 Dependence of visualized image on stroboscope waveform

Fig. 4 Deconvolution algorithm for elimination of stroboscope dependence

.---. 1

E

Z v >,, 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.

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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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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

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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 <
(6)

The remelted zone in A1 alloy castings possesses fine grains and small amount of fine porosity. Contrary to this, the matrix has coarse grains, and contains a lot of coarse porosity. If ultrasonic waves are incident on the surface of a specimen, they first travel within the remelted layer, and then spread into the matrix. When the ultrasonic wave propagates through the remelted zone,

Fig.1 Schematic representation of reflection and back scattering of ultrasonic wave travelling in substance with remelted zone.

UE of Remelted Zone Depth in AI Alloy

187

the intensity of back scattering waves is low because of small attenuation coefficient caused by fine grains. While propagating through the matrix, it becomes higher because of large attenuation coefficient caused by coarse grains. This situation is illustrated in Fig. 1. Thus, it is possible to find a boundary between the remelted layer and the matrix from the intensity distribution of back scattering waves.

EXPERMENTAL PROCEDURES

Experimental material A plate of aluminum alloy castings of A1-4wt%Cu-5wt%Si was used in the experiment. The plate (150mm x 150mm x about 20mm) was partly melted by TIG arc welder to form a remelted zone on its surface. The remelted zone is shown as the hatched area in Fig. 2. The upper surface of the plate was polished with emery paper #500 for the ultrasonic measurement.

Ultrasonic measurement In the ultrasonic measurement, both the direct contact method and the water immersion method were employed. Experimental setups are illustrated schematically in Fig. 3.

Direct contact method. A probe generating a longitudinal wave with a frequency of 5MHz was used. Ultrasonic waves were transmitted from the probe through an oil-based couplant into the specimen. The ultrasonic waves were stored in a digital storage. Water immersion method. Focused probes with a longitudinal wave in frequencies of 10 and 15MHz were used. The diameter and the focal distance of the probes are 10mm and 25mm, respectively. The diameter of the ultrasonic wave is about one millimeter at a focal distance. A water path, the distance from a probe to the top surface of a specimen, was fixed at 25mm. The waveform measurement was carried out at an interval of one millimeter.

Fig. 2 Specimen used in ultrasonic measurement.

Fig.3 Typical ultrasonic measurement systems.

Metallographic observations After the ultrasonic measurement, the plate was sectioned and polished for metallographic examination. The specimen was etched with a solution of 10wt% NaOH to expose the

G. Jiang et al.

188

boundaries between the remelted zone and the matrix. The thickness of the remelted zone was measured with an optical microscope.

RESULTS AND DISCUSSION

Backscattering waves measured by direct contact method By using a longitudinal wave with a frequency of 5MHz, backscattering waves were obtained at positions with and without the remelted zone as shown in Fig. 4. Comparison of Figure 4a with Figure 4b shows that the intensity of the scattering wave just behind the first bottom wave becomes lower, although the intensity of the first bottom wave is almost the same. In addition, the intensity of the second bottom echoes is higher when the remelted zone existed.

nla,

.:, = ,-

. 1 ,

0

~ 0

_

=

-1

-1 I

0

2

4 6 Time (l~S)

8

10

0

2

4 6 Time (llS)

8

10

Fig.4 Ultrasonic waves measured with a probe of 5 MHz by direct contact method: (a) with remelted zone, (b) without remelted zone. These results show that the ultrasonic measurement by the direct contact method is sufficient for detection of the remelted zone. However, the direct contact method to evaluate a remelted zone size has following problems. (a) Since the intensity of the bottom echo depends heavily on a roughness and a flatness of the specimen bottom, evaluation of a remelted zone thickness by using the intensity of the bottom echo will cause significant error (especially in castings). (b) The measured area is almost the same as a probe size in the direct contact method, and it is very difficult to measure the thickness and the width of the remelted zone at small areas or with a good spacial resolution. In order to avoid the above problems, backscattering waves with a water immersion method of focused probes was employed.

Backscattering waves measured by immersion method Waveforms in Fig.5 were obtained with a probe generating a longitudinal wave of 10MHz in frequency. The intensity of the backscattering wave between the surface wave and the first bottom wave changes with the presence of a remelted zone. With existence of the remelted zone, the intensity of the scattering wave just behind the surface wave is very low; while it becomes higher at the matrix. In contrast, without remelted zone, the intensity of the scattering wave just behind the surface wave is high, and it decreases monotonically following elapsedtime (a depth from the surface). Figure 6 shows waveforms measured with a probe generating a frequency of 15MHz. Because of the high frequency, the attenuation of the reflected echo is so large that the backscattering wave behind the surface wave was not observed clearly, and the existence of the remelted zone

189

UE o f Remelted Zone Depth in AI Alloy

was not clarified. 0.1

0.1 A

t--

91

0

==

0

m

-0.1

0

2

4 6 Time(l~S)

8

-0.1

10

0

2

4 6 Time(l~S)

8

10

Fig.5 Ultrasonic waves measured with a probe of 10MHz by immersion method: (a) with remelted zone, (b) without remelted zone.

o,l

i

A

>-

0,tlL

4

>

"go 0

~:~:~-<---"

~---:-.

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~

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,11111, 2

I, I, 4 6 Time (l~S)

I, 8

/ 10

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4 Time

6

(rts)

8

10

Fig.6 Ultrasonic waves measured with a probe of 15MHz by immersion method: (a) with remelted zone, (b) without remetted zone. Energy integral and evaluation o f remelted zone thickness

From the above results, it is considered that the immersion method with a probe generating a longitudinal wave of 10MHz in frequency is effective for evaluation of a remelted zone thickness. The thickness of the remelted zone is obtained by measuring the elapsed time while the ultrasonic wave propagates from the top surface of specimen to the boundary between the remelted zone and the matrix. Since the backscattering echo changes sinusoidally, inferring the boundary position from the change of the scattering waveforms may give rise to a deviation from the true position. Therefore, the intensity of backscattering echo was squared and integrated to decrease sinusoidal change in the echo. Since the energy of wave is proportional to the square of the amplitude, namely, the intensity of the wave, it can be defined as the following integral:

e (x) = Io [e (r

(7~

where P( ~ ) is the intensity of the ultrasonic wave, and x is a distance from the specimen surface. Typical energy integrals of the ultrasonic wave are shown in Fig. 7. In these integrals, the energy integral of the surface wave was included. As shown in Fig. 7, with a remelted zone, an inflection point occurs in the energy integral curve. In contrast, there is no inflection point in

190

G. Jiang et al.

the energy integral curve without remelted zones. Therefore, a remelted zone thickness can be inferred from the inflection point of an energy integral curve. The remelted zone thickness obtained from this method is shown in Fig. 8. A section analysis was also employed to measure the remelted zone thickness with an optical microscope. A typical photomicrograph at the boundary of the remelted zone and the matrix is shown in Fig. 9. The remelted zone thickness obtained with the optical microscope is also shown in Fig. 8.

~

26.3

N.. 35.6

~

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~

6

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4

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, I,

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depth(mm)

8

10

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depth (mm)

Fig.7 Change in integral of square ofultrasonic wave intensity with depth fi'om surface: (a) with rermtted zone, (b) without rernelted zone.

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section analysis

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(a) position from A to B

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2

0 A

5

10

15

20

25

(b) position from A to C

30 C

dR with positions.

As can be seen from Fig. 8, the thickness and the width of the remelted zone measured with the ultrasonic measurement are extremely close to those measured by inspection of the section. The error was less than 6%. Therefore, it is feasible to evaluate nondestructively the remelted zone thickness in A1 alloy castings with the remelted zone of the same phases and structures as the matrix. By using this method, the remelted zone thickness was obtained within 2 minutes for each point. Instead, the section analysis took several hours.

CONCLUSIONS A new method based on ultrasonic measurement is developed to evaluate the remelted zone size in A1 alloy castings. The thickness and the width of the remelted zone are evaluated by measuring the intensity distribution of the backscattering wave, which depends on the size and

UE of Remelted Zone Depth in AI Alloy

191

the distribution of the crystal grain. The remelted zone thickness and width evaluated from the analysis of the backscattering wave intensity are in good agreement with those measured by the destructive section analysis. The error is less than 6%. And by using this method, the remelted zone thickness is evaluated in a much shorter time than the section analysis.

Fig.9 Microstructure at boundary of remelted zone and as-cast matrix.

REFERENCES

1. 2. 3. 4. 5. 6. 7.

E.P. Papadakis and B. W. Peterson (1979) Materials Evaluation 37, 76. E.P. Papadakis (1998) J. Testing and Evaluation 26, 240. R.L. Smith (1982) Ultrasonics 9, 211. J. Lewndowske (1986) Ultrasonics 3, 73. L.M. Shen and H. Kato NDT&E International (submitted). H. Quincy and E. Steve (1997) Materials Evaluation 55, 1323. T. Mihara, M. Obata and T. Furakawa (1989) The Japanese society for Nondestructive Inspection 38, 583.

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

193

ACOUSTIC EMISSION MEASUREMENT BY LASER I N T E R F E R O M E T E R IN CERAMIC COATINGS

Makoto Watanabe*, Manabu Enoki* and Teruo Kishi**

* Department of Materials Science and Engineering, Graduate School of Engineering, The University of Tokyo, 7-3-1, Hongo, Bunkyo-ku, Tokyo, 113-8656 JAPAN ** National Institute for Advanced Interdisciplinary Research, 1-1-4, Higasi, Tsukuba Science City, 305-8562, JAPAN

ABSTRACT This paper describes an application of laser techniques for the detection of acoustic emission (AE) signals on fracture behavior study in elevated temperature. In this research, the AE signals from thermally loaded AlzO3/SUS304 plasma sprayed coatings with different layer thickness were detected using four Laser Doppler Vibrometers (LDV) during thermal cycle between 1000 C and room temperature. AE behavior showed strong dependence on the thickness of ceramic layer. Also, the AE location was determined from the difference of wave arrival time among each laser-focused position, and the fracture mode of some AE events were analyzed using the inverse analysis technique. The results demonstrated that the AE measurement by laser techniques has an advantage under high temperature environment, though there are some limitations and difficulties to be improved, KEYWORDS Acoustic emission, Laser interferometer, Ceramic coating, Inverse analysis.

INTRODUCTION The thermal barrier coatings (TBC), which consist of a ceramic layer applied over a metallic bond-coat layer, have been developed to prolong the life of hot-section components and to improve the thermodynamic efficiency of internal combustion engines and gas turbines. However, the performance of currently available TBCs is considered as not sufficiently reliable in practice. Some of the major technical issues for the development of reliable TBCs and the prediction of lifetime are (i) understanding the failure mechanisms of current TBCs in actual turbine environments, (ii) developing characterization and test methods to measure and understand fracture behavior of TBCs dependent on time, temperature, and environment. Especially, there are lots of difficulties to monitor the fracture at the elevated temperature environments. The one of main purpose of this work is to develop the monitoring method based on the laser techniques in order to solve these issues. Acoustic Emission (AE) method is the best technique for in-situ monitoring of fracture behavior, which provides real-time information on damage process. AE signals are usually detected by attaching a piezoelectric transducer on a sample. However, using a piezoelectric

194

M. Watanabe, M. Enoki and T. Kishi

transducer has several limitations. In particular, since a piezoelectric transducer must be in direct contact with a sample, it cannot be used in harsh environment like high temperature environment. In order to apply AE methodology to thermal cycle test, a much larger waveguide than a specimen is needed between a specimen and a transducer [1,2], which has large effects to the detected signals with the transfer property of it. The application of a waveguide makes it difficult to analyze AE waveforms quantitatively and results in the qualitative analysis. In recent years, numerous efforts have addressed the substitution of laser-based techniques for ultrasonic nondestructive evaluation in place of conventional piezoelectric transducers [3,4]. The main advantages of a laser interferometer are (i) non-contact detection, (ii) small focus area, and (iii) absolute measurement. Since a laser interferometer makes no physical contact with the surface being observed, it can have a possibility to detect a seismic AE event in high temperature environment. A laser interferometer can be used to measure Doppler-shifted scattered light from which the displacement or velocity of being monitored can be calculated and it is expected to perform accurate analysis for AE source wave. Though there are some disadvantages such as the high sensitivity to noise, low S/N ratio, and the difficulty of handling, however, the effective and useful method for fracture evaluation of materials like TBCs in high temperature environment may be constructed by using a laser interferometer as an AE sensor. Until now, there are few reports referring to the detection of AE signals in the practical materials because of the above difficulty of experiments. The objective of this paper is to study the possibility of using laser AE system in high temperature environment and to study the quantitative evaluation of fracture behavior in thermal barrier coatings.

EXPERIMENTAL METHOD A specimen was deposited by plasma splaying of A1203 onto SUS304 substrate with three coating thickness, 100, 500, 1000~tm. The diameter and the thickness of a substrate were 30mm and 5mm respectively. Since a large difference of the thermal expansion coefficient between A1203 and SUS304 was expected to cause fracture to easily occur and to generate comparatively large amplitude AE signals, this system was selected for a first trial of AE measuring by a laser interferometer at the elevated temperature. The specimen was heated from room temperature to maximum temperature 1000~ with an infrared image furnace, and the temperature was controlled by a thermocouple attached on the surface of the specimen in heated zone. The thermal treatment consisted of a constantly heating of maximum temperature with a rate 15~ a 10 seconds holding at maximum temperature, and a cooling to room temperature in air. During the whole period, heating, holding, and cooling, four laser interferometers were simultaneously applied to detect AE signals generated from the fracture of coating layer and, over more, to estimate the location and the fracture mode of an AE event as shown in Figure l(a) and (b). Each laser beam was lead to the specimen with four reflection mirrors and was focused on the opposite side of heated side, which surface was polished enough to get the reflected beam. The laser interferometer is a commercial heterodyne one (Graphtec Co., Ltd., AT0022), which is so called a Laser Doppler Vibrometer (LDV) with He-Ne laser and can directly measure the velocity of objects more than 2/zm/s. A Law Pass Filter (LPF) 300kHz was used to decrease noise level.

195

AE Measurement by Laser Technique

a)

Infrared Image Furnace PC Wave memory ....... :.:~:~i~:~i:~:~ Heating ~/~

FM

Thermocouple

Demodulator 'x/dZ

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.

He-Ne laser

1 /-

...........

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FIGURE 1. Experimental setup for acoustic emission measuring using the laser interferometer during thermal cycle test, a) Schematic of measuring system, b) detail of laser focusing to a specimen. Detected AE waveforms were recorded by the wave memory (JT toshi Co., Ltd., DCM-120,) with sampling rate of 200ns and 2kwords.

RESULTS AND DISCUSION Figure 2 shows a typical temperature history and dtected AE events with the first peak amplitude of each AE waveform for three specimen types. While almost specimens with coating thickness 500gm and 1000~tm were fractured at one or two cycle, for the specimens

196

M. Watanabe, M. EnokJ and T. Kishi

with coating thickness 100~tm, only the partial delamination of the ceramic layer was observed after four or five cycle. There were the obvious differences of AE behavior among the specimen types. In the coating thickness 1000gm, all AE events were only detected in the heating and were not detected in the cooling period. On the other hands, in the coating thickness 500 and 100~tm, all AE events were only measured during the cooling. The failure during heating in the thickness 1000~tm is considered to be caused by a lot of initial cracks, which were induced during the plasma splaying because the thermal stress due to the difference of the thermal expansion coefficient has tends to increase with an increase of the coating thickness. The AE behavior of the specimen of 5001am coating shows the characteristic fracture process of ceramic coatings. Namely, after some AE events (marked as group A in Figure 2) were almost simultaneously generated within a few seconds, next group of AE signals (group B in the figure) were detected in 300 seconds later, group C in more 500 seconds later, but little difference in the temperature, and the ceramic layer were perfectly delaminated with the detection of last AE. The progress of AE events exhibited step-by-step manner, which means the propagation of the generated cracks were interfered and the criteria of crack propagation were satisfied and again a crack began to propagate. The differences of temperature from A to B and B to C were little and it is very important to understand the effect of incubation time for understanding of coatings failure.

AE location

Based on the assumption of homogeneous material since the thickness of coating is enough less than the substrate, the AE location can be evaluated from the difference of longitudinal wave arrival time [5]. The AE location analysis results for the sample of thickness 500~tm is 1200 1000

0.01

1200

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1200

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1st cycle 2nd cycle I--] 3rd cycle <~ 4th cycle O A

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Figure 2. Typical temperature history and detected AE events, a) 100, b)500, c)1000~tm.

5000

E

AE Measurement by Laser Technique

197

shown in Figure 3, in which the circle stands for the specimen area and the triangle symbol stands for the each AE location. The number in the figure means the order of each AE event. After the first fracture was recognized near the edge, the crack propagated along the circumference of the specimen by priority and made progress to the center region finally. From the SEM observation, at the edge of the specimen the fracture occurred at the interface, while the fracture in almost area except the edge was recognized in the ceramic layer just above the interface. Since vertical cracks for the interface were not observed, the detected AE signals were considered to be generated from this delamination at the interface and near the interface. From the results of the SEM observation and the AE location, it is understood that the crack initiated just at the interface of the edge and grew along the periphery of the specimen, while the initiated crack was deflected into the ceramics layer parallel to the interface and propagated over the specimen area. This fracture behavior is considered to be caused by the stress concentration occurs at the interface of the edge of the specimen materials due to the difference of the thermal expansion coefficient.

Fracture mode analysis The fracture mode of an AE event was estimated in order to understand fracture behavior of a material using the moment tensor analysis. In the calculation, it was assumed that (i) the material consists of two different material layers, (ii) each material is homogeneous, isotropic, and elastic, (iii) the cracking occurs at the vicinity in the ceramic layer at the interface, and (iv) the normal vector of crack surface is parallel to the interface normal vector, that is, the delamination parallel to the interface. Figure 4 is the source analysis result for seven events of the 500~tm coating, in which the horizontal axis exhibits the angle between the normal vector of crack surface and the discontinuity vector of the displacement of the crack, therefore, if 0 = 0 ~ it corresponds to the mode I, tensile fracture type and if 0 = 90 ~ it stands for mode II, shear fracture type. The vertical axis stands for the distribution number. Though almost evaluated AE events had 9 AE position ..............~.5"'"'~....................... ...............~"~3

2A

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

.............. 1~110 /':" 13 A AIO

/

,z.

B -

...................... A4 ...."._.

-5 -.9 '.........-...

......

12

z ,!

5 -5

6

9A A6

-10

.:..-

.......""

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

................ ............... ........................L..........................

-i5

X (mm) Figure 3. AE location result for coating thickness 500~tm. The number stands for the order of AE events

M. Watanabe, M. Enoki and T. Kishi

198

mixed mode of mode I and mode II, the component of shear mode tended to be large, therefore it was considered that the microfractures were mainly generated by the shear mode in this experiment. Since the thermal expansion coefficient of the substrate SUS304 is larger than that of the ceramic layer, if the temperature distribution at the holding period is uniform over the specimen, the compression stress perpendicular for the interface is induced during cooling at the edge, as the result, the mode II fracture will occur more easily than the mode I fracture. Though the calculation process to derive the Green's function is extremely complicated, however, it is very significant to be able to estimate the fracture mode of cracks generated during the thermal cycle by using the laser technique and the AE source analysis technique.

N=7

4 ID

z

3 2 1

-

0

0-30

30-60 0( o )

60-90

Figure 4. Frequency distribution for the angle between a crack surface and a discontinuity vector CONCLUSIONS The fracture behavior of A1203 coatings is evaluated by the AE detection at elevated temperature using laser interferometers. AE signals during the thermal cycle could be detected and it was shown that the laser interferometer has a good performance as an AE sensor even if it has some limitations to be improved. Different AE behavior with the thickness of coatings was observed, and the coatings were easily broken with the increase of thickness of coatings. AE location during thermal test can be successfully evaluated by the multicannel AE measuring. Initially the crack was generated at the interface of the edge of the specimen and propagated into the ceramic layer and the internal region. Fracture mode of some AE signals was estimated by the FEM wave analysis and the deconvolution method. Almost AE event occurred under the mixed mode, and the contribution of mode II were large and the cracking was considered to be induced by the shear fracture mode.

REFERENCES

1. 2. 3. 4. 5.

Berndt, C. C. (1989)Journal of Materials Science 24, 3511. Nakahira, H. (1991)Journal of the Society of Materials Science, Japan 40, 24. Palmer, C. H., Review of progress in Quantitative Nondestructive Evaluation 5A, 651. Bruttomesso, D. A. (1993)Journal of Engineering Mechanics 119, 2303. Enoki, M., and Kishi, T. (1988) International Journal of Fracture 38, 295.

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

199

ULTRASONIC EVALUATION OF COCrA1YCOATING WITH LEAKY RAYLEIGH WAVES Koichiro KAWASHIMA, Hiroshi FUJ1TA, Toshiyuki SHIMA Department of Mechanical Engineering, Nagoya Institute of Technology Gokisocho, Showaku, Nagoya 466-8555, Japan Ichiro KOMURA and Takahiro KUBO Toshiba CO. ABSTRACT The mass density and Young's modulus of CoCrA1Y coating on INCONEL substrate were nondestructively evaluated by the velocities and attenuation of the leaky surface waves which were measured with line focused PVDF transducers. The measured values agree within 10% with those measured by the conventional methods. The aging effect on the coating under nitrogen atmosphere at 1273K for 500 hrs. was also detected by the change in the leaky Rayleigh wave dispersion and attenuation. KEYWORDS

Coating, Leaky surface wave, Ultrasonic velocity, Ultrasonic attenuation

INTRODUCTION Many kinds of anticorrosion coating or thermal barrier coating are widely used for steam and gas turbine blades. The health monitoring of these coating during service is important issue for the structural integrity. Several nondestructive techniques, such as ultrasonic testing, thermography, X-rays, are applied for the detection of the damage or cracks in the coating. Among them, the ultrasonic surface wave technique is best for this purpose because the surface wave is propagated only in the surface layer of one wavelength thick. The V(z) technique in acoustic microscopy [1] has been used for velocity and attenuation measurement of the leaky surface waves, however, the penetration depth of the leaky surface wave of most metals or ceramics is less than 20 ~tm when the operating frequency is over 200 MHz. For evaluation of plasma sprayed coating of a few hundred micron thick, the measurement at lower frequency is required. For frequency range of 10-50 MHz, we can measure precisely the velocities of the leaky surface waves by using time domain pulse signals excited by PVDF focused transducers [2-4]. In addition, the mass density and Young's modulus [5] in the surface layer can be nondestructively estimated by solving Viktrov's equation [6] of the leaky Rayleigh wave with the measured velocities and attenuation of leaky surface waves. The technique has been applied for the measurement of the mass density and Young's modulus as well as aging of plasma-sprayed CoCrAIY coating on INCONEL 738LC substrate. In addition, the thermal degradation of the coating is evaluated with the velocity dispersion of the leaky Rayleigh wave and attenuation coefficient.

PRINCWLE OF VELOCITY AND ATrENUATION MEASUREMENT

Measurement of leaky surface wave velocities[4] The excitation and reception of the leaky surface waves by a line-focused PVDF sensor is modeled as shown Fig.1. Among infinite beam paths, we can detect in time domain only the signals of the direct

K. Kawashima et al.

200

reflection and of the paths for three critical angles on the water/solid interface, as shown in Fig.2. For small defocus z, we identify only signals of the direct reflection and leaky Rayleigh wave, however, three leaky surface waves, namely, creeping, SV and Rayleigh wave, are identified at large defocus. When the sensor has spherical or cylindrical surface of the focus F and the sensor axis is perpendicular to the sample plane, the time delay of each surface wave for an increment of the defocus is given by

Ati = 2Az( tanOiiV

Vw cosoil) - miAz

(1)

where 0 i is the critical angle of each wave, Vw is longitudinal wave velocity in water, V/is each wave velocity. With Shell's law, each wave velocity (energy velocity) is given by

Vi = Vw [l_ Vwmiz

-1/2

4

(2)

When the sample surface has some scattering sources, the leaky wave velocity shows dispersion. The phase velocity of the leaky Rayleigh wave is calculated by measuring the phase difference at two defocuses, zl and z2, namely

2nfaL Ar

O)

v~(f)

-

ar

= r ( f ) - [r ( f )

4ffgf (Zl -

-

--Z2)]

(4)

V., cos0~

(5)

AL - 2tan0R(z , - z2)

where the second term in the square brackets denotes the phase correction for the water path shown in Fig.1.

Leaky Creeping Wave

~D

Leaky Rayleigh wave .

Surface SV Wave Time Fig.1 Beam paths corresponding to critical angles excited by line-focused PVDF sensor.

Fig. 2 Leaky creeping, SV and Rayleigh wave signals on time domain.

Attenuation measurement of leaky Rayleigh wave The amplitude of the leaky Rayleigh wave is expressed in terms of the defocus z, attenuation coefficient a~ in water and that a at the solid/water interface [7]. A(z) -- A o exp[-2aw (F - z / cos0R)]exp(-2oz tan 0R)

(6)

from which we obtain the attenuation coefficient c~. a =

0% sin 0R

1 d[lnA(z)] dz 2 tan 0R

(7)

Ultrasonic Evaluation of Coating

201

ESTIMATION OF MASS DENSITY AND EIASTIC MODULI The governing equation of the leaky Rayleigh wave traveling along the interface between an isotropic elastic solid and water is given by

4kZqs_(k2 +s2~

=

i pL

qkt 4

,q2 = k 2

_ _

kl2,s2

= k 2

-

-

kt 2

(8)

_k

where kl, ko k are the wave number of the longitudinal, transverse and leaky Rayleigh waves of a solid and kw is the wave number of the longitudinal wave in water, p and Pw are the mass density of the solid and coupling water. The real and imaginary parts of k give the velocity and attenuation of the leaky Rayleigh wave. When we measure k~,k, and k with the technique shown later, we can estimate p with the known values of kw and p~ With the values of kl, k, and p, the elastic moduli are calculated by G

pV, 2, E

pV, 2

4~: - 3

~2 _ 1

2~2-1 2(~ 2 - 1)' ~ - V, / V/

(9)

For water-free surface, eq. (8) gives the well known equation of the Rayleigh wave.

/ /6

/ +8[

:0

10,

This equation is used instead of eq. (8) to determine Vtwhen the measurement of Vt is difficult on time domain with a surface wave transducers.

EXPERIMENTAL METHOD The digital signal acquisition system for leaky surface waves is shown in Fig.3. The transducer is translated along X, Y and Z axes by i ~tm and rotated around these axis by 0.002 degree. Two linefocused PVDF transducers were used; the nominal frequency of 10 and 36 MHz, the focal length of 15 and 5 mm, respectively. The received signals were fed to an A/D converter and acommlated many times into a digital memory to enhance S/N ratio. The signal of the leaky surface waves were recorded for two samples of plasma-sprayed CoCrAIY coating on INCONEL 738LC substrate with or without thermal aging at 1273 K for 500 hrs. in nitrogen atmosphere. The surface was finished by grinding after coating. The coating thickness is 0.2 mm and 0.125 mm for the sample without or with thermal aging. The micrograph of the samples are shown in Fig.4. After thermal aging, the aluminum nitride particles precipitate in the coating.

202

K. Kawashima et al.

Pulser/reciver

~~,

I,,

I

M~176c~176

1

Personal computer

Personal computer

Fig.3 Digitalsignal acquisition system.

Fig.4 Micrograph of plasma-sprayed CoCrA1Y coating :(a) no aging, (b) after aging at 1273K for 500 hrs.

EXPERIMENTAL RESULTS Received waveforms

As an example, Figs. 5 and 6 show the received waveforms of the sample before and after thermal aging, where the nominal center frequency is 36 MHz and the propagation distance of the surface waves is about 4.0mm. The signal of the longitudinal creeping wave is identified in Fig.5, however, it is masked by noise in Fig.6. The separation of leaky SV and Rayleigh waves is impossible for these samples. Thus, the transverse wave velocity is calculated by eq. (10).

Ultrasonic Evaluation of Coating 50

I

'

I

'

i

'

I

,

203 I

X -50 6

5.8

6.2

I

6.4

6.6

Time [~tsl

Fig.5

Received waveform of coating without aging.

50 ~

r

l

T

l

vV ' - - - - 2

0

g -50 5.8

6

6.2

6.4

6.6

Time [~tsl Fig.6 Received waveform of coating with aging. Based on these waveforms, we calculated the velocities of the longitudinal creeping wave and leaky Rayleigh wave by using the cross correlation method. The results are shown in Table.1. Apparently, the leaky Rayleigh wave velocity of the aged sample is higher than that without aging. This change results from the precipitation of acoustically hard aluminum nitride. Table 1. Leaky surface wave velocities and attenuation. Frequency (MHz) 10 36

VR(No aging) (m/s)

VR(Aged) (m/s)

3119.0 3138.3

3199.5 3344.3

Vt (No aging)

(m/s)

6073

Figure 7 shows the attenuation coefficient of the leaky Rayleigh wave, measured by 36 MHz transducer. Though there are some scattering in attenuation point by point, the attenuation coefficient on the sample before aging is 0.63 +--0.03/mm at 32 MHz. With the velocities shown in Table I and this attenuation, the mass density of the coating before aging is estimated by eq.(8). The estimated values are compared in Table 2 with those measured by the conventional methods. The estimated mass density and Young's mudulus agree within 10% with ones measured by the conventional methods. Figure 8 shows the dispersion curves of the samples without and with aging. We have used 10 and 36 MHz transducers, therefore, we could not measure reliable dispersion curve in the frequency of 16-25 MHz. Nevertheless the dispersion curves show nice continuity for the frequency rage from 8- 50 MHz. At 10 MHz, the wavelength is about 0.3mm which is greater than the coating thickness, 0.2ram and 0.125 nun, therefore the dispersion curves are affected by the substrate. On the contrary, the wavelength at 40 MHz is about 0.08ram which is much smaller than the coating thickness, thus the dispersion curves show the coating characteristics. Apparently, the aging results in higher velocity and higher

204

K. Kawashima et al.

attenuation, as shown in Figs. 7 and 8. Thus we can detect the change in the coating layer by the measurement of the velocity and attenuation of the leaky Rayleigh wave. I

_

'

I

.

'

I

9

2

9

3400 I

~l

o-I

,

I ' 1 ' 1 '

0.8 3300

_

.~~

; I

q).7

3200-

' . I z

"

0

,~

0

0.6 28

311

32

34

3100

'~

3000 I

10

,

I

20

,

I

30

,

I

40

,

50

Frequency [MHz]

Frequency [MHz] Fig.7 Attenuation coefficient

No aging

Fig.8 Dispersion curves.

Table 2. Mass density and Young's modulus of CoCrA1Y coating.

Mass density (kg/m3) Youn~;'s modulus(GPa)

Ultrasonic measurement 7.7 • 103 230

Conventional measurement 7.3 • 103 200

CONCLUSION The mass density and Young's modulus of CoCrAIY coating on INCONEL 738LC substrate have been estimated with the leaky wave velocities and attenuation. The estimated values agree within 10 % with those measured by conventional methods. The velocity and attenuation of the leaky Rayleigh wave of the aged sample are higher than those without the aging. Consequently the degradation of coating can be evaluated by the velocity and attenuation of leaky Rayleigh wave.

REFERENCES 1. Briggs, A.. (1992). Acoustic Microscopy, Oxford University Press. 2. Xiang, D., Hsu, N.N. and Blessing, G.V.(1996). Ultrasonics, 34, 651. 3. Gondard, G., Tady, F., Noroy, M.H. and Paradis, L. (1996). in Review of Progress in QNDE, 15, 1605. 4. Kawashima, K., Fujii, I., Takenouchi, N. (1997). Trans. JSME, Ser. A, 63, 2444. 5. Kawashima, K., Fujii, I., Takenouchi, N. (1998). in Nondestructive Characterization of Materials VIII, 707. 6. Viktrov, I.A.(1967).Rayleigh andLamb Waves, Plenum Press. 7. Kushibiki, J. and Chubachi, N. (1985). IEEE Trans. SU-32,189.

Nondestructive Characterizationof MaterialsX Green et al. (Eds) 9 2001 ElsevierScienceLtd. All rights reserved.

205

QUANTITATIVE EVALUATION OF SiC/NiP COMPOSITE COATING BY SURFACE ACOUSTIC WAVE SPECTROSCOPY AND NANO-INDENTATION

IKUO IHARA, TAKESHI SAWA AND KOHICHI TANAKA Department of Mechanical Engineering, Nagaoka University of Technology, Nagaoka, Niigata 940-2188, Japan

ABSTRACT In this study an ultrasonic quantitative evaluation of a composite coating is performed and the ultrasonically determined elastic constants are compared with mechanically determined ones in order to examine the present ultrasonic technique for its validity. An inversion technique with surface acoustic wave data and a nano-indentation technique are here used to determine elastic constants of coatings ultrasonically and mechanically, respectively. The experimental results for a NiP coating have shown that the estimated Young's moduli by ultrasonically and mechanically coincide with each other in an error of 2 %. This demonstrates that the present ultrasonic technique has been successfully applied to the elastic constant determination of such coating. Furthermore, SiC/NiP composite coating (SiC particle size and its content are about 0.2 ~tm and about 15 % by volume, respectively) of 10 ~tm in thickness was examined. The estimated results by the surface wave spectroscopy indicated a 32 % increase in Yotmg's modulus and a 7% decrease in density due to the randomly dispersed SiC particles. The validity of this result is also discussed in comparison with mechanically and theoretically estimated results. KEYWORDS Surface acoustic wave, Nano-indentation, Composite coating, Inverse analysis, Elastic constant

INTRODUCTION Coatings have been widely used in the field of surface engineering. For mechanical and tribological applications, particulate metal matrix composite coatings are now in high demand owing to the capability to enhance wear resistance as well as mechanical strength. For design and manufacturing of such particle reinforced composite coatings, the elastic properties of the coatings should be known accurately because they are closely related to the tribological and the mechanical characteristics. It is known that the material properties of coatings depend on the deposition process and is not always the same as bulk properties. One of significant factors that determines the overall material properties of such particulate composite coatings is the spatial distribution of the particles within the matrix. It would be normally difficult to estimate the elastic properties of the composite coatings theoretically. Therefore, the problem of determining the elastic properties of such composite coatings is of technological importance for the practical use of the coated materials. Surface acoustic waves (SAWs) have been reported as an effective means for probing the characteristics of materials surface [1-3]. The

206

I. Ihara, T. Sawa and K. Tanaka

combination of a SAW measurement technique and an inverse analysis allows quantitative determination of the elastic properties of coatings [1-5]. It is therefore worth applying the SAW technique to the evaluation of the composite coatings. One of the purposes of this work is to determine quantitatively the elastic constants of a composite coating by SAW. Although, to date, many investigations on elastic constants determination using the SAW techniques have been made [ 1-5], there is little known about the validity of the estimated elastic constants because of the difficulty in verifying it. It is essential to assure that the estimated value by the SAW is appropriate. Therefore, another purpose of this work is to examine the validity of the ultrasonically determined elastic constant. In the beginning, elastic constants determinations of a NiP coating and a SiC/NiP composite coating are performed using a SAW spectroscopy technique and then the determined Young's moduli are compared with the mechanically determined ones using a nano-indentation technique. Furthermore, the validity of the experimentally estimated Young's moduli of the SiC/NiP composite coating is discussed in comparison with the theoretically estimated one.

SAW SPECTROSCOPY The velocity of a SAW propagating over a thin film coated material depends on frequency. This frequency dependence of the SAW velocity, so-called SAW dispersion, is strongly influenced by the materials parameters of the layered structure such as elastic constants, mass densities, film thickness and the interface conditions between the film and the substrate. Therefore, when some of the materials parameters of the layered structure are known, the remaining materials parameters could be determined from the SAW dispersion curve [1-5]. Because of the absence of an analytic expression for the SAW dispersion curve, an appropriate inversion technique should be used for determining some tmknown materials parameters from the SAW dispersion curve. In the present work, the inverse problem to determine elastic constants and density of a coating film is def'med, where the coating thickness and the materials properties of the substrate are known and each material is isotropic [4]. The inverse problem is here reduced to the optimum problem which determines a set of parameters (Young's modulus E, Poisson's ratio v and density/9) minimizing the objective function r defined as r =~

)VsAw(i)--VsAw(i )t 1• W ( II"

(1)

N i=l

where V'SAW(/)and 1,'SAW(/) are discrete data of the calculated and measured dispersion curves, respectively, N is the number of data points, and IT'(/) is weighting coefficient assigned to each data point. In this work, W(/) is taken to be unity. The Complex method [6], a numerically interactive searching procedure, is used to solve the optimum problem [4, 7]. This method, starting with 2n (n is the number of unknown parameters) feasible points which is given by using random numbers, determines the optimum parameters within constrained regions by a typical descent procedure. In general, the most essential aspect in solving an inverse problem is the stability and the uniqueness of the solutions when only a limited amount of experimental data is available. In order to investigate the convergence of the solutionS for the present inverse problem, the behaviors of the objective function r in the solutions spaces to be explored are examined by

Evaluation of Composite Coating

207

using a numerical example. Figure 1(a) shows the contour plot of r in the vicinity o f the true solutions in the error field for E - p, while v remains constant to be the true value. Figures 1(b) and 1(c) show the contour plots for E - v and p - v, while p and E remain constant to be the true values, respectively. The distributions of r in the error fields were calculated by the use of the SAW dispersion curve data for a NiP coated steel. An obvious global minimum which corresponds to the true solution (r=0) is observed in each figure, and no local minimum is observed except for the global minimum. There was no local minimum even in the contour plots for the exploring area within + 30%. Also the contour lines in each figure are continuous around the global minimum points. This suggests that the stability and the uniqueness of each solution exist in the inverse problem. Actually, it was confirmed by performing the inversion o f the artificial dispersion curve calculated from a known set of parameters that the three parameters of coating can be successfully determined with a high accuracy [4]. However, it is also found through the investigations with the error fields that it must be extremely difficult to fred the stable solutions when the measured dispersion data have a certain amount of error and/or when the objective function r does not converge to a small value in the optimum process. In this work, to improve the accuracy of the solutions obtained by the inversion, we performed ten times inversions independently and then took an average of the obtained values.

Fig. 1. Changes of the objective function r with the materials properties of NiP coating deposited on a steel, where coating thickness = 10 lam and frequency range = 20 - 100 MHz are used in the calculation, and r = 0 corresponds to true solutions: (a) error field on E and p ; (b) error field on E and v; (c) error field on p and v.

NANO-INDENTATION The nano-indentation is an effective means for probing the mechanical properties o f materials on the nanometre scale[8, 9] and can be easily adapted for determining the elastic constants of coating films [10]. In this method, the applied load P and the penetration depth h of the indenter are continuously measured as an indentation test is made, and then the mechanical properties are estimated from the obtained P-h curve. Figure 2 shows a schematic representation of such a P-h curve for an indentation experiment. Basically, the elastic constants of materials are related to the indentation load-displacement behavior through the following equation [8, 9]: S = aE',fA

(2)

208

I. Ihara, T. Sawa and K. Tanaka

Fig.2. Typical load-displacement curve obtained by an indentation experiment: Pm~x denotes the maximum indentation load; hmax the indenter displacement at the maximum load; hf the final depth of the contact impression after unloading; and hs the displacement of the surface at the perimeter of the contact. where S is the experimentally measured stit~ess during the initial stage of unloading, a is a geometrical factor equal to 1.166 for a triangular pyramid indenter [8], E* is the effective Young's modulus, and A is the projected real contact area between the indenter and the specimen surface. The effective Young's modulus E* is related to the Young's moduli and Poisson's ratios for the specimen and the indenter through the equation:

1/E* = ( 1 - v ~ ) / E s + ( 1 - v ~ ) / E I

(3)

where Es and vs are Yotmg's modulus and Poisson's ratio for the specimen and El and vl are the same parameters for the indenter. The projected real contact area A for a triangular pyramid indenter is calculated from the equation: A = 3-4"5tan 2 65 ~h~

(4)

where hA is the vertical distance along which contact is made and is calculated from the equation [9, 10]: hA =hmax - 0.726 (hmax - h s ) + A h T

(5)

where AhT is the tnmcation length of the indenter tip. In this work, the value, AhT = 8.7 nm, which was determined from the experiment data for a steel standard specimen, is used. Thus, we can determine the effective Young's modulus E* using Eq. (2) with the stiffness S determined from the P-h curve and the real contact area A obtained from Eqs. (4) and (5).

EXPERIMENTAL PROCEDURE Coatings used here are a NiP coating and a SiC-particle-reinforced NiP composite coating (SiC/NiP, SiC particle size and its content are about 0.2 ~ and about 15 % by volume, respectively). Owing to the randomly dispersed SiC particles, the SiC/NiP coating is expected

Evaluation of Composite Coating

209

Fig.3. Photographs of cross sections for coatings used: (a) SiC/NiP coating; (b) NiP coating.

to enhance its mechanical properties. Each coating is deposited on a steel substrate by electroless-plating at 90 ~ and then is annealed at 300 ~ The layer thickness is 10.0/an for SiC/NiP coating and 10.5/an for NiP coating. In the X-ray diffraction measurements for the coatings, not only broadened diffraction pattern from an amorphous-like Ni phase but also sharp peaks from a partially crystallized Ni3P phase were observed. Since each coating is formed under the same deposit condition, it is considered that the material properties of the NiP matrices for both coatings are almost the same each other. Figure 3 shows the cross sections for the coatings. There seems to be no difference between them in appearance. The elastic constants determinations by the SAW spectroscopy are performed, based on the SAW dispersion measurement and the inverse analysis of the measured data. The key to the accurate inverse analysis lies in the quality of the measured dispersion curve. In this work, an ultrasonic micro-spectrometer (UMSM) which provides a superior SAW velocity measurement with a wide frequency range as well as a high accuracy was used [4, 11]. Figure 4 shows a block diagram of the UMSM. In the measurement using the UMSM broadband impulsive waves at incidence angles from 22 ~ to 40 ~ are injected from the spherical acoustic lens to the specimen surface through water and the reflected waves are received by the planar transducer. Applying Fourier transform analysis to the measured ultrasonic signals, the ultrasonic reflection coefficient at the water-specimen interface is measured as a function of both the frequency and the incidence angle. The frequency range is from 10 to 140 MHz. The SAW dispersion is then obtained by analyzing the phase data of the reflection coefficients [4, 11]. The relative accuracy of the measured velocity is about 0.03 % and it corresponds to the velocity variation of+ 1.0 rn/s in the measurement.

oscilloscope

-stage

I [

Stage controller

Fig.4. Schematic block diagram of an ultrasonic micro-spectrometer.

I. Ihara, T. Sawa and K. Tanaka

210

The nano-indentation tests were performed using the conventional load controlled nanoindentation tester, ENT-1100 (ELIONIX Co., Tokyo) with the Berkovich indenter (a triangular pyramid indenter). The measurement resolutions of the applied load and the relative displacement of the indenter are 0.2 gN and 0.3 tun, respectively. Since the specimens used here are thin films on a substrate, the measured P-h curve may be influenced by the elastic deformation of the substrate. In this work, to check the influence of the substrate stiffness on the indentation tests, twenty five P-h curves which have different maximum displacements of the indenter (hm~x30-500 nm) were measured by altering the applied load (Pr~x= 0.3-30 mN).

RESULTS AND DISCUSSIONS Figure 5 shows the measured SAW dispersion curves for NiP and SiC/NiP coatings. It is obvious from this result that the two dispersion curves differ from each other. This reveals that the dispersion curve is strongly affected by the SiC particles dispersed in the NiP matrix. We can safely say that the dispersion curve for the SiC/NiP coating reflects the average properties of the mixture consisting of the SiC particles and the NiP matrix, since the SAW wavelength (about 3 0 - 100/ma) is much larger than the coating thickness (10pro) as well as the SiC particles size (about 0.2/an). The inverse analyses are then applied to each dispersion curve to determine the elastic constants and the density of the coating. The constraints which are upper and lower limits of the range to be searched in the inversion process are within +30% of the values given in a literatures or the theoretical values by a mixture rule: 100 < E < 240 GPa, 0.2 < v < 0.45, 4300 < p < 10000 kg/m3 for SiC/NiP coating, and 40 < E < 150 GPa, 0.15 < v < 0.44, 3950 < p < 8700 kg/m3 for NiP coating. The obtained results from ten times inversions are listed in Table 1. We can see from this result that the Young's modulus for SiC/NiP coating is higher than that for NiP coating by 32%. This enhancement of the Young's modulus is likely due to the distribution of the SiC particles whose Young's modulus (about 410 GPa) is quite higher than that for the NiP. Because of the small density of the SiC (about 3200 kg/m3), the result that the density for SiC/NiP coating is lower than that for NiP coating by 7% is also reasonable. Thus, we can say that the estimated values by the SAW are qualitatively valid. In nano-indentation tests, twenty five P-h curves that have different maximum penetration depths were measured for each coming by altering the applied load. Figure 6 shows some of the measured P-h curves for the two coatings. It can be clearly seen from the result that the 3200

//SiC/Ni-P coating

3000

..~ o

m.,

2800 2600 Ni-P coating

u

2400

r~

2200

20

40 60 80 Frequency (1~-~)

Fig.5. Measured dispersion curves for NiP and SiC/NiP coatings.

100

Evaluation of Composite Coating

211

Table 1.95 % confidence limits of the determined elastic constants and densities for coatings by surface acoustic waves.

NiP coating SiC/NiP coating

Young's modulus (GPa) 117.0 + 5.1 154.8 + 11.2

Poisson's ratio 0.41 + 0.02 0.31 + 0.03

Density (k~/m3) 8274.2 + 263.1 7676.6 ~: 384.3

inclination of the curves for SiC/NiP coating is steeper than that for NiP coating. This suggests that the elastic modulus of the SiC/NIP is higher than that of the NiP. The effective Young's moduli E* are calculated from the twenty five P-h curves for each coating using Eq. (2) and the 95% confidence limits are then obtained to be 125.8+17.7 GPa for the NIP and 152.0~28.8 GPa for the SiC/NIP. As expected, the SiC/NiP coating has a higher Young's modulus. It seemed in the present nano-indentation tests that there was little influence of the rigidity of the steel substrate since the estimated E*s were almost independent of the hn~x. However, it should be noted that the scatter in the measured values for the SiC/NIP coating has a tendency to increase while the hn~x is less than 100 nm. This is likely due to the inhomogeneous structure that is a mixture of SiC and NIP. In order to compare the estimated results by the nano-indentation to those by the SAW spectroscopy, the Young's moduli for the NIP and the SiC/NIP coatings are calculated using Eq.(3) with E1 = 1050 GPa, ~ = 0.1, and vs = 0.41 for NiP and vs = 0.31 for SiC/NIP, respectively. The obtained results are summarized in Table 2 together with the other results estimated by ultrasonically and theoretically. Although there are several kinds of methods for theoretically estimating the elastic constants of an inhomogeneous material, we use here the self-consistent estimate which is a well-known method to calculate the overall elastic constants of an elastic solid containing randomly distributed micro-inclusions, assuming that all inclusions are spherical and the matrix and the inclusions are both linearly elastic and isotropic [12]. It is also assumed in the estimation that the inclusion content is 15% by volume and the Young's modulus and Poisson's ratio are 117 GPa and 0.41 for the NIP matrix and 410 GPa and 0.15 for the SiC particle, respectively. We can see from the results in Table 2 that the measured values by the different techniques, SAW and nano-indentation, almost agree within the experimental errors, furthermore, the two measured values for the SiC/NIP coating are close to the theoretically estimated one. The relatively small

35

~

' , ' , .

30

,

i

. . . .

i

. . . .

i

. . . .

. . . . . siC/NiP coating .......... NiP coa

~

. . . .

,

.

,.

//

400

soo

.

.

.

25

o

<

5 0

o

loo

200

300

,

,

,

600

Penetration depth h (urn)

Fig. 6. Measured load-penetration depth curves for coatings. Different lines for each coating are for different maximum applied loads Pn,~x.

212

1. lhara, T. Sawa and K. Tanaka

Table 2. Comparison among the determined Young's moduli by ultrasonically, mechanically and theoretically. (GPa) Surface acoustic wave Nano-indentation Theoretical NiP coating 117.0+ 5.1 118.8+ 18.9 SiC/NiP coating 154.8 + 11.2 160.6 + 35.0 142.3

value for the theoretical one may arise from the uncertainties of the assumptions used in the estimation. Nevertheless, the fact that the two experimental results almost agree with each other confirms the validity of the measured values and of the measurement techniques employed. Based on this confn'mation, we can reasonably say from the results by the SAW estimation that owing to the dispersed SiC particles, the Young's modulus of the SiC/NiP coating is higher than that of the NiP coating by 32 %. CONCLUSIONS In this work the comparative study of SAW spectroscopy and nano-indentation on the estimation of coatings has been made to examine the SAW technique for its validity. In the experiments with SiC/NiP and NiP coatings, the ultrasonically determined Young's moduli agree with the mechanically determined ones. Furthermore, the determined values for the SiC/NiP are close to the theoretically estimated one. It has been confirmed from the results that the elastic constants determined by the SAW are quantitatively valid. The results of the SAW estimations have also demonstrated that the Young's modulus of the SiC/NiP is 32 % higher than that of the NiP. The results of this work show that the two techniques employed here can be applied to evaluation for other technologically important composite coatings. ACKNOWLEDGMENT The authors would like to thank Nihon Kanizen Co. for supplying the specimens. REFERENCES 1. Briggs, A. (1992). Acoustic Microscopy. Clarendon Press, Oxford. 2. Achenbach, J. D., Kim, J. O. and Lee, Y.-C. (1995). In: Advances in Acoustic Microscopy 1, pp. 153-208, Briggs, A. (Ed). Plenum, New York. 3. Kundu, T. (1992) J. Acoust. Soc. Am. 91, 591. 4. Ihara, I. and Koguchi, H. (1995). In: Simulation of Materials Processing: NUMIFORM95, pp. 575-580, Shen, S.-E and Dawson, P. (Eds). A.A.Balkema, Rotterdam. 5. Makarov, S., Chilla, E. and Frohlich, H.-J. (1995) J. Appl. Phys. 78, 5028. 6. Box, M. J. (1965) Computer Journal 8, 42. 7. Ihara, I., Koguchi, H., Uchida, N., Aizawa, T. and Kihara, J. (1997). Proc. IEEE Ultrason. Syrup. (Toronto), pp. 545-548. 8. King, R. B. (1987) Int.J.Solids Structures, 23,1657. 9. Oliver, W. C. and Pharr, G. M. (1992) J. Mater. Res., 7-6, 1564. 10. Sawa, T., Akiyama, Y., Shimamoto, A. and Tanaka, K. (1999) J. Mater. Res. 14,6, 2228. 11. Nakaso, N., Ohira, K., Yanaka, M. and Tsukahara, Y. (1994) IEEE Trans. Ultrason. Ferroelect. Freq. Contr., 41-4, 494. 12.Nemat-Nasser, S. and Hod, M. (1993). Micromechanics: overall properties of heterogeneous materials, pp.232, North-Holland, AmsterdarrL

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

213

DETECTION OF THE NANO-SCALE SUBSURFACE DEFECTS

BY ULTRASONIC ATOMIC FORCE M I C R O S C O P Y Toshihiro TSUJI and Kazushi YAMANAKA Department of Material Processing, The University of Tohoku Aoba-yama 02, Sendai 980-8579, Japan ABSTRACT In electric devises, we have to evaluate and detect subsurface defects on the nano-scale. For this purpose, we think that ultrasonic atomic force microscopy (UAFM) is useful because of the possibility of the quantitative evaluation of the elastic property on the nano-scale.

In

this study as a preliminary experiment, we applied UAFM to subsurface lattice defects of highly oriented pyrolytic graphite (HOPG). dislocations of HOPG.

Then we were able to image subsurface

We found that the load dependence of the resonance frequency

measured on a dislocation had a local minimum at a specific contact load and were able to explain it by a reversible lateral displacement of the dislocation.

From these results, we

conclude that UAFM is useful for analysis of subsurface defects on the nano-scale. KEYWORDS Ultrasonic Atomic Force Microscopy, Nano-scale, Subsurface Defects, Resonance Frequency, Contact Stiffness, HOPG, Dislocation, Reversible Displacement INTRODUCTION In present electric devises, we have serious problems on damaged subsurface layers on the nano-scale and have to detect and evaluate them.

We had developed ultrasonic force

microscopy (UFM) and observed subsurface dislocation of HOPG.

But the quantitative

analysis was difficult since the contact between tip and sample showed complicated nonlinearity [ 1].

Then we developed ultrasonic atomic force microscopy (UAFM) in which the

contact can be kept linear [2].

UAFM can measure not only the topography on the nano-

scale but also the quantitative effective elasticity of samples by an inverse analysis taking into account the tip shape and the adhesion force between the tip and sample [3]. In this study, we applied UAFM to cleaved surface of HOPG and tried to quantitatively analyze dislocations.

As a result, we made a new discovery on dislocations.

214

T. Tsuji and K. Yamanaka

EXPERIMENT Fig.1 shows a schematic illustration of UAFM. We have developed UAFM attaching a high frequency vibrator to the cantilever support of a contact mode AFM. We measure the deflection of a cantilever using a laser beam detected by the two segment photodiode. When we drive the vibrator by a swept frequency signal of a network analyzer, we process the detected signal by the network analyzer and obtain the cantilever vibration spectra. In addition, when we drive the z axis by a DC signal of an external synthesizer, we can precisely measure the load dependence of the resonance spectra. On the other hand, when we drive the vibrator by a constant frequency signal, we process the detected signal by a lock-in amplifier. Then we scan the sample in the x-y direction and we obtain the amplitude distribution of a cantilever vibration. At the same time, we control the height of the sample by the detected signal lower than 1 kHz to measure the topography. We used highly oriented pyrolytic graphite (HOPG) as a sample. Fig.2 shows a schematic illustration of cleaved surface of HOPG. It has layered structure of atomic planes composed by connected tings of six carbon atoms. If we cleave HOPG, we observe steps with a height in the range from 0.3nm to about 200nm and atomically fiat terraces between the steps. We applied UAFM to terraces and carried out the experiment in dry air with the temperature of 23~ and the relative humidity of 30%.

Cantilever Deflection Vibration Signal (Linear Detection) LPF~ Photodiode ~..~:..

shaped s ~

]

~ Laser .Ultrasonic <~ii!~> z ~ Vibrator

II

a ~

[

~ ~

"~

r

Swept frequency Signal~

Network Analyzer (NWA)

/ Constant frequency( r Signal

[Z-Axis Controller

Lock-in Amplifier

HighFrequency Generator

I Ref.

TXYZ-Axis Control DC signal

Piezo electric scanner for sample

ISyn, [[Topography 11~

Fig. 1 An implementation of ultrasonic atomic force microscope.

(

~

J

215

Detection of Subsurface Defects by UAFM

Fig2 A schematic illustration of cleaved surface of HOPG. ANALYSIS OF THE DISLOCATION OF HOPG BY UAFM Fig.3 shows the cleaved surface of HOPG observed by UAFM. 500nm.

The image area is 500nm x

The contact load calculated from the deflection of the cantilever was 0nN, with an

adhesion force of about-60nN.

Fig.3(a) is the contact mode AFM topography.

scale shows that the maximum height is 10.4nm. about 10nm and atomically flat terraces.

The gray

We observed a step with the height of

We observed the step with a width since the side

wall of the tip gradually ran over the step.

Fig.3(b) is the amplitude distribution of the

cantilever deflection vibration at the first resonance frequency measured at the position A (118.4kHz).

We observed dark lines at the position B, C and D.

Then, to evaluate the elastic property at these positions, we measured the load dependence of the cantilever vibration spectra.

Fig.4 shows the load dependence of the cantilever

deflection vibration spectra measured at the position A and B.

When the contact load was in

the range from 0nN to -25nN, the spectrum was shifted to lower frequency. contact stiffness at the position B was smaller than at the position A. known that HOPG has edge dislocations such as shown in Fig.5.

Therefore the

On the other hand, it is

Since the distance between

atomic planes at the position of the dislocation is longer than that at other positions, the elasticity and the contact stiffness at the position of the dislocation is smaller than that at other positions.

Therefore, we estimate that dark lines observed in Fig.3(b) were subsurface edge

dislocations of HOPG. Moreover, we estimate that the positions B and D were on a same dislocation since they were on the same straight line. about 10nm.

In addition, the left terrace was higher than the right one by

Therefore, in this image, we estimate that we were able to observe a subsurface

dislocation at the depth of more than 10nm.

216

T. Tsuji and K. Yamanaka

Fig.3 Images of cleaved surface of HOPG. (a) Contact AFM topographic image. (b) Amplitude distribution of cantilever deflection vibration at the resonance frequency measured at position A (118.4kHz). Image area is 500rim x 500rim and contact load was 0nN.

0

--5O ~

on

.lo o rn

,9

-100

...,,

t~

O

-150 115

120

125

Frequency, kHz

Fig.4 The load dependence of the cantilever deflection vibration spectra.

Detection of Subsurface Defects by UAFM cleaved surface

B

atomic plane of HOPG

tip

217

A

edge dislocation

Fig.5 A schematic illustration of an edge dislocation of HOPG. Fig.6 shows the load dependence of the resonance frequency measured from the spectrum in Fig.4.

The horizontal axis is the contact load and the vertical axis is the resonance frequency.

The resonance frequency measured at the position A monotonically increased as the contact load increased. To explain this load dependence, we carried out an inverse analysis of the resonance frequency.

In this analysis, we assumed that both the tip and the sample are homogeneous

elastic bodies, the shape of the tip is z = cr", where n is the tip shape index, and took into account the adhesion force between the tip and the sample [3]. Then, we simultaneously estimated the parameters such as the thickness of the cantilever t, the tip shape index n and the adhesion force Fc by assuming the Young's modulus and the Poisson's ratio of HOPG,

Ei~opa=3OGPa and VHopc=0.24, and those of the tip of single crystal Si, Esi=166GPa and Vsi =0.22.

As a result, we were able to fit the calculated resonance frequencies to measured

ones with the maximum error of about 0.13kHz.

The estimated values of t, n and Fc are

shown in Fig.6. Therefore we were able to explain the load dependence of the resonance frequency measured at the position A by the contact between two elastic bodies.

But the resonance frequency at

the position B showed a unique behavior and had a local minimum at a specific load and then rapidly increased to the value of resonance frequency measured at the position A with no dislocation. To explain this behavior, we make a finite element analysis in the next section.

218

T. Tsuji and K. Yamanaka 119.0~- ....

' .........

L 118.5[-

118.0 ~-

, .y~."

L

~,

' .........

Measured frequency 9A

I-

L k

117 51-

. ,~,.~j~o'~o " ii 9 . .-~Ta~Z~ ~'o ~

d.r~,, o o ,?~ ",fl" ~176176-

. ~

o -

" I- ,-~-" L ,.~[_ ~ Calculated ~ 117-07 frequency 8

' ...... _

8

effective

elasticity

(S'_~-g_r_aphite_) i.-6-~'8~a~-Ifia

thicknes s

of cantilever t=4.230/z m tip shape index n=3152 pull off force

-

-

-

_i

Fc=-96.67nN

116.5 . . . .

i ......... -50

i ......... 0 Load F,nN Fig.6 The load dependence of the resonance frequency.

i ...... 50

FINITE ELEMENT ANALYSIS OF SUBSURFACE DEFECTS Fig.7 shows the cross-section of models for the finite element analysis along the indent axis. In the model Fig.7(a), we assume the low elasticity layer which represents the dislocation at the depth of 3nm.

In the model (b), to take into the effect of the finite width of the

dislocation, we limited the radius of the low elasticity layer to 3nm from the indentation axis. The Young's moduli and Poisson's ratios of the substrate and the low elasticity layer are E,b=30GPa, v,b=0.24, E~w=15GPa and v~=0.24. We carried out the finite element analysis of the model (a), (b) and another model with the homogeneous elasticity of the Young's modulus E =30GPa and the Poisson's ratio v=0.24 to indent a rigid sphere of 50nm radius to the depth of lnm. Fig.8 shows the theoretical load dependence of the resonance frequency at the first deflectional mode calculated from the finite element analysis.

We calculated the resonance

frequency by the parameter of the cantilever used in previous section.

The solid curve

shows the resonance frequency of the model with the homogeneous elasticity of the Young's modulus 30GPa.

In the case of the model Fig.8(a) shown by the chained curve, we

reproduced the decrease of the resonance frequency at low load but were not able to reproduce the approach of the resonance frequency to the value of the homogeneous model at high load.

In the case of the model Fig.8(b), taking into account the effect of the width of

the dislocation shown by the doted curve, we were able to reproduce the approach of the resonance frequency when the load was higher than 100nN, but not able to reproduce the local minimum of the resonance frequency at a specific load.

Therefore, we concluded that

Detection of Subsurface Defects by UAFM

219

we are able to partly explain the unique behavior of the resonance frequency but not perfectly by the finite element analysis. In order to improve the explanation of the observed phenomena, we tried more detailed observation of the images of dislocations. Then, we found that the dislocation moved laterally by a few tens of nm as the load increased and the dislocation returned to the original position as the load decreased. Therefore, we are able to explain the load dependence of the resonance frequency by the reversible displacement of the dislocation, which will be published elsewhere [4].
i ~R i

i !

Rigid

=50 / ~

:!>~R=50/ i

ff ~'/

]

~

sphere

E=15GPai v=0"24 i

~

//

U

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

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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

I

I

I

I

I

I

I

'

'

I

/.,'!/ IJ;

108 ///

I

I

I

I

I

-

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/.,/// /,'//

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.

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-:

,," i"

//./ g~= 111 /////

1070

I

/

~'

n,

I

elastic body E=30GPa, z/=0.24

,--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"

<

0

i

u

l

0 Frequency [MHz]

Fig.5 Frequency dependence of attenuation coefficients of epoxy resin and CFRP.

the matrix. In such situations where the reinforcing phases are so large, the above defined normalized frequency grows to an order of 10-1. Although direct comparison is refrained from because of different reinforcement geometry, the experimental results of Kinra et al. [8] qualitatively supports the present discussion. In order to assess the present theory from an experimental point of view, ultrasonic attenuation measurements were carried out for a unidirectional CFRP plate consisting of epoxy resin reinforced by carbon fibers with fiber volume fraction 55%. Longitudinal ultrasonic pulses were transmitted into the CFRP plate via a delay-line medium (PMMA) in the propagation direction normal to the fibers. Together with the diffraction loss, the reflection/transmission response at the transducer front surface was correlated [11]. A transducer for longitudinal wave with the center frequency 5 MHz was used. Fast Fourier Transform technique was employed to analyze spectral response of reflected pulses to compute frequency-dependent attenuation of the CFRP. Figure 5 shows the measured attenuation coefficient of the longitudinal wave in the CFRP as open circles, along with the data for epoxy resin as triangles with their fitting line (broken). It can be observed that not only in the epoxy resin but also in the CFRP the attenuation coefficient shows almost linear frequency dependence. Also, it can be confirmed that in the measured frequency range the composite attenuation is indeed less than that of the matrix. In Fig.5, the theoretical prediction of the attenuation in the CFRP is plotted in the case of r 0.55 as a solid curve, with the matrix absorption ( 1 - r shown as a chained line. The frequency range of the measurement (2 to 8 MHz) corresponds by the above normalization roughly to ( a / c , ~ ) f <0.01. In Fig.5 it is seen that the model prediction conforms to the measured result regarding its magnitude relative to the matrix attenuation as well as its frequency dependence. Moreover, it is observed that in the illustrated frequency range the present theoretical model predicts relatively small scattering effects, indicating that the major governing factor of the attenuation in CFRP is the absorption in the matrix. It is remarked finally though that for such high fiber volume fractions discussed here, due to the aforementioned reason no sound justification is at hand to employing the simplistic independent scattering model. When the scattering is relatively weak, however, the present model is thought to give a reasonable result. Extension of the model to incorporate multiple scattering effects in general circumstances constitutes a subject of the future investigation.

230

S. Biwa, Y. Watanabe and N. Ohno

CONCLUSION In the present investigation, a theoretical model of ultrasonic attenuation behavior in polymer-based fiber reinforced composites has been formulated accounting for both scattering by the fibers and absorption in the viscoelastic matrix. According to the present model employing the notion of independent scattering, the composite attenuation coefficient can be expressed in a simple form using the matrix attenuation, the fiber volume fraction and the scattering cross section of the fiber. Based on the derived model, attenuation of the longitudinal wave in a unidirectional CFRP has been analyzed. As a result of the analysis, it has been shown that depending on the magnitude of the normalized frequency (alternatively, the ratio between the fiber radius and the incident wavelength) the composite attenuation may or may not exceed the matrix attenuation. The model also illustrates that in the frequency range of practical interest the major factor of the attenuation in CFRP is the viscoelastic absorption in the matrix, so that the composite, when reinforced by elastic fibers, yields lower attenuation than the matrix. The frequency dependence of the predicted attenuation coefficient of CFRP has been verified by the corresponding measurements.

ACKNOWLEDGEMENTS The authors wish to express their gratitude to Mitsubishi Rayon Co., Ltd., Products Development Laboratories for supplying CFRP and epoxy samples for the attenuation measurements. It is also acknowledged that one of the authors (S. B.) received a financial support from the Ministry of Education, Science, Sports and Culture, Japan, under the Grant-in-Aid for Encouragement of Young Scientists.

REFERENCES 1.

Lhermitte, T. D., Handley, S. M., Holland, M. R., Miller, J. G. and De Mol, R. (1992), In: Nondestructive Testing '92, Hallai, C. and Kulcsar, P. (Eds), pp.924-928, Elsevier Sci. Pub., Amsterdam. 2. Hosten, B. (1998), J. Acoust. Soc. Am. 104, 1382. 3. Jeong, H. (1997), J. Compos. Mater 31, 276. 4. Papadakis, E. P. (1968), In: Physical Acoustics IV, Part B, pp.269-328, Mason, W. P. (Ed), Academic Press, New York. 5. Bose, S. K. and Mal, A. K. (1974), J. Mech. Phys. Solids 22, 217. 6. Truell, R., Elbaum, C. and Chick, B. B. (1969), Ultrasonic Methods in Solid State Physics, Academic Press, New York. Brauner, N. and Beltzer, A. I. (1988), Ultrasonics 26, 328. 8. Kinra, V. K., Petraitis, M. S. and Datta, S. K. (1980), Int. J. Solids Structures 16, 301. 9. Datta, S. K., Ledbetter, H. M. and Kriz, R. D. (1984), Int. J. Solids Structures 20, 429. 10. Pao, Y.-H. and Mow, C.-C. (1971), Diffraction of Elastic Waves and Dynamic Stress Concentrations, Rand Corp., Crane, Russak & Company Inc., New York. 11. Papadakis, E. P., Fowler, K. A. and Lynnworth, L. C. (1973), J. Acoust. Soc. Am. 53, 1336. ,

Nondestructive Characterization of Materials X

Green et al. (Eds) 9 2001 ElsevierScience Ltd. All rights reserved.

231

ATTENUATION INDUCED FROM DISTRIBUTED DEFECTS IN A N E L A S T I C S O L I D

M. Kitahara, D. Kishibe, T. Takahashi and N. Kishi Dept. of Civil Engineering, Graduate School of Engineering, Tohoku University Aoba-yama 06, Sendai 980-8579, Japan

ABSTRACT Attenuation caused by the scattering is studied for the distributed cracks in an elastic material. The attenuation is evaluated from the scattering cross-section by giving the number of defects in the unit volume, provided that the distribution of cracks is dilute. Here, the scattering cross-section is calculated from the far-field representation of the scattered waves by cracks. The cement-based specimen that contains penny shaped cracks is prepared and the measurement of scattered waveforms is performed. After the processing of measured waveforms, the measured attenuation is obtained and it is compared with the numerically evaluated attenuation. A relatively good agreement is obtained for the measured and numerically evaluated attenuations in the effective range of transducer bandwidth. KEYWORDS Attenuation, scattering cross-section, penny shaped cracks, cement-based material. INTRODUCTION The ultrasonic attenuation can be used to characterize the distributed defects in materials and the comprehensive review of the ultrasonic attenuation has been given by Beltzer [1]. Experimental observations of the attenuation for ultrasonic waves provide useful information of the damaged state in the material. Often, an increase in the attenuation of ultrasonic waves is an indication of stiffness degradation of the materials [2]. In this paper, a criterion of the dilute distribution for cracks is thoroughly studied by the multiple scattering calculations for cracks and the criteria is used to prepare the specimen which contains penny shaped cracks. The ultrasonic attenuation induced from the scattering by dilutely distributed cracks is experimentally observed and the observed attenuation is compared with the numerically calculated attenuation. An agreement of the observed and calculated attenuations is discussed in the effective frequency bandwidth of the transducer. It is to be remarked that the multiple scattering effect is negligible order in the present setting of the experiment. As for the method to calculate the multiple scattering effect of cracks, we refer to the reference [3].

M. Kitahara et al.

232

The ultrasonic attenuation in an epoxy matrix containing lead inclusions has been reported by Sayers and Smith [4]. Recently, the laser ultrasonics [5] and the electromagnetic acoustics [6] have been advantageously applied to evaluate the attenuation of the material. ATTENUATION

FOR MATERIAL CONTAINING

CRACKS

We consider the elastic material that contains the distributed cracks as shown in Fig.1. Here the incident wave u is propagating along the x3-coordinate axis. The following assumptions are made for the distributed cracks. (1) The distribution of cracks is dilute so that the multiple scattering effect is approximately neglected. (2) The shape of cracks is all penny shaped with same radius a. (3) The normal vectors of all cracks point to the x3-coordinate direction. (4) The distribution of cracks is random.

Fig.1. Elastic material containing dilutely distributed penny shaped cracks.

For the investigation of the wave fields in Fig.l, it is convenient to introduce the complex wave number [1]

kL(w) -cO/CL(W)+ ia(a~)

(1)

where CL(CO)is the phase velocity of the longitudinal wave and a(w) is the attenuation. Then the plane wave propagating along the x3-axis can be written in the following form.

(2) From Eq.(2), the intensity < I > averaged in time reduces to < I >=<

Io > e - 2 ~

(3)

where the bracket < 9 > means the time average and < I0 > is the intensity at the reference state x3 = 0. The time averaged power of the scattered wave < ps > can be written

Attenuation Induced f r o m D&tributed Defects

233

< pS > = P(w) < I >

(4)

where P(co) is the total scattering cross-section [7]. When the distribution of cracks is dilute, the multiple scattering effect of cracks becomes small and the independent scatterer approximation can be used. Then the differential equation governing the change of the intensity < I > in the Xa-coordinate direction can be written as d

+ NP(~) < I > = 0

dx3

(5)

where N is the number of cracks in the unit volume. The solution of Eq.(5) is in a simple form. < I > = < I0 > e -NPx3

(6)

The attenuation a(w) is obtained

(7)

o~(w) = ~ N P ( w )

from Eqs.(3) and (6). We define the nondimensional parameter e as (8)

e - Na 3

where a is the radius of penny shaped cracks. The nondimensional attenuation coefficient o~(w)a can be written a(w)a = ~_ P(w) 2 7ra2

(9)

from Eqs.(7) and (8). SCATTERING

CROSS-SECTION

FOR CRACKS

For the evaluation of the attenuation, it is important to calculate the total scattering cross-section P(w) in Eq.(7) or Eq.(9). A general expression of the total scattering crosssection is considered for the elastic material with cracks. For the plane longitudinal incident wave with the propagation vector p uiIN -- Pi e x p ( i k L p " X),

(]p] -- 1)

(10)

the time averaged intensity < I > of the incident wave reduces to the following form. 1 < I > = ---~wkL()~ + 2#)

(11)

The expression for the differential cross-section d P ( w ) / d ~ (where d~ is the solid angle) becomes

M. Kitahara et al.

234

2 ^ SC SC, dP(~z) = lim x Im(xioij Uj )

(12)

where & is the unit vector pointing to the observation point x and x - Ix. The scattered wave u~ C in Eq.(12) can be rewritten as

~(~)=

~: A~s exp(iksx) s=L,T,H

(13)

X

where A s is the scattering amplitude of the longitudinal (c~ = L) and transverse (a = T, H) wave components and it has the form

A~ = iksf~f~I~(Sc)

(14)

where f s is the polarization vector of the c~(= L, T, H) wave component. The term In~(.) is the integration of the crack opening displacement A u over the crack surface Sc I:(5~) =

CkjnlX'l /s e-ik~$.y 4~pc~ c nk(y)Auj(y) dSy

(15)

where n is the unit normal vector of the crack surface and y is the position vector on the crack surface Sc. Introduction of Eq.(13) to Eq.(12) leads to the following expression for the differential cross-section

dP d~

kL ATI2 = IALl2 + ~T(I +l A H I2)

(16)

where A s is the scattering amplitude in Eqs.(14) and (15) for the a-wave component. The total scattering cross-section P(w) can be obtained from the integration of the differential scattering cross-section with respect to the solid angle on the unit sphere. i ' r dP P(w) = ] ] ~-~ sin OdOdr

(17)

The integration in Eq.(17) can be carried out numerically. APPLICABILITY

OF INDEPENDENT

SCATTERER APPROXIMATIONS

Approximate criteria to adopt the independent scatterer theory are examined by performing the multiple scattering calculations for cracks. In the calculations, the horizontal and vertical arrangements are considered for the configuration of two circular cracks. The radius of cracks is a and the distance d of two cracks is chosen as a parameter to determine the criteria. Fig.2 shows the horizontal arrangement of two circular cracks. The incident longitudinal plane wave is propagating along the x3-coordinate direction. Figs.3~5 show the normalized scattering cross-sections for two horizontal cracks in the case of d/a = 0.5, 1 and 2. The normalized total scattering cross-section P(~z)/Tra 2 for two horizontal cracks is shown

Attenuation Induced from Distributed Defects

235

by the square [] mark. For the reference, two times the value of the total scattering crosssection for a crack is depicted by the dot 9 mark. The white circles and triangles show the contributions from the longitudinal and transverse waves, respectively. In the case of the horizontal arrangement, the multiple scattering effect is small for the total scattering cross-sections, even if the distance d/a is 0.5, as shown in Fig.3.

Fig.2. Two horizontal cracks. Fig.3. Scattering cross-section.

Fig.4. Scattering cross-section.

Fig.5. Scattering cross-section.

Fig.6 shows the vertical arrangement of two circular cracks. Figs.7~9 show the scattering cross-sections for vertical cracks in the case of d/a = 1, 4 and 5. For the vertical arrangement, the multiple scattering effect is large in the distance of d/a = 1. At about two times the distance of the crack diameter, d/a = 4, the multiple scattering effect becomes relatively small in the low frequency range of akL= 0 ~ 2, as shown in Fig.8.

236

M. Kitahara et al.

Fig.6. Two vertical cracks. Fig.7. Scattering cross-section.

Fig.8. Scattering cross-section. ATTENUATION

Fig.9. Scattering cross-section.

MEASUREMENT

The experimental set-up to measure the attenuation is shown in Fig.10. The square pulse is sent to the piezoelectric transducer. The transducer radiates the incident ultrasonic wave to the specimen and the forward-scattered wave is received at another transducer. The signal is sent to the receiver and it is stored in the digital oscilloscope. The oscilloscope

Fig.10. Measurement system.

Attenuation Induced from Distributed Defects

237

is connected to PC and the data processing is done in the personal computer. The cement-based specimen shown in Fig.11 is prepared for the ultrasonic measurement. A part of specimen which has no crack is used to measure the reference waveform u0.

~0(=) = ~(x~, z~)~-:0(~/:~i(~/~0(~//~-~
(is)

The forward scattering waveform u l from the cracks U 1 (X) : ~(Xl,

X2)e--Ctl(W)X3ci[(W/cLl(W))x3-Wt]

(19)

is measured in the part of the specimen that contains the cracks. Taking the absolute values of Eqs.(18) and (19) and the natural logarithm of the ratio of these two absolute values leads to the expression for the attenuation.

~(w)a

-

-

{O~I(W ) -- ao(w)}a -- a

[u0(w)l dlnlul(w)l

(20)

In this expression, the thickness d of the specimen is used for the propagation distance in the x3-coordinate direction and the attenuation a(w) is nondimensionalized by the crack radius a.

Fig.11. Preparation of cement-based specimen with cracks.

MEASURED AND CALCULATED ATTENUATIONS The measured amplitudes in the frequency range are shown in Fig.12 for both of the reference and forward scattering waves. The measured attenuation evaluated from Eq.(20) is plotted by triangles in Fig.13. The numerically calculated attenuation from Eq.(9) is also shown by the solid line for comparison. For both of the experiment and the numerical calculation, the nondimensional parameter for the average crack numbers is chosen as - N a 3 = 0.0078. The nominal center frequency of the transducer is 0.5MHz. Relatively good agreement is obtained in the effective frequency range of the transducer.

M. Kitahara et al.

238

- - ~ measurement numerical ~alysis 1'

o. 04O. 45-.

04o

o.35-:. O 0. 30-:

0.10-i O. 05-i

o.oo_i~~i~ 0.o

o.1

0.2

I

0.3

0.4

0.5

reference

~

scatteredwave

wav

r .~

LZ,I_I-

0.6

o.7

:

O. 025 O. 02 0.015 0.01

O. 005

0.8

,,,,,,,,

0.9

1.o

Frequency (MHz) Fig.12. Amplitudes in frequency range.

'] " "

0.03

, ,.,.~

d j7

O. 15-:

O. 035

~

0

0

_J 0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

Frequency (MHz) Fig.13. Measured and calculated attenuations for cement-based material with cracks.

ACKNOWLEDGEMENT A part of this work was supported by the Grant-in Aid for COE Research(No.liCE2003), The Ministry of Education, Science, Sports and Culture. R E F E R E N C E S

1. Beltzer, A.I. (1989) Wave Motion 11, 211. 2. Zhang, Ch. and Achenbach, J.D. (1991) Int. J. Solids Structures 27, 751. 3. Kitahara, M., Kishi, N. and Nakagawa, K. (1999). In:Review of Progress in Quantitative Nondestructive Evaluation 18, pp.37-44, Thompson, D.O. and Chimenti, D.E.(Eds). Kluwer Academic/Plenum Publishers, New York. 4. Sayers, C.M. and Smith, R.L. (1983) J. Phys. D: AppI. Phys. 16, 1189. 5. Cho, H., Sato, H., Takemoto, M., Sato, A. and Yamanaka, K. (1996) Jpn. J. Appl. Phys. 35, 3062. 6. Hirao, M. and Ogi, H. (1997) Ultrasonics 35, 413. 7. Gubernatis, J.E., Domany, E. and Krumhansl, J.A. (1977) J. Appl. Phys. 48, 2804.

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

239

FDM ANALYSIS OF ULTRASONIC NONLINEARITY IN PARTIALLY DEGRADED MATERIAL

K. C. Kim, H. Yamawaki and T. Saito

National Research Institute for Metals, Sengen 1-2-1, Tsukuba, Ibaraki 305-0047, Japan K. Y. Jhang

Hanyang University, School of Mechanical Engineering, 17 Haengdang-dong, Seongdong-ku, Seou1133-791, Korea

ABSTRACT In recent, the fact that the nonlinear acoustic effect has a strong correlation to the micro deformation of materials has been considered as useful for NDE of material degradation. However, the most past researches considered only the case that the material degradation was distributed evenly along the path of the probing wave. Contrarily, in this paper, the nonlinear behavior of ultrasonic wave in partially degraded material is considered. For this aim, FDM(finite difference method) model for the nonlinear wave equation was developed with the restriction to the 1-D longitudinal wave motion and how the partial degradation in material contributes to the measurable nonlinear parameter was analyzed quantitatively. In order to verify the rightness of this simulation method, the relation between the measurable nonlinear parameter and the continuous distribution of degradation obtained from simulation was compared with experiment results and two results showed similar tendency. It can be known from simulation result that the degradation level, the range of degradation and the continuous distribution of degradation have strong correlation with the measurable nonlinear parameter. As it was possible in these simulations that only special part is assumed as degraded one, the quantitative evaluation of partially degraded material will be obtained by using this method. KEYWORDS Nonlinear acoustic effect, material degradation, FDM, measurable nonlinear parameter.

INTRODUCTION In recent, several theoretical and experimental researches have reported that the nonlinear acoustic effect of the ultrasonic wave has a strong correlation to the micro deformation of materials.[1-7] Where, the acoustic nonlinearity is considered as a kind of material constant so that it changes according to the change of material properties. Therefore, if it can be measured then the change of material can be evaluated. Parameter [3 obtained by nonlinear elasticity is usually used as such quantitative evaluation of nonlinear acoustic effect. The value of 13 in the degraded material is different from that in the virgin state. This property is very useful for the NDE of material degradations. However, the most past researches considered only the case that the material degradation was distributed evenly along the path of the probing wave, since the typically used method to measure the parameter [3 can calculate only the value averaged along the wave path. Of course there was an

240

K.C. Kim et al.

effort to localize the acoustic nonlinearity, but it considered only the localization method using the localized pumping wave in addition to the probing wave.[2,8] Contrarily, in this paper, the nonlinear behavior of ultrasonic wave in the partially degraded material is considered. This is more close to the actual situation as like as in the stress concentration or local damage. The analysis was carried out by numerical technique, since the analytical approach is so complicated and difficult to solve in this case. Numerical techniques for the ultrasonic visualization are power methods that provide clear and instant comprehension of the ultrasonic phenomena in solids. Such techniques have been used for the study of elastic wave, and in nondestructive testing, for detailed study of isotropic materials, and partially applied for anisotropic materials[9-11 ]. In this paper, FDM (Finite Difference Method) is applied to show the nonlinear propagation behavior of ultrasonic wave in partially degraded material. Especially, we focus on the 1-D longitudinal ultrasonic wave propagation in partially degraded material. For this aim, the nonlinear wave equation derived on the basis of the nonlinear elasticity is shown and its FDM model is developed. Medium is represented as discrete 1-D nodes, a part of which is assumed to be degraded so that it has different value of nonlinear parameter from that of the other normal nodes. The waveform of ultrasonic wave propagated in partially degraded material was showed and how the partial degradation in material contributes to the measurable nonlinear parameter was analyzed quantitatively by through several simulations. Also, in order to verify this simulation method, the relation between the measurable nonlinear parameter and the continuous distribution of degradation obtained from simulation was compared with previous experiment results[2].

NONLINEAR WAVE EQUATION To begin with, let us consider the equation of motion for ultrasonic wave in an infinite elastic solid in the absence of dissipation, written as 0 2U i

OTt.J

P 0----7 = Ox--~-.

(1)

where p is the mass density, Ui is the displacement in the Xi direction, and Tij is the stress tensor. If only the 1-D motion in thin circular rods is considered, then the above equation simply becomes 02U P cot2

OT Ox

(2)

where the stress T may be written in terms of the strain g using a nonlinear version of Hooke's law as follows[2-6]: T = c~(1 + p~ + ...)

(3)

where c is the stiffness constant and 13 is the nonlinear parameter. Also, strain can be represented by 0u g=~ Ox

(4)

Using Eq. (3), (4) and the fact that the longitudinal wave speed is V0 = x f c / p , Eq. (2) can be rewritten as

Ultrasonic Nonlinearity in Degraded Material 02U Ot2

= v 2

02U Ox2

241

(5)

where, V 2 = Vo2(1 + 2flo~ +...)

(6)

Eq. (5) becomes linear wave equation when only the first term of the right hand side of Eq. (6) is considered.

FDM MODEL Generally, finite difference equation of 1-D linear wave equation can be given as follows, [11]

Ut+ - 2 U t +U t T2

- = Vo2

Ux+ -2Ux +U x_ X2

(7)

where T and X are temporal and spatial increment, respectively. On the other hand, as for the nonlinear wave equation, V0 in Eq. (7) should be replaced with V in Eq. (6). Also, if we replace with the finite differential form of Eq. (4) then we can obtain final FDM model of 1-D nonlinear wave equation as follows.

U,+ - 2U, + U t T2

Ux+ - 2U x + U x (

-=V

) - 1+2fl o UX+2x-Ux_

X2

2

(8)

where the nonlinearity was considered up to first order so that the higher order terms of ~ in Eq. (6) was neglected. 130indicated the nonlinear parameter at the virgin state of material. However, the value of nonlinear parameter will be changed in the degraded material, thus Eq. (8)' is used at the degraded part.

Ut+-2Ut

-

= Vo z

Ux+-2Ux +Ux

-

1 + 2fld

Ux+-Ux ] 2X

-

(8)'

where ~ d shows the nonlinear parameter of degraded material. Fig. 1 shows the 1-D nodal representation of medium with partial degradation. Every normal node is calculated by equation (8) and the degraded nodes are calculated by equation (8)'. x

9 : normal node 0 : degraded node

Fig. 1. The 1-D nodal representation of medium with partial degradation

COMPUTER SIMULATION AND DISCUSSION To prove the applicablity of this method, some calculations were presented here. Genreal numerical conditions are time step At=l.4ps(700 nodes), spatial step L =0.05~tm(200 nodes) and sound velocity 6300m/s; model material is aluminum with p-2680kg/m 3 and initial wave form is sine wave of a period. In order to know the relationship between partial degradation and nonlinearity, 3 kinds of computer simulations are performed as shown in Fig. 2.

242

K.C. Kim et al.

X -- I.

v w w

..... Degraded point" 1 node

(a) The first simulation model : the nonlinear parameter in one node is changed ~d 3 9

A

A

T - ~ ~

I 9

~- ~

Discontinuous distribution o f 13

I,=

w w ~

-

L_v_J Degraded point" N nodes (b) The second simulation model 9the node number calculated by 13d is changed

~d

.... _...O ~

n

u

o

u

s

~

distribution of 13(Gussian)

~o -

"

"

"-

......

0

0

0

0

0

C

,.,

,.,

,.,

"

,.,

-

"

"

Degraded point N nodes (c) The third simulation model 9the distribution ofl3d is assumed as gauss distribution in the degraded nodes

Fig. 2. The explanation of simulation method. (a) The first simulation model 9the nonlinear parameter in one node is changed. (b) The second simulation model 9the node number calculated by 130 is changed. (c) The third simulation model 9the distribution of 130 is assumed as gauss distribution in the degraded nodes. That is, first, by changing the nonlinear parameter of Eq (8)' in one node as like Fig. 2 (a), we tried to increase the degradation level(the degree of degradation). Secondly, by increasing the number of node calculated by 130 as like Fig. 2 (b), we assumed the increase of the range of degradation. Thirdly, by assuming the distribution of 130 as gauss distribution in the degraded nodes as like Fig. 2 (c), we investigate the effect of the continuous distribution of degradation. Before the simulation, the initial and propagated waveform was compared in the case of 130=0 in Eq. (8) and it was confirmed that two cases are exactly same in waveform and nonlinearity. The nonlinearity of propagated waveform is evaluated by the measurable nonlinear parameter [~m which is calculated from the amplitude of received waveform. The measurable nonlinear parameter [~mCan be defined as follows, [1-6] 8A2 i~m = A?k2x

(9)

The calculation method of measurable nonlinear parameter ~m from the amplitude of received waveform was explained in the reference [2]. Where A1 and A2 are the amplitudes of the fundamental frequency component and 2 n~ order harmonic component, k is the wave number, and x is the propagation distance. That is, measurable nonlinear parameter ~m depends on the ratio of the amplitude of the 2 n~ order harmonic component and the power of fundamental component when other variables are constant. Firstly, the nonlinearity of propagated waveform according to the change of nonlinear parameter 130 in Eq (8)' shown in Fig. 2 (a) is calculated and showed in Fig. 3. The value of used 130 was represented Eq (10) ~d = /30(1 +a), (a = 0,0.1,0.2.-.3)

(10)

Where, used 130 is 3.47x 10l~ and represent the nonlinearity of the virgin state of material. In this case, the percentage of degradation node for the whole node is 1%. Here, the degree of degradation is changed by magnitude of nonlinear parameter 13d.

Ultrasonic Nonlinearity in Degraded Material

243

1.5 1 "~

-~ <

0.5 o -0.5 -1 -1.5

t5 I 1. 2 2.5 time(see) x 10 l ~ (a) the wave form used in simulation (solid line : initial wave form, dotted line : the waveform o f ultrasonic wave propagated in partially degraded material) ~ , , ~ , I

0

1.1[

0.5

1

& 1.05l

/ 1

.o__o.~_.o._o_.e._~-~-o--~"~176

i 1.5

i 2

i 2.5

i 3

i 3.5

4

13gl30 (b) The relationship between the degradation level and the normalized measurable nonlinear parameter

Fig. 3. The result of computer simulation; (a) the waveform used in simulation (b) the relationship between the degradation level(13d/130) and the normalized measurable nonlinear parameter(13m/130) In Fig. 3(a), solid line shows initial wave form and dotted line shows the waveform of ultrasonic wave propagated in partially degraded material and Fig. 3(b) shows the relationship between the degradation level(13dl30) and the normalized measurable nonlinear parameter(13m/130). In order to normalize the nonlinear parameter, all nonlinear parameters are divided by [30. Here, 13m/130 represents the calculated nonlinear parameter estimated from received signal which passed through the degraded part. In this case, the relationship between the degradation level(13d/[30) and the normalized measurable nonlinear parameter(13m/130) is not linear but somewhat nonlinear. Secondly, in order to know the relationship between the degradation range and the normalized measurable nonlinear parameter, computer simulation according to the increase of degraded node number shown in Fig. 2 (b) is performed. In this case, the nonlinear parameter 130 and 13dof in Eq (8) and (8)' was 3.47• 10-l~ and 6.94• 10-1~ respectively. That is, a in Eq. (1), in this case is 1. Fig. 4 shows simulation result of Fig. 2(b). Fig. 4 (b) shows relationship between the percentage of degradation node and the normalized measurable nonlinear parameter(13m/130). Here, the percentage of degradation node means the percentage of degradation node for whole node. Circles are values evaluated from the calculated datum. A solid line is curve fitting of these circles. The relationship between the normalized measurable nonlinear parameter(13m/[30) and the percentage of degradation node is linear. Here, if the percentage of degradation node is 100%, ~mand 13dare 2130 so that the value of 13m/130should become 2. Then, because the degradation level in this case is 60%, the value of 13m/130is near 1.6. From this result, we can confirm that this simulation result is right. If we compare Fig. 3 with Fig.4, it can be known that the magnitude of change in the normalized measurable nonlinear parameter(13m/[30) is different. That is, the range of degradation has bigger effect on the normalized measurable nonlinear parameter(13m/130) than the degradation level(13d/130). Thirdly, by assuming the distribution of 13 in degraded nodes as gauss distribution, we assumed the continuous distribution of degradation. The used gauss distribution was shown in Fig. 2 (c). tz in Eq. (1), in this case is 1. Here, the average value of magnitude of gauss distribution is agreed with the magnitude of degradation node of Fig. 4. Fig. 5 shows the simulation result of Fig. 2 (c). Fig. 5 (b) shows relationship between the percentage of continuous degradation node and normalized measurable nonlinear parameter(~m/130). Circles are values evaluated from the calculated datum. A solid line is curve fitting of these circles. From this relationship, the

244

K.C. Kim et al.

relationship between the normalized measurable nonlinear parameter(13m/130) and the continuous distribution of degradation can be explained. Then, if we compare this result with Fig. 4, it can be noticed that the normalized measurable nonlinear parameter(13m/130) of Fig. 5 (b) has higher value than that of Fig. 4 (b). This result shows that the nonlinearity in the continuous degradation is bigger than that of evenly distributed degradation, but both tendency is similar. So, the continuity or discontinuity is not an important point in our modeling 1.5

!

|

i

i

0.5

1

1.5

2

1 -~ 0.5 E < -0.5 -1 -1

"50

2.5

time(sec) x 10 l~ (a) the wave form used in simulation (solid line : initial wave form, dotted line : the waveform of ultrasonic wave propagated in partially degraded material) 2 1.8 1.6 ~

1.4 1.2

1 0.8

~ 10

0

~ 20

~ 30

J 40

t 50

60

Percentage of degradation node(%) (b) The relationship between the percentage of degraded node and the normalized measurable nonlinear parameter

Fig. 4. The result of computer simulation; (a) the waveform used in simulation (b) the relationship between the percentage of degraded node and the normalized measurable nonlinear parameter(13m/130)

1 -~ 0.5

~ o

~

< -0.5 -1 -1"50

I

I

i

I

0.5

1

1.5

2

2.5

time(sec) x 10 I~ (a) the wave form used in simulation (solid l i n e initial wave form, dotted line- the waveform of ultrasonic wave propagated in partially degraded material) 2

,

,

,

,

,

,

1.8

~6

~1.4

~_o~~~,~~..,~~-o

-~~

1.2

1 0.~v 0

~ 10

~ 20

~ 30

J 40

~ 50

60

Percentage of continuous degradation node(%) (b) The relationship between the percentage of continuously degraded node and the normalized measurable nonlinear parameter

Fig. 5. The result of computer simulation; (a) the waveform used in simulation (b) the relationship between the percentage of continuously degraded node and the normalized measurable nonlinear

parameter(13m/130) From these three kinds of results, the normalized measurable nonlinear parameter(13m/130)seems to be strong influenced by the range of degradation. Fig. 6 shows results of the normalized

Ultrasonic Nonlinearity in Degraded Material

245

measurable nonlinear parameter ( ~ m / ~ 0 ) calculated according to the increase of the percentage of degradation node in case of Fig.3 (b).

The percentage of degradation node = 60% o

o

&

1.4

o o

o

1.3 o

08 0.9

8~~

o o

~

~

o

o o

o o

o

~ o

o

o

~

o

o

o

o

o

o o

o o

o o

o

o o

o

o

o o o

o

o

o

o

o

o

o

o

o

o

~

o

o

~

o

o

~

~

~

o

o

o

~

o

~

o

o

o o

~ o

o o

~ o

i 1.5

o

o

o

~

o

o o50%o

~

~

o

o

o

o

o

o

40% o

o

~

~

~

~

o

o

~

~

~

~

o

o

~

o

o

o

30% 20%

~

~

~

o

10%

~

~176176176 ..................

000

o o

~

o

o

o o

o

~

o

o

o o

o

o

o

o

o

o

I

~130

2

2.5

Fig. 6. The tendency of normalized measurable nonlinear parameter(13m/130) according to the increase of the degradation level (13dl30)and the range of degradation From Fig.6, we can see the relationship becomes completely nonlinear as the increase of degradation range. From this result we can say that the degradation range has bigger effect on the normalized measurable nonlinear parameter(13m/130) than the degradation level(13d/fl0). As an example, 60% degradation With 1.5 degradation level results almost 30% change in normalized measurable parameter, which is much bigger than at 1% degradation with degradation level of 4.0. Unfortunately, however, we can measure only the normalized measurable nonlinear parameter(13m/130). We can not distinguish the degradation level(13d/130) and the degradation range from the normalized measured nonlinear parameter(13m/130). So, when a researcher get a small value of the normalized measurable nonlinear parameter(rim/130), he may consider the degradation is very weak and the target is still safe, in spite that there may exist a concentrated severe degradation in the target. So, in order to apply parameter [3 to actual NDE, this dangerous case should be considered. For example, the localization of 13 can be used as the method to know the distribution of [3.

COMPARISON WITH EXPERIMENT RESULT In order to verify the rightness of these simulations, the experiment result obtained from our previous research [2] was showed in Fig. 7 and compared with simulation result. In the previous experiment, the specimen was fatigued by cyclic loads so that the degradation level may be most in the center and decreases according to the distance from center. That is, this tendency of degradation is similar to the gauss distribution. So, third simulation result and previous experiment result were compared. In the previous research, the nonlinear parameter was estimated for the transmitted ultrasonic waveform obtained from 6 kinds of SS41 specimen fatigued by cyclic loads. In Fig. 7, horizontal axis means the repeated cycles of loads and vertical axis the normalized measurable nonlinear parameter(13m/130). Circles in figure represent the mean value of the normalized measurable nonlinear parameter(13m/~0) calculated from the results for three kinds of different exciting voltages and lower and upper bars show minimum and maximum value of them. This tendency between the repeated cycles of loads and the normalized measurable nonlinear parameter(13m/130) is similar with the simulation result. From this result, it can be known that the simulation method here mentioned is applicable to expect the ultrasonic wave behavior in partially degraded material and evaluate the degraded material.

K.C. Kim et al.

246 2.5

. . . . . . . .

i

. . . . . . .

I

. . . . . . .

&

-g 1.5

0.5

. . . . . . . .

10 0

I

101

. . . . . . . .

I

. . . . . . . .

I

10 2 10 3 Number of cycles

. . . . . . . .

10 4

10 5

Fig. 7. The normalized measurable nonlinear parameter(13m/130) estimated from SS41 specimens fatigued by several cycles

CONCLUSION 1) FDM(finite difference method) model for the nonlinear wave equation was developed with the restriction to the 1-D longitudinal wave motion 2) Using this simulation method, 1-D longitudinal wave propagation behavior in rod was showed. 3) The tendency between the normalized measurable nonlinear parameter and the continuous distribution of degradation obtained from simulation shows similar with the experiment result. From this result, it was confirmed that the proposed simulation method is right. 4) It was revealed that the degradation level, the range of degradation and the continuous distribution of degradation have strong correlation with the normalized measurable nonlinear parameter. Specially, the range of degradation has bigger effect on the normalized measurable nonlinear parameter than the degradation level and the continuously distributed degradation has a few bigger effect on the normalized measurable nonlinear parameter than the uniformly distributed degradation.

REFERENCES

.

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

Na, Jeong K., Cantrell, John H., and Yost, William T. (1996) Review of QNE 15, 1347. Jhang, K. Y. and Kim, K. C. (1999) Ultrasonics, 37(4), 39. Hikata, A., Chick, Bruce B. and Elbaum, Charles (1965) J. Appl. Phys. 36 (1), 229. Hikata, A., Chick, B. B. and Elbaum, C. (1963) Appl. Phys. Let. 3 (11), 195. Yost, W. T., Cantrell, John H. and Jr, Breazeale, M. A. (1981) J. Appl. Phys. 52 (1), 126. Shkolnik, Iosif E. and Cameron, Timothy M. (1996) The 14th ISNA, pp. 316-320. Sutin, A. (1996) The 14th 1SNA, pp. 328-333 Sato, T. W., Ma, H. Ninoyu, Jhang, K. Y. and Kousgi, Y. (1993) NDT & E lnternatinal, 26 (3), 119. Yamawaki, H. Saito, T., Masuda, C. and Fukuhara, H. (1994)Jpn. J. Appl. Phys. 33, 3126. Yamawaki, H. and Saito, T. (1996) Materials Science Forum 210-213, 589. Sharpe, R. S. (1982) In Research Techniques in Nondestructive Testing Vol. VI, pp. 107150 Academic Press Kim, K. C. and Jhang, K. Y. (1996) Proceedings of Asian Pacific Conference for Fracture and Strength '96, pp 923-928 AULD, B. A. (1973) InAcoustic Fields And Waves in Solids, pp. 68-84, John Wieley&Sons, Inc Truell, R., Elbaum, C. and Chick, B. B. (1969) In Ultrasonic Methods in Solid State Physics, pp. 38-52, New York Academic Press TenCate, J. A. and Abble, Koen E. A Van Den (1996),JASA 100, 1383. Hikata, A., Chick, B. B. and Elbaum, C. (1965) J. AppL Phys 36 (1), 229. Birks, A. S. (1991) In Nondestructive Testing Handbook 7 : Ultrasonic Testing, ASNT

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

247

QUANTITATIVE EVALUATION OF INTERLAMINAR-TOUGHENED CFRP COMPOSITES BY ULTRASONIC MICRO-SPECTROMETER

Y. OKABE and N. TAKEDA

Komaba Open Laboratory (KOL), The University of Tokyo 4-6-1 Komaba, Meguro-ku, Tokyo 153-8904, Japan

ABSTRACT The interlaminar-toughened CFRP composites T800H/3900-2 have interlaminar resin layers, whose thickness is about 30~tm, to suppress the growth of delamination. In this research, the elastic constants of the resin layer and the CFRP layer in T800H/3900-2 were evaluated quantitatively using an ultrasonic micro-spectrometer (UMSM). At first, the CFRP layer, whose thickness was about 140~tm, was made from only one prepreg sheet. Then, the acoustic reflection coefficients were measured at the surface of the specimen by the UMSM. Through the fitting of a theory to the reflection coefficients, all elastic constants of the transversely isotropic CFRP plate were obtained appropriately. Next, the interlaminar-toughened CFRP composite was ground to expose the interlaminar resin layer. Similarly, the elastic properties of the resin layer were evaluated quantitatively from the acoustic reflection coefficients measured at the surface of the exposed resin layer. As a result, elastic properties of the interlaminar resin layer could also be obtained in reasonable ranges.

KEYWORDS Interlaminar-toughened CFRP, acoustic reflection coefficient, elastic properties, quantitative, eval/~ation, ultrasonic micro-spectrometer

INTRODUCTION Interlaminar-toughened CFRP composites have interlaminar resin layers to suppress the growth

248

Y. Okabe and N. Takeda

of delaminations. One example of the composites is T800H/3900-2 [ 1]. This material system has interlaminar epoxy resin layers, in which tough and fine polyamide particles are dispersed, and the thickness of the layers is about 30gm as shown in Fig. 1. The elastic constants of interlaminar layers are needed for analysis of fracture process in the composites [2]. But the measurement of the properties by tensile tests is very difficult because the layers are very thin. Moreover, it is difficult to fabricate a bulk heterogeneous resin layer separately. In this research, the inversion of acoustic reflection coefficients is applied to determine the elastic constants of the interlaminar resin layer and the CFRP layer in T800H/3900-2 using an ultra-

Fig.1 Interlaminar-toughened CFRP composite T800H/3900-2 ([0/90]s).

sonic micro-spectrometer (UMSM) [3]. At first, the CFRP layer, whose thickness was about 140gm, was made from only one prepreg sheet. Then, the acoustic reflection coefficients were measured at the surface of the specimen by the UMSM. Through the fitting of a theory to the reflection coefficients, all elastic constants of the transversely isotropic CFRP plate were determined. Next, the interlaminar-toughened CFRP composite was ground to expose the interlaminar resin layer. Similarly, the elastic properties of the resin layer were evaluated quantitatively from the acoustic reflection coefficients measured at the surface of the exposed resin layer.

DETERMINATION OF ELASTIC CONSTANTS OF A CFRP LAYER

Materials A lamina made from only one prepreg sheet of T800H/3900-2 consists of a CFRP ply that has resin layers at both sides. This lamina was ground to remove the resin layers, and the CFRP plate without the resin layers was used as a specimen. Figure 2 shows the coordinate system of the specimen. Since this specimen is assumed to be transversely isotropic, there are five independent components of stiffness matrix C~1, C22/C44, Cs5 and C~2. The density of the specimen p is 1.64g/cm 3.

Experimental Apparatus

] x 3 Normal to Surface Incident P1ane 1 ~ /'fin a [ 0: Incident Angle ~ ~ ( ~ ~ _1 ~p:AzimuthalAngle

. , , - v~ / ~ X2 Specimen

k

Direction of Fibers

The reflection coefficients were measured by an UMSM. The UMSM uses two types of transducers:

Fig.2 Coordinate system of specimen.

Application of Ultrasonic Micro-Spectrometer to CFRP

249

a planar transducer used for vertical incident measurements [Fig. 3 (a)] and Spherical-Planar-Pair (SPP) lenses that can change the incident angle [Fig. 3 (b)] [3]. The SPP lenses consist of a spherical transducer and a planar transducer. A broadband ultrasonic pulse wave is transmitted from the planar transducer into water and reflected at the surface of the specimen. The reflected wave is received by the spherical lens focused on a small spot on the surface. Then the reflection coefficient is obtained by Fourier transformation for the reflected wave. This instrument can change the incident angle by tilting the SPP lenses. The specimen was immersed in water to face water at each side.

Fig.3 Schematic illustration of transducers used by UMSM: (a) Planar Transducer, (b) SPP Lenses.

Measurement of Thickness and C22 The thickness of the specimen should be known exactly for the quantitative evaluation of elastic constants by ultrasonic waves. Thus, the thickness was measured using the planar transducer. A reflector was set at the bottom of the water tank to return the longitudinal wave, which was transmitted through the specimen, to the planar rod. Then the time difference At was measured. The At is the difference in the propagation time between a wave propagated only in water and that transmitted through the specimen. Next, the longitudinal wave reflected at the surface of the specimen was measured as shown in Fig. 3 (a). The spectrum of the reflected wave is shown in Fig. 4. The spectrum has a period of minimum points Afclearly, which was caused by the interference of the wave reflected at the upper surface and that reflected at the lower surface. From the obtained At, Afand the wave velocity in

6 5

water vw, the thickness of the specimen d can be calculated using the following equations.

d = Vw(At + 1/Af)/2

(1)

Moreover, one component of stiffness matrix C 22 (2)

As a result, d and C22 were determined to be 142~tm and 16.2GPa, respectively.

1

0 0

(= C33) is given as C22 = p (2dAj02

,N2

50

100 150 Frequency (MHz)

200

Fig.4 Magnitude of the reflection coefficient measured at the incident angle of 0 ~

250

Y. Okabe and N. Takeda

Determination of C44 When the incident plane is set to be a 2-3 plane (~p = 900), the reflection coefficient depends on only two components of stiffness matrix C22 and C44, theoretically. Hence, the reflection coefficient was measured at the 2-3 plane using the SPP lenses as the incident angle was changed from 22 ~ to 38 ~ by increments of 0.2 ~ The frequency range was from 20 to 100MHz. The measured reflection coefficient was normalized by that measured for sapphire. The magnitude of the reflection coefficient is shown in Fig. 5. Many dip curves appear in the graph by the excitation of leaky Lamb waves. These dip curves reflect the elastic properties of the CFRP plate. Since C22 is already obtained, C44 is inverted through the fitting of a theory to the dip points. The reflection coefficient is calculated using A.H.Nayfeh's equation [4]. At first, the dip points are detected from the measured reflection coefficient. Next, the spectrum is calculated at each incident angle, and the dip points that correspond to measured ones are detected from the calculated spectrum. Then the objective function F is constructed as the sum of the squares of the differences between measured dip point frequenciesfM and calculated onesfc at all incident angles 0:

0

The objective function F is minimized by the optimum value of C44, and the downhill simplex method [5] is used for nonlinear optimization. In this optimization, only the dip points, whose incident angle 0 is from 22 ~to 26 ~ and frequency is from 20 to 60MHz, were used. Through the above procedure, C44 was determined to be 3.70GPa. Furthermore, Fig. 6 shows the reflection coefficient calculated using the obtained value of C44. Compared with Fig. 5, the calculated reflection coefficient in Fig. 6 reproduces not only dip points but also the change in magnitude depending on the frequency and the incident angle.

Fig.5 Magnitude of the reflection coefficient measured at the 2-3 plane (~p = 90~

Fig.6 Magnitude of the reflection coefficient calculated using the optimized value of C44 at the 2-3 plane (~p = 90~

Application of Ultrasonic Micro-Spectrometer to CFRP

251

Determination of the Other Three Components The reflection coefficient at a 1-3 plane (~p = 0 ~ depends on four components of stiffness matrix C~, C22, C55 and C~2. Hence, the reflection coefficient was measured at the 1-3 plane similarly. The magnitude of the measured reflection coefficient is shown in Fig. 7. This graph also has many dip curves. Thus, C~1, C55 and C~2, were optimized from the dip curves using eq. (3). In this optimization, the dip points .within 22 ~ to 38 ~ in the incident angle and 20 to 50MHz in the frequency were used. As a result, C55 converged on definite value. However, C1~ and C~2 could not converged, because the two parameters affect each other. Thus, the longitudinal Young's modulus E 1measured by tensile test was used as a restriction. The specimen for the tensile test was made through the same process. The obtained value of E 1by the tensile test was 146GPa. As for the longitudinal Young's modulus of unidirectional laminates, it has been shown that the value measured using ultrasonic waves is almost the same with that measured by tensile test [6]. From the optimization using the E~ as a restriction, all the three components converged on definite values: C~ = 158GPa, C55 = 7.88GPa and C12 = 12.4GPa. Furthermore, Fig. 8 shows the reflection coefficient calculated using the obtained stiffness matrix. This calculated reflection coefficient also reproduces the measured one very well. These agreements suggest that the stiffness matrix components were optimized accurately.

Fig.7 Magnitude of the reflection coefficient measured at the 1-3 plane (~p- 0~

Fig.8 Magnitude of the reflection coefficient calculated using the optimized stiffness matrix at the 1-3 plane (~p = 0~

Comparison with Tensile Test Results The elastic modulus and Poisson's ratios calculated from the obtained stiffness matrix are shown in Table 1. For comparison, those of T800H/3631 measured by tensile tests [7] are also shown. T800H/3631 has similar constituents of CFRP layers in T800H/3900-2. Except for E 1, all the values obtained by the UMSM are larger than those measured by tensile tests. This difference is

Y. Okabe and N. Takeda

252

a typical phenomenon for the elastic constant measurements of viscoelastic materials by ultrasonic waves owing to the viscoelasticity effect, which depends on the frequency [6], [8]. However, these values are used as the properties of the CFRP layer for the determination of the elastic properties of interlaminar resin layers, because the reflection coefficient at the surface of the resin layer was measured within the same frequency range. Table 1 The elastic constants calculated from the obtained stiffness matrix and those of T800H/ 3631 measured by tensile testsl Elastic moduli (GPa) E1

E2

G12

G23

Optimized from reflection coefficients

146

11.2 7.88

3.7

Tensile tests for T800H/3631

148

9.57 4.50

-

Poisson'sratios vl2

v23

0.416 0.514 0.356

-

DETERMINATION OF ELASTIC CONSTANTS OF A RESIN LAYER

Materials For the measurement of reflection coefficient on the surface of the interlaminar resin layer, the T800H/3 900-2 cross-ply laminate as shown in Fig. 1 was ground to remove the outermost 0 ~ ply, and the resin layer at the 00/90 ~ interface was exposed.

Measurement of reflection coefficients Using the SPP lenses, reflection coefficient was measured at the surface of the exposed resin layer, as the incident angle 0was changed from 12 ~ to 28 ~ by 0.2 ~ The incident plane was set to be parallel to the fiber direction of 90 ~ ply. The configuration of the measurements and the coordinate system are shown in Fig.9. Figure 10 shows the magnitude of reflection coefficient mea-

Fig.9 Configuration of the measurements and

Fig. 10 Magnitude of reflection coefficient

the coordinate system.

measured at the surface of the resin layer.

Application of Ultrasonic Micro-Spectrometer to CFRP

253

sured at the surface of the resin layer. In this graph, four dip curves appear by the excitation of leaky Lamb waves that propagate along the resin layer. Therefore, the elastic properties of the resin layer are determined through the fitness of a theory to the dip curves.

Analysis For the analysis, the specimen is assumed to consist of the isotropic resin layer and the 90 ~ ply that is the transversely isotropic semi-infinite body as shown in Fig.9. The elastic constants of the CFRP layer obtained from reflection coefficients are used as those of the 90 ~ ply. On the other hand, the resin layer has four unknown properties: Young's modulus E, Poisson's ratio v, the density p, and the thickness d. Thus, the four parameters are determined simultaneously from the dip curves in Fig. 10. The four parameters were optimized within constrained ranges. As a result, all the parameters converged on definite ranges. Table 2 shows the constrained ranges and the convergence ranges obtained from 20 results of the optimization. But the convergence ranges are wide. Particularly the range of the density is large. This is because four parameters were optimized at once. The other reason may be that the specimen was slightly different from the cal-

Table 2 Convergence ranges of the four

culation model owing to the waviness of the lower

parameters optimized within the con-

interface of the layer and the dispersion of fibers at

strained ranges.

the interface. However, all the convergence ranges

Constrained Convergence Ranges Ranges

are reasonable. Furthermore, Fig. 11 shows the calE (GPa)

0.5 - 10.0

v

0.05 - 0.50 0.203- 0.358

calculated reflection coefficients reproduce the

p (g/cm3)

1.00- 1.50

1.04- 1.46

change in magnitude depending on the frequency

d (~tm)

5.0 - 20.0

9.56 - 14.5

culated reflection coefficient using a combination of the converged values. Compared with Fig. 10, the

2.09 -4.70

and the incident angle. Moreover, in order to confirm the precision of the convergence, the optimization was carried out for only E and v by fixing p and d to be 1.26g/cm 3 and Table 3 Convergence ranges of only E and v. The other two parameters p and d were fixed to be 1.26g/cm 3 and 12.4~tm respectively. Constrained Convergence Ranges Ranges E (GPa)

0.5 - 10.0

3.65 - 3.82

v

0.05 - 0.50 0.233- 0.258

Fig. 11 Magnitude of reflection coefficient calculated using a combination of the converged values: E = 3.68GPa, v = 0.247, p = 1.23g/cm 3, d = 12.4gm.

254

Y. Okabe and N. Takeda

12.4~tm respectively, which are mean values of the 20 results. The results are shown in Table 3. These ranges are very narrow. Thus, if p and d can be determined exactly by other measurements, E and v will be determined more accurately.

CONCLUSIONS The elastic constants of the CFRP layer and the resin layer in interlaminar-toughened CFRP composite T800H/3900-2 were determined from the acoustic reflection coefficients measured by the UMSM. All elastic constants of the CFRP layer, which was assumed to be transversely isotropic, were obtained definitely. Although the obtained values were larger than those measured by tensile tests for T800H/3631, these values are usable as dynamic elastic constants. On the other hand, the convergence ranges of the resin layer properties were large, because the four parameters were optimized simultaneously. But, if the density and the thickness of the resin layer can be determined exactly by other measurements, Young's modulus and Poisson's ratio will be determined more accurately.

REFERENCES 1. Odagiri, N., Kishi, H. and Nakae, T. (1991). Proc. Amer. Soc. Composites 6th Technical Conf., pp.43-52. 2. Takeda, N. and Ogihara, S. (1998) Composites PartA, 29A, 1545. 3. Tsukahara, Y., Nakaso, N., Ohira, K. and Yanaka., M. (1996). In: Advances in Acoustic Micros-

copy, Vol.2, pp.103-165, Briggs, A. and Arnold, W. (Eds). Plenum Press, New York. 4. Nayfeh, A. H. (1995). Wave Propagation in Layered Anisotropic Media, Elsvier Science, Amsterdam. 5. Press, W. H., Teukolsky, S. A., Vetterling, W. T. and Flannery, B. P. (1993). Numerical Recipes

in C, Cambridge University Press, Cambridge. 6. Okabe, Y. and Takeda, N. (1997). Proc. 5th Japan International SAMPE Symposium, pp. 11191124. 7. Takeda, N., Kobayashi, S., Ogihara, S. and Kobayashi, A. (1998) Adv. Composite Mater., 7, 183. 8. Okabe, Y., Takeda, N., Yanaka, M. and Tsukahara, Y. (1999) IEEE Trans. Ultrason. Ferroelec.

Freq. Contr., 46, 1269.

CERAMICS AND CONCRETE

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

257

Q U A N T I T A T I V E D A M A G E E V A L U A T I O N OF C O N C R E T E C O R E S A M P L E S BY A C O U S T I C E M I S S I O N Masayasu OHTSU* and Tsuyoshi IIDA** * Professor, Graduate School of Science & Technology, Kumamoto Universitv2-39-1 Kurokami, Kumamoto 860-8555, Japan ** Nishimatu Corp., Tranomon 1-20-10, Minato-ku, Tokyo 105-8401, Japan ABSTRACT Damage estimation of structural concrete is studied, introducing acoustic emission (AE) measurement and damage mechanics. AE activity of a concrete-core sample during a uniaxial compression test is analyzed, based on the rate process theory. AE activity under compression is very high from low stress level, when concrete contains a number of microcracks. In contrast, AE activity of sound concrete keeps low until impending final failure. This discrepancy of the activity is quantitatively represented by the rate of AE activity, which is found to be substantially correlated with pore distribution of concrete. Then, the degree of damage is quantitatively defined on the basis of damage mechanics. Correlating the rate with the danmge parameter, quantitative estimation of the damage is proposed and its applicability is confirmed. KEYWORDS Acoustic emission, damage estimation, rate process, damage mechanics, concrete-core test. INTRODUCTION Nondestructive evaluation (NDE) techniques for concrete structures are recently in great demand for inspections responsible for maintenance. This is because it is seriously realized that the concrete structures are no longer maintenance-free. At particular, the increase of aged and damaged structures in Japan has emphasized on the need for repairing and retrofitting the concrete structures. In advance to repair and rehabilitation, inspections of the structure in service are normally necessary and knowledgeably performed to estimate the current state of the structure. In concrete, the damage of defects and cracks is primarily nucleated due to chemical reaction, mechanical stress, and fatigue. To estimate the damage of a concrete structure in service, a detailed inspection is usually conducted by taking core samples out of the structure. Both chemical and physical properties are then measured. As for mechanical damage, a compressive strength and Young' s modulus are determined and then are compared, if possible, with those of the specification. Besides, there exist few mature techniques to estimate the damage of concrete. To this end, measurement of AE (acoustic emission) activity in a uniaxial compression test of a core sample has been proposed [1]. It is found that AE generating behavior is closely associated with the presence of microcracks in concrete and can be analyzed quantitatively, introducing the rate process theory [2]. In the present paper, to estimate the damage quantitatively, damage mechanics is introduced. Then, it is attempted to estimate the damage parameter from AE rate process analysis. RATE PROCESS ANALYSIS When concrete contains a number of critical microcracks, active AE nucleation is expected under compression due to crack propagation from existing defects or microcracks. In contrast, AE

258

M. Ohtsu and T. Iida

activity in sound concrete is known to be stable and low prior to final failure. Thus, the presence of critical microcracks in concrete is closely associated with AE generating behavior under compression [3]. To formulate AE activity under compression, the rate process theory is introduced [1]. Probability function f(V) of AE occurrence from compressive stress level V(%) to V+dV (%) is formulated as a function of the incremental number of AE events dN as, f(V) dV = dN/N,

(1)

where N is the accumulated number of AE events up to stress level V(%), which is normalized by the compressive strength. When concrete contains a number of microvoids, AE activity under uniaxial compression is found to be high even at low stress level, while it is quite low for sound concrete. To discriminate these AE activities at low stress level, the function f(V) is assumed as the following hyperbolic function, f(V) = a/V +b,

(2)

where a and b are empirical constants. In Fig. 1, some results obtained in experiments are given. From uniaxial compression tests of core samples, histograms on the function f(V) are given, directly estimated as dN/(N'dV) where hyperbolic functions are determined by the leastsquare-error approximation. In Case A where AE activity is high at low stress level, the probability function is modeled as that the coefficient a is positive. Hereinafter, the coefficient is called the rate. In contrast, the function is modeled as that the rate a is negative in Case B, representing low AE activity at low stress level.

I

0.0 0.3

I,

,|,l,i

0.4

oxpo~imonL

|, _ 17

0.5 0.6 0.7 Stress level V Fig. 1

9

li ~

1

,

experiment

o

0.8 Stress level V

Rate process analysis of AE activity.

Substituting Eq. 2 into Eq. 1, a relationship between the number of total AE events N and stress level V(%) is obtained as, N = C V" exp (bV).

(3)

Here, C is the integration constant. DAMAGE MECHANICS From damage mechanics [4], the state of damage is represented by the scalar damage parameter f2. Based on the concept of continuum damage mechanics, Young's modulus E of a damaged material is expressed as E = E*(1- f~), where Young's modulus E* is that of an intact material.

(4) In the uniaxial compression test of a

Quantitative Damage Evaluation of Concrete by AE core sample, a relation between stress and strain is obtained as shown in Fig. 2.

259 Initial

Young's modulus Eo is associated with the current degree of damage f2o, as derived from Eo = E*(1 - f2o).

During the test, the damage increases from f2o to f2c, where f2c is the degree of

damage at failure and Young's modulus reaches Ec = E* (1 - f2c).

In this case, evolution of

damage f2c - f2o during the test must be associated with AE generating behavior. ! |

.",xi" ,',. .,-,.-" -!

-

Fig. 2

~

"-'.,,:'Ec ..,.

/ L-,

o["

,

I

Stress-strain relation and Young's modulus.

Concerning AE activity, an amplitude distribution is mathematically presented [5], n(a) = k a -m.

(5)

Here n(a) is the number of AE events of the amplitude from a to a+da. k and m are empirical constants. In particular, m is known as such a parameter for precursor in seismology that the value of m decreases prior to main shock. Assuming that the crack w)lume is proportional to the amplitude of AE event, the damage evolution is derived as, ff2c- ff2o = ~ n(a) a da = (l-m) AoNc / (2-m),

(6)

where Ao is the threshold amplitude of the measuring system. Nc is the total number of AE events. Substituting Eq. 3, eventually the evolution of damage is represented, f2c- f2o = (l-m) Ao C (100) a exp(100b) / (2-m).

(7)

RESULTS AND DISCUSSION

Rate a and pore distribution Based on Eq. 7, a correlation analysis was conducted in respect to the damage evolution (f2c f2o) and empirical parameters m, a, and b on the data of a variety of core samples. Then, it was found that the correlation coefficient between the rate a and the damage evolution was the highest of all other coefficients. Thus, it was decided to estimate the damage only from the rate a. Core samples were t ~ e n from an aqueduct of a nuclear power plant shown in Fig. 3. To drill core samples, three sites were selected : at the gate (Site A), 20 m inland (Site B), and 30 m inland from the gate (Site C). Core samples were mold-cut of 7.4 cm diameter and 14.5 cm height. A uniaxial compression test of the sample was conducted as shown in Fig. 4. AE events generated under compression was counted up to final failure by AE processor (LOCAN; PAC). AE sensor of I MHz resonance was attached at the middle of the sample, grinding the

260

M. Ohtsu and T. Iida

Site C SileB ~ ite A

|-

m

-

~

..:~..~_.--:

m

m

m

-

>

...-:=-._..::z:=

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

m

Core-drilled locations

1 Fig. 3

Locations of core samples in a nuclear power plant. Load |

,1

Loading Plate i AE Sensor

Pre-Amplifier

Concrete Sample

I 1 Fig. 4

] [

Locan AE [System

Experimental set-up for the uniaxial compression test.

attached surface as flat. For event-counting, the threshold level was set to 150 mV and the deadtime to 1 ms. The frequency range was from 10 kHz to 300 kHz, and total amplification was 60 dB. Three samples at each site were tested. Then, the uniaxial compressive strength and the rate a were determined as the averaged value of the three. In order to quantify microsca~pic damage, distribution of pore radii was also measured by the mercury intrusion method from concrete fragments at the three sites. After determining the pore distribution, the volume of pore radii over 0.5 btm was determined. It was reported that microvoids over 0.5 btm are predominantly responsible for deterioration of concrete [2]. Results of the pore volumes over 0. 5 btm radius, the rate a, and the compressive strength are summarized in Fig. 5.

From Site A to Site C, the pore volume over 0.5 btrn radius decreases

Quantitative Damage Evaluation of Concrete by AE

261

with the increase of the distance from the sea. This implies that the heaviest damage was introduced in concrete at the gate (Site A), where concrete was frequently deteriorated by seawater. Apart from the seaside, it is expected that the damage of concrete decreases. In accordance with the increase of the pore volume over 0. 5 ~tm radius, the rates a increase from Site C to Site A. Thus, the increase of the pore volume over 0.5 ~tm radius corresponds remarkably to the increase of the rate a. It suggests that with the increase in the rate a, the volume of the microvoids responsible for damage increases. Because the ratcs a are still negative, the degree of damage is supposed to be minor in these concrete samples. As a result, compressive strengths vary with little relation to the rate a. ,

0.03.

,"

.,

,

,

,

/ ---O- pore volurn~ over 0 5 lxm radius 9

/

- O - compressive strength

9

00+

j /

v

.#

J38

9 |

I

a ~"

9. . . . . . . . 35

Values of the rate a Fig. 5

Relation among the pore volume, the compressive strength, and the rate a in the core samples from a nuclear plant.

Damage due to freezing and thawing To estimate the damage quantitatively, concrete samples of 10 cm diameter and 20 cm height were made. The compressive strength at 28 days was 45.4 MPa. Controlled damage was driven into these samples, by the freezing and thawing test. After 28 day moisture-cure, a freezing and thawing cycle from -16 ~ Cels. to 3 ~ Cels. was repeated with three hours. Concrete samples were damaged by 100 cycles, 200 cycles and 300 cycles. Then, uniaxial compression tests of these damaged samples were carried out. Table 1 Mechanical properties of concrete samples

Freezing and

Compressive

Young's

thawing cycles strength (MPa) modulus (GPa)

Relative Young's modulus ,,,

o cycle

45.4

31.1

I00.0

100 cycles

42.4

31.1

141.5

200 cycles

45.7

29.3

96.6

300 cycles

43.7

29.6

57.7 ,

,

262

M. Ohtsu and T. Iida

Results of the compressive strength and Young's modulus are given in Table 1. Relative Young's modulus was determined from the resonance vibration of the sample during the freezing and thawing process. All these are averaged values of three samples. With the increase in freezing-thawing cycles, the compressive strength, Young's modulus, and relative Young's modulus decrease with some scatters. The damage level f2o at each cycle was determined from Young's modulus of each sample as Eo = E*(1 - f2o), setting E* to Young' s modulus of 0 cycle. According to Eq. 4, the damage evolution is determined from, f~c- f~o = (Eo- Ec)/E*.

(*,)

In an actual case of core samples drilled from a structure, Young's modulus E* of intact concrete can not be known. Therefore, a relationship between Eo- Ec and the rate a was investigated. Results of samples damaged by freezing and thawing are given in Fig. 6. A linear correlation is approximated as shown in the figure, where Eo - Ec is represented, (9)

Eo - Ec = E*(f2c- ff2o) = X a + Y.

Then, it is assumed that the initial Young's modulus Eo is identical to E* in the case that the rate a is equal to 0. As a result, E* is determined as, E* = Ec + Y.

(10)

From these equations, variation Eo/E* was obtained as shown in Fig. 7. It is clearly observed that the ratio of initial Young's modulus Eo to intact Young's modulus E* decreases monotonously with the increase in the freezing and thawing cycles. Thus, reasonable agreement between concrete samples of controlled damage and the damage estimated from AE rate process analysis. 20

I

'

I

'

I

.

I

Eo-Ec=Xa+Y X = -61.02 Y= 13.15

.

r,.9

'

15

0

0

-

10

m

I

-0.01

,

I

0

,

I

i

0.01

I

0.02

The rate a Fig. 6 Correlation between the rate a and Young's modulus in concrete samples.

0

100

200

300

The number of cycles Fig. 7 Ratios of Young's moduli Eo/E* in the freezing and thawing test.

From Eq. 4, the damage f2o is obtained as, f2o = 1 - Eo/E*.

(11)

Quantitative Damage Evaluation of Concrete by AE

263

Thus, the damages f2o estimated from AE rate process analysis are compared with the damages f2o determined from Table 1. Results are summarized in Fig. 8. All three samples of each cycle are plotted. As can be seen from Table 1, the damage does not monotonously increase with the increase in the freezing-thawing cycles. Still reasonable agreement are observed between the damages introduced in the freezing and thawing experiments and those estimated by the procedure based on the rate process analysis of AE. The initial damage f2o obtained in the experiment is always larger than that estimated from AE. The inital damage in the experiment was determined from the comparison with Eo at 0 cycle. Eo/E* at 0 cycle is, however, larger than 1.0 as seen in Fig. 7. Thus, the initial damage might be overestimated in the experiment. These results confirm an applicability of the procedure, where the damage of concrete at the current (initial) state can be estimated from AE rate process analysis without knowing the original state of construction.

,

,/1

0.1 ~ 11j t!j

.O

~ -0.1

~

-0.1

(~

0 cycle 100 cycles 200 cycles 300 cycles

'

01.1

g2o estimated from AE Fig. 8

Comparison of the damage obtained in the experiment and estimated by AE rate process analysis.

CONCLUSION Quantitative estimation of the damage in structural concrete is in great demand for inspection of concrete structures. To this end, AE measurement is applied to a uniaxial compression test of a concrete-core sample. Then, AE generating behavior is quantitatively estimated by the rate process analysis. Besides, the damage level is defined on the basis of the damage mechanics. Correlating the rate a of the rate process analysis and the damage parameter in damage mechanics, the procedure is developed and proposed to estimate quantitatively the damage of a concrete-core sample taken out of the structure. Results obtained are concluded, as fl)llows: 1) According to the results of the uniaxial compression tests of core samples taken from a aqueduct of a nuclear power plant, it is confirmed that with the increase in the rate a, the volume of the microvoids responsible for damage increases. This demonstrates that the rate a could give a quantitative reference on damage level of concrete. 2) A procedure to estimate the damage of the current state is studied, introducing damage mechanics and correlating the rate a with the damage parameter. From the correlation between

264

M. Ohtsu a n d T. Iida

the damage evolution and the rate a, Young's modulus of the intact state is estimated. By using concrete samples of controlled damage, which was introduced by the freezing-thawing process, the decrease of Young's moduli and the current damage are estimated by the procedure based on the rate process analysis. It is confirmed that these values estimated are in reasonable agreement with the actual values determined from the experiments.

REFERENCES 1. Ohtsu, M. (1992),"Rate Process Analysis of Acoustic Emission Activity in Core Test of Concrete," Concrete Library of JSCE, 20, 143-153. 2. Ishibashi, A, Hidaka, E., Farahat, A. M. and Ohtsu, M. (1995),"Deterioration Evaluation by AE in Concrete Samples," Proc. 6th Int. Conf. Structural Faults and Repair, 2, 69-74. 3. Matsuyama, K. and Ohtsu, M. (1992), "Rate Process Analysis of AE Activity to Evaluate the Deterioration of Concrete by Core Tests," Proc. AECM-4, ASNT, 132-138. 4. Kachanov, M. (1980), "A Continuum Model for Medium with Cracks," J. Engineering Mechanics, ASCE, 106(5), 1039-1051. 5. Mogi, K. (1962), "Magnitude Frequency Relation for Elastic Shocks accompanying Fracture of Various Materials and Some related Problems in Earthquakes, " Bull. Eartq. Res. Inst., 40,831853.

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

265

ESTIMATION OF DAMAGE PARAMETER AND CRACK VOLUME IN ROCK-LIKE MATERIALS BY SIGMA-AE PROCEDURE M. SHIGEISHI and M. OHTSU

Department of Civil and Environmental Engineering, Kumamoto University, 2-39-1 Kurokami, Kumamoto, Japan

ABSTRACT The moment tensor analysis of acoustic emission (AE) can provide quantitative information on kinematics of crack motions. Inplane unconfined compression tests of mortar and concrete plates having a slit are carried out. The slit, which directions are 0 ~ 30 ~ or 45 ~ to the loading axis, is prepared as an internal crack. From detected AE signals during the tests, the sources, that are, micro-cracks are located. In addition, the orientations of micro-cracks, types (tensile, shear or mixed), and volumes are also identified using the SiGMA-AE analysis. Furthermore, the damage parameters are estimated by the moment tensor components, determined in SIGMA procedure. The results of comparison between the calculated crack volumes and the estimated damage parameters show that there is a close correlation. KEYWORDS Acoustic emission, moment tensor, crack kinematics, damage mechanics

INTRODUCTION Acoustic emission (AE) phenomenon is an emission and propagation of elastic waves generated by releasing internal energy, as micro fracturing in an elastic material [ 1]. A procedure, which is named 'SIGMA', to determine AE source kinematics, such as AE source location, crack surface orientation, cracking motion direction and crack type, has been developed. Then this procedure has been applied to many cases of AE experiments in civil engineering [2]. Moreover, the SIGMA procedure could be applied to quantitative evaluation of crack volumes at AE sources. This paper shows the extended procedure for the quantitative evaluation of crack volumes and the results of application to the practical AE waveform data detected during the test of cement-mortar and concrete plates under inplane unconfined compression. Also, the results are investigated using the damage variables based on damage mechanics.

EVALUATION OF MICRO-CRACK VOLUME BY SIGMA-AE ANALYSIS

Acoustic Emission Moment Tensor Analysis Crack kinematics at the point y is modeled using the crack motion vector b(y, t) and the unit vector n normal to the crack surface F as shown in Fig. 1. The vector b is known as Burgers vector and represented as b(y)iS(t), where b(y) is the magnitude of crack displacement, I is the direction unit vector of crack motion, and S(t) is the source time function of cracking. The integration on the crack surface F is expressed using the second-tensor m on the crack kinematics, which is moment tensor:

M. Shigeishi and M. Ohtsu

266

fF C pqkl [b(y )l k S(t)]~tdS = [C pq~llkn I ]~ V

S(t)= m pq S(t ),

(1)

where Cpqklis the tensor of elastic constants and AV is the crack volume which is product of the magnitude b(y) and the crack area AF or AV = b(y) AF. In an isotropic material, the moment tensor is derived,

mpq = [Cpqtll,nt]~V = [~l,n,~pq +,lpnq +,lqnp]~V,

(2)

where ~, and ~t are Lame constants, and 8 is Kronecker's delta symbol. Referring to the generalized theory [3] of AE based on the elasto-dynamic field on the boundary element method (BEM), the elastic displacement u(x, t) due to the crack motion b(y, t) represents AE wave,

ui (x,t): Gip.q(X, y,t)m pq * S(t) ~

(3)

(amplitude:A)

bl~.-

j/

n~crack

surface;F

Fig. 1. Crack model and AE observation Moreover, the amplitude A (x) of the first P-wave portion at observation point x in the far field is simply represented,

A(x)= Cs Ref (s, r)ypm pqTq/R,

(4)

where Cs is coefficient containing the sensor sensitivity calibration. R is the distance between the source and the sensor. Ref(s, r) is the reflection coefficient associated with the sensor sensitivity directions s and direction of wave incidence r from the source. When the source location and the amplitudes of the first motion are known, the independent six components of the moment tensor can be determined by solving the simultaneous equations of Eq. (4) at each observation point. Applying eigenvalue analysis to the determined moment tensor, the crack kinematics can be obtained. A quantitative classification of the crack types into shear mode, tensile mode and mixed mode is developed. The eigenvalues of the moment tensor are decomposed into a shear component X, a deviatric component (CLVD) Y, and a hydrostatic component Z [4]. Three eigenvalues are normalized and uniquely decomposed into three ratios X, Y, and Z, (the maximumeigenvaluee2 / the maximum eI) 1.0 = X +

Y4-Z,

the intermediateeigenvaluee2 / the maximum eI = 0 - 0.5Y + Z, the minimumeigenvaluee 3 / the maximum eI = -X - 0.5Y + Z.

(5)

D a m a g e Parameter and Crack Volume

267

From the unit eigenvectors el = 1 + n, e2 = 1 • n and e3 = 1 - n corresponding to three eigenvectors, the vectors I and n can be recovered from the following relations, ! = ~[2 + 21knk .e, + ~/2- 21,n, .e 3, n=42+21kn , ' e , - 4 2

(6)

21,n,.e 3.

It should be noted that the moment tensor components are calculated as relative values and normalized. Hence, original magnitudes of the moment tensor must be reproduced for evaluation of the micro-crack volume. It is considered that the original moment tensor mpq is the product of the maximum component max(m) in the original moment tensor and the relative moment tensor m "pq, or mpq = max(m) m ~gq. In fact, Eq. (4) is represented as, A ( x ) R / C s Ref(s, r)7'p~'q = m pq - - m a x ( m ) . m pq.

(7)

Other, by relation k = 2~tv / (1-2v), Eq. (2) can be rewritten as,

mpq = AV~l(2v/1-2v)lkn,~pq +lpnq +lqnp]

(8)

Then, the maximum component max(m) in Eq. (7) can be derived by substituting the unit vectors ! and n in Eq. (6), into Eq. (8).

mpq :

max(m).m;q

=AV~tmax([(2v/1-2v)l, nk~)pq+lpnq +lqnpD. m~q.

(9)

Furthermore, substituting Eq. (8) into the Eq. (6), the equation for micro-crack volume b A V can be determined by, A(x)R/(Cs Ref(s,r)~tmax([(2v/1-2V)lknk~ m +Ipnq +lqnpD~[pm'pq~[q)=AV.

(10)

D a m a g e Variables on Acoustic Emission Source In the damage mechanics, the second-order damage tensor, ePkl is defined [5] as,

EPk,= 1/V* ~ AF(bkn, + b,n, )")/2,

(11)

where V* is the representative volume and AF is the area of one crack. Taking into account only one crack, the damage tensor e l kl for one crack is derived from Eq. (11), ~.lk, =l/V~e([b(y)lk]n ' +[b(y)lt]nk)dS/2:AV(lkn , +ltnk)/2V*.

(12)

The scalar damage variable for one crack D is readily obtained as, D = nk~.l~tn, = AV/V* lkn k .

(13)

From Eq. (2), the trace component is obtained, mu = (3~,+ 2~t~knkAV.

(14)

268

M. Shigeishi and M. Ohtsu

Consequently, the damage evolution can be represented as,

~O (i)= 1/V*~[AVIknk](i)= 1/IV*(3~,-1-2~[)~_~mkk (i)

(15)

.

i

This implies that the damage evolution process is estimated relatively from the accumulation of the trace components of the moment tensors. In addition, the relative crack volume can be obtained from Eq. (14) as,

(16)

(3k + 2~t)AV = mkk/[lknk ].

EXPERIMENTAL APPLICATION TO ROCK-LIKE MATERIALS AE Measurements in Inplane Unconfined Compression Test of Plates According to the SIGMA procedure and the damage variable analysis, evaluation of micro-crack volume was applied to the practical AE waveforms, which were recorded in the compression tests of plate specimens made by cement-mortar or concrete. Cement-mortar and Concrete plates were employed for inplane unconfined compressive loading tests. The sketch of the experiment is shown in Fig. 2(a).

Fig. 2. Inplane unconfined compression test overview The mixture proportion of cement-mortar was arranged so that the mass ratios of cement, sand and water were 1 : 2 : 0.6. Table 1 shows the mixture proportion of concrete. The mechanical properties of cement-mortar and concrete are shown in Table 2. Table 1.

Mixture proportion of cement-mortar and concrete Gmax W/C s/a Weight per unit mass (kg/m3) (mm) (%) (%) W C S

AE Slump Air nnl-mir (cc) (cm) (%)

G Concrete 20 50 48 172 346 834 1 021 104 8.0 Cement-mort 342 570 1 140 Gmax: maximum size of coarse aggregate, W/C: water-cement ration in weight, s/a: sand aggregate ratio in volume, W: water, C: cement, S: sand aggregate, G: coarse aggregate (crushed stone), AE: air entraining agent

5.0

The plate specimen was of dimension of 100 mm x 100 mm x 15 mm and contained a through-thickness slit of 1 mm wide and 20 mm long at the center of the specimen. The different inclination angles of the slits were prepared to be parallel, 30 degrees, or 45 degrees

Damage Parameter and Crack Volume

269

to loading axis. The configuration of the specimen and sensor locations is shown in Fig. 2(b). Four wide-band PZT piezoelectric AE sensors on both sides of the specimen detected AE waveforms. Total gain of the measurement is 40 dB. One AE waveform is digitized into 1024 words in length at 1 MHz sampling frequency. Tabel 2.

Mechanical properties Compression strength (MPa) Concrete 38.2 Cement-mortar 39.0

of cement-mortar and concrete Young Poisson's Velocity of elastic Density (g/cm3) modulus (GPa) ratio wave ~rn/sec~ 30.3 0.21 4 420 2.48 21.9 0.19 3 880 2.16

Characterization of A E Sensor using Davies-bar Technique To determine the micro-crack volume by the procedure explained above, it is indispensable that the AE sensor measures an absolute quantity for displacement. Hence, the dynamic characteristics of the AE sensor are examined. A simple method for characterizing AE sensors using a Davies-bar as a propagating media of elastic waves was presented [6]. Fig. 3 shows the principle of this method.

Fig. 3. Davies-bar technique AE Sensor calibration The sensor response corresponding to the input signal can be characterized in the frequency domain. When a projectile strikes the end of a bar, a compression pulse as elastic wave is generated. Moreover, this wave propagates along the bar as far as the bar deforms elastically. The equation of the one-dimensional elastic wave is represented as, 32u/3 t 2 = c 2 32u/3 z 2

(17)

C = #t-E/ p ,

(18)

where u is the displacement of the particle of the bar in the axial direction, t is time. z is spatial coordinate parallel to the axis of the bar. C is the velocity of the longitudinal elastic wave in the bar. E is Young' s modulus and 19 is the density of the bar. The wave reflects at the other end of the bar, where AE sensor has attached on, and propagates in the opposite direction as a tensile pulse. At the instance of reflection, the wave produces the displacement given as, d(t)= 2Cje(t)dt,

(19)

where d(t) is the displacement and e(t) is the strain of the incident pulse at the end of the bar. It is assumed that the cross-sectional area and the acoustic impedance of the bar are larger than those of the sensor, and that the influence of the presence of the sensor on the wave reflection is negligible. The waveforms of strain-gauge output, which are glued to the bar at the distance

270

M. Shigeishi a n d M. Ohtsu

L apart from the end, and the AE sensor signals are recorded by digitizing into N in length at A t sampling rate. The spectrums, F e and FA are given by Fast Fourier Transformation (FFT) of the strain-gauge output and the AE sensor signal, respectively. From the spectrum of the gauge outputs as the velocity, the displacement allowing the time function to the AE sensor is represented as, (20)

e. = l/2=f F e .

The sensor sensitivity S can be given as, (21)

s = FA/E= F A / ( F E / 2 ~ f ) = 2nf FA/F e .

In practice, the response of the whole system including Davies-bar with the sensor is calibrated. To detect the vibration of the Davies-bar, a laser interferometer could be useful for this purpose. The AE sensors, which were used in compression test of the specimens, were characterized using Davies-bar made of cement-mortar of the same mix proportion as the cement-mortar specimens. The bar was 1 350 mm in length and 20 mm diameter Two strain gauges were bonded 400 mm distant from the end. Elastic wave generated and propagated by striking of hammerhead, which was made of steel and the same diameter as the Davies-bar. The signal amplification of the strain gauges was defined so that the recorded 1 V was equivalent to 5000/~ strain. ,

~

0.5

~ 9

0

,

,

,~

t

g

o

9 -0.5

-0.5 -I

't

0.5

,

0

I

i

I

0.001 Elapsed Time (sec)

0.002

-I

i

0

I

(a) .

0

0.002

0.001 Elapsed Time (sec)

(b) .

100 200 Frequency (kHz)

.

.

300

0

I

.

,

.

I

100 200 Frequency (kHz)

.

I

300

(c) (d) Fig. 4. Davies-bar experiment results: (a) of output of strain gauge; (b) of output of AE sensor; (c) of spectrum of strain gauge output; (d) of spectrum of AE sensor output

Damage Parameter and Crack Volume

271

The typical signal of the gauge output and AE sensor output recorded in the test are shown in Figs. 4(a) and 4(b). The spectrums of those are also shown in Figs. 4(c) and 4(d). Figure 5 shows the characteristic curve of this AE sensor. The sensor sensitivities calculated at 60 kHz are summarized in Table 3. ~" 10

.

I

.

I

.

I

Table 3 Sensor Product name PAC IIT-1131313 PAC UT-1000 PAC UT-1000 PAC UT-1000

8

.,,~ r~

=

4

o

2

~

0 0

sensitivities Serial number lc)R (c.h 1~ 171 (ch.2) 126 (ch.3) 167 (ch.4)

Sensitivity iV/m) 2.1q• 1.94x103 1.93x103 1.78x103

!

100 200 Frequency (kHz) Fig. 5. Sensor characteristics curve

300

SiGMA-AE Application and Volume Estimation Applying the 'SiGMA-2D' analysis [7], AE sources were identified. The typical results are shown in Fig. 6. The arrow symbol (+--~) indicates the micro-crack classified into a tensile type and the arrow direction represents the crack opening direction. The cross symbol (• indicates shear type micro-crack and their two directions are of crack normal and crack motion. The observed crack opening of the specimen is shown by dotted lines in the results. The crack developments from the tip of slit in concrete ran zigzag through the coarse aggregate and the located AE sources are scattered around the wider area. In contrast, the crack developments in cement-mortar are very smooth and AE sources concentrate near the crack.

Fig. 6. SiGMA-2D results: (left) of a mortal plate; (right) of a concrete plate; with a 30 deg inclining internal slit Figures 7 shows the relations between the estimated crack volumes AV and the classified crack types in cases of specimen having a slit made inclination 30 degrees to the load axis. Comparing these two cases, the evaluated crack volumes of cement-mortar plate has a tendency to be larger than that of concrete one. This difference could be caused by presence of coarse aggregate, which hinders the crack growth, so that the crack volumes become low. From

M. Shigeishi and M. Ohtsu

272

also the SiGMA-2D results in Fig. 6, it is confirmed that the crack are prevented from developing by the coarse aggregate. Moreover, Figure 7 shows that the remarkable large cracks were classified into tensile-mode or mixed-mode. ,

0.004

0.002 ;~ 0.001

c,

o

,

~

,

0ooo

o

~ 0.0002

~

;~ 0.0001 I "

o

o0 Fig. 7.

,

~40

60

o,

o

'

o o

o0~ ~ * ~40' ~60]

100

'

1

o

"l

80

100

Shear ratio (%) Shear ratio (%) (a) (b) Crack volume estimation: (a) in a mortal plate (30 o); (b) in a concrete plate (30 o)

Damage Mechanics Parameters Figure 8 shows the distribution of the damage parameters mkk and mkfllknk in the classification of crack type. Here, all vertical values are only relative. However, the parameter values in the cement-mortar specimen are large, compared with those in case of the concrete specimen. Moreover, these distributions are very similar to the results of crack volume estimation in Fig. 8. These results confirm that the damage parameter derived from Eq. (16) can be regarded as the relative crack volume. [• -9] t• r ] , o , 2~ 800 ' 1 6f , o, , ' K_~

[•

0 .9]

o~

k

cP

oO

I

__~' . . . . ex'-,~,~"--.~-2,

[•

~1~

o o

'

1

I. . . . ~,..._ll,_r _o , 0 20 40 60 80 100 0 - 20- 40 60---- 8 0 100 Shear ratio (%) Shear ratio (%) (a) (b) Fig. 8 Damage parameters: (a) in a mortal plate (30 o); (b) in a concrete plate (30 o) 9 ~

i[ l

]l/[~t.i

i [ !]1]1 I l l

]1 ] L I I I [ ] i i , , i [ i I

IL]hrlKJ[

I

REFERENCES 1. 2. 3. 4. 5. 6. 7.

Ohtsu, M. (1998). Theory and Characteristics of Acoustic Emission. Morikita Publishing. Ohtsu, M. (1996). Magazine of Concrete Research. 48, pp. 321-330. Ohtsu, M. (1991). Journal of Geophysical Research. 96, pp. 6211-6221. Knopoff, L.; and Randall, M. J. (1970). Journal of Geophysical Research. 75, pp. 4975-4963. Kachanov, M. (1992). Applied Mechanics Reviews. 45, pp. 304-335. Ueda, K. and Umeda, A. (1991). Transactions of the Japan Society of Mechanical Engbieers. 57, pp. 143-147. Shigeishi, M.; Ohtsu, M. (1996). Journal of Material Science Japan. 45, pp. 1055-1060.

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

ELECTROMAGNETIC

273

AND ACOUSTIC

BY THE CRACKS

EMISSION

CREATION

GENERATED

IN SOLIDS

J. SIKULA, V. LOKAJICEK*, B. KOKTAVY, V. SEDLAKOVA, P. KOKTAVY AND J. PAVELKA Department of Physics, Brno University of Technology, Czech Republic Fax. +4205 4114 3398 E-mail: [email protected] 'Geophysical Institute, Czech Academy of Sciences, Prague Keywords: non-destructive testing, acoustic emission, electromagnetic emission, crack, granite ABSTRACT Electromagnetic (EME) and acoustic (AE) emission, which are generated in stressed solids, can be used to indicate cracks and defects position, dimension and their orientation. When a crack is created in a solid, the faces of the crack are charged due to loss of the chemical bounds. The faces of the crack make up plates of an elementary capacitor and the charges located on them constitute an electric dipole or quadrupole. The magnitude of the electric charge configuration and the dipole moment depend on the time. The EME signal is given by a redistribution of electric charges. As a EME sensor two conducting plates are fitted to the specimen making the parallel plate capacitor. The granite samples with dimensions 4x4x 1 and 8x4x l cm 3 were examined. They were loaded with a ramp increasing stress up to their failure. Electric voltage measured on parallel plate capacitor is time dependent due to vibration of crack walls. From the high frequency component, the length crack was estimated. We found crack length 0.3 - 1.5 mm. Crack position was determined in experiment with four capacitor and four acoustic emission sensors. The simultaneous investigation of EME and AE signals enable us to localize the crack position in solids and estimate crack dimension.

INTRODUCTION Cracks creation was studied by simultaneous measurement of electromagnetic (EME) and acoustic (AE) emission. This method is based on electromagnetic and ultrasonic effects related to generation of cracks in solids. On the faces of the new created crack electric charges constitute electric dipole system. The electric dipole system is a source of voltage induced on metal electrodes located on the sample faces. The induced voltage is directly proportional to the crack width and its active area. Occurrence of electric signal is correlated with acoustic emission events. The values of EME and AE signal amplitude are correlated, too. The frequency of EME signal is given by crack walls vibrations and we found frequency in the range from 1 to 6 MHz. The corresponding crack length is of the order 0.3 to 1.5 mm. The

274

J. Sikula et al.

second source of EME signal is due to sample mechanical vibrations with a frequency given by sample boundary conditions ( 5 0 - 500 kHz). The recorded electric signal is then modulated by an ultrasonic wave. The electrical conductivity of solids affects the electric charge relaxation and then EME signal disappear when crack walls are discharged. Simultaneous measurement of AE and EME emission can improve localisation of the crack position, cracks plane orientation and its dimension. EXPERIMENT The electromagnetic phenomena which are generated in stressed solids can be used to indicate cracks and defects in solids. Zamada [1], Lokaji6ek [2] and Mori [3] studied correlation between electromagnetic and acoustic signals. From experimental results it follows, that EME signals occur simultaneously with acoustic emission events. The EME signals with large amplitude occurs together with large AE events. Both signals increase just before the sample rupture (see Fig. 1.). Events frequency is very low (fE = 0.2 S-~) in the range of Hooke's law and reaches high value (fE = 4 S~) near sample breakdown. 160

0.20 .-.~ . . . . . . . . . . . . . . . . . . . . . . .

i

i

! ........

~--m=--8

0.15

120

..........

~..........

? ..........

0.10 80

iiiiiiii

40

iii

20

40

60

80

i U = Uo.exp(-th ) ! U o = 0.15 mV

i

......

0

-0.05

:~---

100

100

,

110

Dependence

of AE

.........

120

,

,

130

140

150

t/~s

a l MPa

Fig.].

.......... 'i

i..........

0.05

J

0

i,i ................................................

~ -

and EME

events

Fig.2. Electric voltage time dependence with exponentially time dependent average value

on mechanical strain for sample 4

t :: ~::~i:::.',:~ ii~'~

,~ .,:,"

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o

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'

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r :,,,,,=, h' r!::!:i:~ I,,

' :

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:

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60

80 t/~s

100

-0.2 [ l - - t ....... ~(',, f-tT--r-~[~ - -Z-F-:--ri~ ...... rt: I II ~ [ I / '. "I ~ t ' t,: ' t , : i AE f 6 t = 111~s ' i t ' -0. 3 1000 1050 1100 1150 1200 t/~s

IV'l,"'/': i'~;~rIiiL:fll',l',~;l~t:~!/!: !;ll!ll; 1111~1I li I:il~',~ iI~ -0.15 I-~iiill---J---i'~ ............. i ............ ~. . . . . . . . . . . . . .t l I , ~ i' t1'! "l ',1:~;/'l:'~ !; it'IIl!~;i i~ Illfil It i'(V',il II# ~, t~ I 1' i'Jl I~ '11'1 I"" '1'' I ir

:: i:~i

~0~3O 40

Fig.3. Electromagnetic EME signal time dependence

120

Fig.4. Electromagnetic EME and Acoustic AE signal time dependence

Electromagnetic and Acoustic Emission in Solids

275

A standard experiment set up was used which allowed to load the samples up to 100 kN. The measuring set up was described in [4]. In our experiments two conducting plates are fitted to the specimen and the electric voltage arising at this capacitor is amplified. The voltage across the plate capacitor depends on time (see Figs.2.,3.,4.). The EME signal appears instantaneously when the crack is created. Then the EME signal precedes AE signal due to different velocity of propagation. In Fig.4. this delay is 11 its and crack distance from AE sensor is 40 mm. Similar experiment made on sample dimension 8x4xl cm 3 provided with four capacitor sensors allow us to localise crack position more precisely. In Fig.5. the EME and AE signals vs. time are depicted. In this case EME signals appears at the same time and AE signals are delayed due to different distance between crack and AE sensors. Fig.6. shows position of AE sensors (No. 1-4), EME sensors (No. 5-8) and schematically the crack position (small circle denoted A). Capacitor sensors were connected to 4 low noise preamplifiers with frequency band from 10 to 700 kHz and then the high frequency component created by crack wall self-vibrations is lost and only low frequency component generated by whole sample mechanical vibrations is visible.

I

I

AE

i

3.

'

4. 5.

6.

EME 7.

0

i I

/-2_

I

!

~ I AAA^A

I

AI

~

/

i

3/vv AAP

l

1:o10 Xl

"c220"c:~4 t[ps]

30

40

Fig.5. Time dependence of EME and AE signals for event 7123

Fig.6. Crack position

DISCUSSION

We have two measurable quantities: a) mechanical displacement measured by AE sensors, and b) the electric voltage V(t) on load resistor R connected to parallel plate capacitor C (see [4]). We will suppose, that when the crack is produced, the electric charges +q and -q appear at the faces of the crack and their movement due to vibration induce time dependent voltage V on parallel plate capacitor. Our model is based on energy balance on source: qE.vdt - electrical energy created by charged crack walls motion, and load: R i 2 d t - electrical energy dissipated on amplifier input resistance and electrostatic energy CVdV of capacitor sensor.

276

J. Sikula et al.

Using this energy balance, it holds qE.vdt = Ri 2dt + C V d V dV

- -

dt

V

+ -

r

=

q

E.~.

(1)

(2)

CV

where r = R C is time relaxation constant given by sample resistance R and parallel plate capacitor capacitance C. When a crack is produced a certain portion of the mechanical energy is converted into vibrations. The time dependent displacement of the crack faces results in an AC electromagnetic signal component whose frequency corresponds to that of mechanical vibrations of the specimen. Due to this effect the crack walls are vibrating and crack width l is a function of time. The charges +q and -q are at the equilibrium distance l o and their displacement u(t) is given by u = uo e -St sin cot

(3)

where 6,, co are damped harmonic motion constants. The electrical conductivity is discharging the elementary capacitor representing crack walls. We will suppose the electric charge exponential time dependence in the form q =qo e -/~

(4)

Then the electric dipole moment p(t) is given by p = qo e-~ (l o + 2Uoe-8' sincot)p ~

where lo is the thickness of the elementary capacitor given by cracks walls distance.

Fig.7. Parallel plate capacitor and dipole moment orientation for a crack

(5)

Electromagnetic and Acoustic Emission in Solids

277

From (2) and (5) we have differential equation for voltage V(t) dV

V

dt

r

--+--

= ge -~ sin(cot + ~o)

(6)

where g

.

_

2q0v 0 cos a

.

(7)

Cd

7 = f i + f l describes EME signal damping due to sample electrical conductivity ([3) and mechanical damping (8). The constant g is given by the value of electric charge qo, velocity amplitude Vo, parallel plate capacitor thickness d and the angle o~ between electric field intensity E and electric charge velocity ~', describing orientation of crack walls with respect to parallel plate capacitor (see Fig.7.).

The electric voltage V(t) on capacitor C is then solving (6) given as a sum of two components: 1) homogeneous equation dV

V

+

....

o

(8)

V(t) = Voe ~

(9)

dt

r

which has solution in the form

where V0 is the integration constant. This DC component of electromagnetic signal will be dominant if crack walls are perpendicular to the plates of measuring capacitor C. In this case vectors of electric field E and charge velocity ~ are perpendicular and c o s a = O. ii) total differential equation which has solution ge-rt sin(co.t+(,o) V(t) = Voe-'/~ - ~coo2 + co2

(10)

where Vo is integration constant and coo = 7"- 1/ r is cut off angular frequency.

Crack length We found, that the frequencies of the EME signals for granite samples were f c = 1 to 6 MHz. The crack length l c can be estimated from this frequency and for granite sample with velocity of mechanical wave propagation 3 km/s we find crack length l C = 1.5 - 0.25 mm. Crack's eigen vibrations are attenuated due to mechanical damping and electrical sample conductivity. In our experiment this high frequency crack vibration was observed for a time interval of the order 10 9s. After this initial time interval, the main source of crack wall

278

J. Sikula et al.

vibration is given by the all sample vibration due to acoustic emission energy. This frequencies are given by sample geometry and frequency of EME signal was about one order lower than for initial crack walls vibration.

CONCLUSION Mechanical stress creates electrical potentials on the walls of the sample. Electrical voltage is a measurable quantity for such effects and generally this is created by redistribution of electric charges, piezoelectric phenomena and by friction during crack walls sliding. There are two sources of acoustic emission events: tensile and shearing processes. Acoustic emission events appear either when new cracks are generated or by shearing process in existing cracks. We suppose, that electric charge redistribution is a main effect, which is manifested in our experiments. Electric charge redistribution is related to time dependent electric dipole moment and then electromagnetic field must be generated. Measured voltage has two components, DC and AC given by crack walls vibration and mechanical parameters of measured sample. High frequency component is of the order form 1 to 6 MHz and give information on crack length. We estimate the length of the crack of the order of 1 mm. Sensitivity of the value of electrical potential on crack orientation with respect to parallel plate capacitor can be used to study orientation of sample's layers. Low frequency component below 500 kHz give information on mechanical vibration spectrum and this component is correlated to acoustic emission signals. For our samples, this frequency was of the order of 100 kHz. The second time constant, which is given by parameters of mechanical vibrations and electrical conductivity is dependent on humidity. ACKNOWLEDGEMENTS This work has been supported as a part of a research project CEZ: J22/98:261100007 and by grant MSMT KONTAKT ME 285. REFERENCES

1. I. Zamada, K. Masuda, H. Mizutani, Electromagnetic and acoustic emission associated with rock fracture, Phys. Earth. Planet Int., 57, pp. 157-168, 1989. 2. T.Lokaji6ek and J. Sikula, Acoustic emission and electromagnetic effects in rocks, Progress in Acoustic Emission VIII, pp. 311-314, 1996. 3. Y. Mori, K. Sato, Y. Obata and K. Mogi, Acoustic emission and electric potential changes of rock samples under cyclic loading, Progress in Acoustic Emission/X, 1998. 4. J. Sikula, B. Koktav~, P. Vagina and T. Lokaji6ek, Cracks and quality testing by acoustic and electromagnetic emission, Proc. of the 23rd European Conference on Acoustic Emission Testing, TUV Austria, Wien 6-8 May, pp. 130-133, 1998.

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

279

N O N D E S T R U C T I V E C H A R A C T E R I Z A T I O N OF THE H Y D R A T I O N P R O C E S S IN TRICALCIUM SILICATE

B. B. Djordjevic*, L. L. Rouch* E. Shchukm , E. Amehna, I. Vidensky* Center for Nondestructive Evaluation The Johns Hopkins University Baltimore, MD 21218, USA 9

,+

9

+

ABSTRACT Early stages of hydration mechanisms and mechanical properties of the developing cement microstructure were studied using monoclinic and triclinic Tricalcium Silicate (C3S). The characterization studies utilize ultrasonic waves and X-ray diffraction methods to observe microstructure evolution and phase transitions during hydration of the C3S. Ultrasonic attenuation has shown number of microstructure transitions and significant changes in hydration rate affected by water ratio and temperature. X-ray diffraction studies indicate very large residual stress in the microstructure lattice that is self-induced during hydration process. C3S and water are principal factors in portland cement. The ultrasonic sensors and x-ray diffraction studies permit in-situ monitoring of the hydration process kinetics. These methods can be used to monitor process and effects on the final material physical properties due to early hydration dynamics. KEYWORDS Hydration of cement, ultrasonic monitoring, ultrasonic attenuation, X-ray diffraction, residual stress INTRODUCTION Cement invented by Romans is a dominant building material and major component of our civil infrastructure [ 1]. Formation of concrete by hydration process of cement mixtures is believed to be an understood process. In fact, the chemistry of formation, hydration, mechanical stability and durability of concrete materials is scientifically incomplete. To better understand and analytically define the mechanical and chemical integrity of the cement, one must understand the details of the complex interactive process of the chemical, physical and mechanical activities during the hardening stages and development of strength in concrete. The reaction between tricalcium silicate(C3S) and water is the principal factor in setting and hardening of portland cement [2]. Many investigators studied this reaction. However, the true mechanisms and the kinetics of the process are still not known [3]. The process of cement cure impacts durability and stability of the concrete material that is often compromised during modem construction methods using process accelerator additives [4]. Presence of water and activation energy are critical elements of the cement hydration process [5] and rate of the concrete formation. * - The Johns Hopkins University, Baltimore, USA, + - Moscow State University, Moscow, Russia.

280

B.B. Djordjevic et aL

Chemistry of phase transitions during hydration process influences quality of mineral binders and bridging process in cements. Strength, durability, and stability of the concrete depend upon microstructure formation mechanism. Recent work [6] showed that bridging of particles forming new phase crystals requires particle compression to achieve bridging crystallization. To better understand the hydration hardening, we must understand individual particles bridging, what drives the bridging and how to control the process? It is reasonable to assume that bridging of the individual particles during hydration and sol-gel transitions is at cystalographicaly favorable nucleation sites. Similar to processing of other materials, in concrete we can distinguish between microstructures based on coagulation or phase contacts. In a coagulation structure, the bridging is limited to simple "touching" (weak contacts; 10-9 to 10-8 N) and is reversible. In a phase structure, the bridging of particles is by valence bonding. Such bonding is relatively strong (above 10-6 N) and is typically irreversible. The transformation from coagulation to phase contacts is usually the result of processing. In cementitious materials, the very fine phases are formed by nucleation of new, highly dispersed hydration products forming a condensational structure. A transition from coagulation to phase contact in such dispersions, highly divided structures reflects the very essence of strength development. During formation of the bridges between particles, the cement crystalline structure develops internal mechanical stresses. Expansion or contraction during crystallization phase transformation, and interference between adjacent bridging crystals causes stresses. These stresses can promote bridging of particles and thus increase structural strength. Concurrently, these stress conditions can act as residual stresses that cause drastic decreases of both strength and durability of the final product. Local compressions in contact zones between particles are balanced with tensions in the adjacent areas; thus, micro-stresses (mode 11 stresses) arise in the structure. The stress state can be monitored by broadening of the X-ray diffraction lines. The measurements with hydrated gypsum and other mono-mineral binders showed that these micro-stresses grow to the level of the material strength. During hydrated material drying, the macro-stresses (mode I stresses) are also created by capillary forces [7]. Creation of internal stresses is part of the structure formation process. However, if internal stresses are not allowed to relax, the hardened concrete matrix develops large internal residual stress with commensurate decrease in the resistance to fracture and cracking. Such residual stress state should impact concrete stability in aggressive environment and aging performance. ULTRASONIC MEASUREMENTS Ultrasonic attenuation measurements are extensively utilized for monitoring the material microstructure changes. Ultrasonic C-scan is commonly used to map ultrasonic attenuation signals. Figure 1 is a time series of such ultrasonic C-scan maps for portland cement sample at different stages of hydration. The 1.0 MHz ultrasonic signals measure overall and local sound attenuation changes due to cement microstructure formation. The ultrasonic C-scan images show non-uniform and rapid microstructure change in the

Hydration Process in Tricalcium Silicate

281

the formation of the concrete. To enhance attenuation measurements, we have developed small 0.5 MHz probes with elastomer coupling to improve sound attenuation measurements during early hydration of C3S. Figure 2 shows a schematic set-up for measurement of ultrasonic attenuation. Attenuation measurements during cure using 0.5 MHz signals demonstrated strong sensitivity to microstructure changes and development of interconnections between the particles. Examples of observed attenuation changes in the C3S samples are shown in Fig. 3 and 4. Measured signal changes often exceeded 3 orders of magnitude.

Figure 1. Ultrasonic C-scan map of portland cement sample at different hydration stages. This ultrasonic attenuation map of the sample indicates rapid microstructure change and inhomogeneous hardening of the material. Light areas indicate low attenuation, dark areas are representative of high attenuation. Time in minutes after addition of water are shown on the pictures.

Figure 2 Ultrasonic test configuration for attenuation measurements for hydration of the Tricalcium Silicate

282

B.B. Djordjevic et al.

The ultrasonic attenuation has shown a strong response associated with the hardening of the C3S. A typical ultrasonic attenuation signal as a function of the time shows several transitions. These transitions can be associated with grain growth kinetics and microstructure material changes during hydration process. It isnot yet clear what specific microstructure changes are responsible for the ultrasonic interactions. However, the fact that these changes occur indicates the hydration reaction is complex and dynamic process. We have also observed overall ultrasonic signal response that is greatly dependent on the sample preparation, water ratio and temperature. Moreover, identically processed samples were remarkably similar in their ultrasonic response. The through transmission set up in these experiments was across about 5-mm thick disk. Signal transmission was monitored over days with fixed test geometry. The most significant changes are observed in the first 10 hours of the hydration [8].

2.5

2

i 1.5

1

0.5

TIME 10 hours/division

Figure 3. Ultrasonic signal intensity change for a through transmission transducer configuration across 5 mm C3S sample.

.~

0.8

m 0.6

T I M E 10 hours/division

Figure 4. Ultrasonic signal intensity change for a through transmission transducer configuration across 5 mm C3S sample.

Hydration Process in Tricalcium Silicate

283

X-RAY STUDIES X-ray diffraction lines are poorly defined in many cements because of the highly dispersed, multi-component, and partially amorphous structure. The C3S samples and the hydration products do not show any useful X-ray diffraction lines (6). However, moderate addition of fine disperse crystalline powder of barium sulfate (BaSO4) can be used as indicator (sensor) for sensing strains and stresses change during hydration process. BaSO4 crystals belong to rhombohedric (trigonal) symmetry group and the stronger diffraction lines are better suitable for measurements of the line broadening. The X-ray line has dual peak at 25.21 ~. Typical micro-photometric curve of the line can be established directly from X-ray diffraction pattern of the fine BaSO4 powder. For purpose of this study, many repeated Xray patterns of the fine disperse BaSO4 powder were used to establish the baseline data. The study of intemal residual stresses during hydration hardening of C3S, (monoclinic) was performed using 15% of BaSO4. The mixture thus incorporated the inert sensor with measurable diffraction lines. The angular shift and line broadening were recorded on hiresolution film during hydration hardening from one to seven days. The experimentally measured line widths were analyzed for geometrical and physical broadening to separate strain induced broadening due to mode I and 11 stresses. Micro-stresses (mode II) arise in the structure due to local compression in contact zone between new phase particles balanced by tension in adjacent area. Macro-stresses (mode I) arise in the course of drying of the hydrated material, due to capillary forces. Micro-stresses cause the X-ray diffraction line broadening B and macro-stresses are observed as the shift of lines A| Fine disperse BaSO4 powder (0.1 lam) added to the mix as a sensor for evaluation of internal stresses in hardened C3S samples. Experimentaly measured broadening of x-ray B can be defined by: B2= b z + 132 B -experimentally measured width of the line, b - instrumentally induced part of the broadening, 13- physical broadening of the line. Where 13= (138 + 213a)2 / (13~+ 413a) and 13~- line broadening due to dispersion of the powder, 13a- line broadening directly caused by deformation of the crystal lattice. For these experiment we found that: b = 4.6xl03rad for BaSO4 line (O = 25.21 ~ was constant in all experiments; 13~= 16. lxl 0 3rad was found from fully hydrated (large excess of wate0 samples of C3S + 15% BaSO4 and is denoted as BaSO4 standard. Using microphotometric curves derived from diffraction data as shown in Figure 5 we were able to estimate the internal residual stresses in sensor crystals to be as high as 0.50 GPa after 7 days of hydration. Shown in Figure 6 is microphotometric data used to determine the deformation broadening 13a and shift A| of BaSO4 line for hydrated "dry" C3S samples.

284

B.B. Djordjevic et al.

Figure 5. Typical diffraction image from the BaSO4 used to develop microphotometric curves. The diffraction lines on original film are more distinct and have more contrast

!

J

iii

C

i I

&

ii

|'L t t

i i

C ID

1 Reference J

9

l l Day J

,Iv '1 ,1 "1

.! /

[7 Day

,i 7

I

-

k

]

I

J - V

,,

10 .2 tad

ip~

O

Figure 6. Microphotometric data for C3S (monoclinic)+15% BaSO4, water/solid ratio 0.4 sample "Dry" samples - hydration hardening at 100% humidity during 1 and 7 days, and 30 min drying in open air Internal residual stress can be estimated from lattice deformations related to stresses mode II and mode I as follows: Ad/dll ~ ~d, and Ad/di ~ AO With internal residual stresses mode I I an ~ E 13d and internal residual stresses mode I: ol ~ E A| Using this analysis, table in Figure 7 estimates the level of internal residual stresses for sensor crystals.

Hydration Process in Tricalcium Silicate

[GPa] ax [GPa]

ffIl

1 day "wet" 0.14 0.47

1 day "dry" 0.21 0.63

7 days "wet" 0.25 0.45

285

7 days "dry" 0.14 0.50

(The ot values may include also some part of the oriented OII component) Figure 7. Table of estimated internal residual stress levels in the C3S samples. The measurements show internal stresses can reach 50 MPa and more, i.e. they are as high as the material strength. Residual stresses has been observed to depend on the hydration conditions including dispersion of the initial material mix, water/solid ratio, time (various stages of hydration), fillers, and on surface-active additives. Thus, measurement of the internal residual stresses appears to be sensitive method for the detection and control of the concrete material latent instability. CONCLUSIONS Ultrasonic attenuation and X-ray diffraction studies were performed to monitor hydration process in C3S. Ultrasonic attenuation has shown a strong response associated with the hardening of the C3S. A typical ultrasonic attenuation signal as a function of the time shows several transitions. These transitions can be associated with grain growth kinetics and microstructure material changes during the hydration process. The X-ray data indicates that internal hydration stresses can reach 50 MPa and more, i.e. they are as high as the material strength. Ultrasonic attenuation and X-ray diffraction measurements have been observed to depend on the hydration conditions: dispersion of the initial material mix, water/solid ratio, time (various stages of hydration), fillers, and, especially, surface-active additives. Thus, these tools appear to be the unique method for the characterization and further study of the concrete material. REFERENCES Gani, M. S. J. (1997) Cement and Concrete, Chapman & Hall, London, Bogue, R. H. (1955) The Chemistry of Portland Cement, Reinhold Pub. Corp., New York. Glasser, F.P., Lachowski, E.E., Macphee, D.E.,(1987) J. Am Ceram. Soc. 70, 481. Peterman, M. B., Carrasquillo, R. L., (1986) Production of High Strength Concrete, Noyes Publications, New Jersey. Fitzgerald, S. A., Neumann, D. A., Rush, J. J., Bentz, D. P., Livingston, R. A., (August 1997) An in-situ qusi-elastic neutron scattering study of the hydration of tricalcium silicate, Chemistry of Mat. Shchukin, E.D., Rybakova, L.M., Kuksenova, L.I., Amelina, E.A. (1997) Colloid J. 59, 90 Amelina, E.A., Kuksenova, L.I., Parfenova, A.M., Rybakova, L.M., Bessonov, A.I., Shchukin, E.D.,(1997): On the mechanism of the effect of organic additives on the properties of cement hardening structures: l. The effect of organic additives on the mechanical properties and internal microstresses (Mode II). Colloid J. 59 96-101. Skalny, J., Odler, I. J. (1972) Colloid lnterface Sci. 40, 199 Malhotra, V. M., Carino, N. J., Ed (1991) CRC Handbook on Nondestructive Testing of Concrete, CRC Press, Boston.

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LASERTECHNIQUES

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

289

Visualization of Elastic Waves by Laser Ultrasonics Junji TAKATSUBO, Masaaki IMADE and Shigeyuki YAMAMOTO Chugoku National Industrial Research Institute 2-2-2 Hiro-suehiro, Kure, Hiroshima, 737-0197, Japan

ABSTRACT We have developed a laser ultrasonic system for visualization of elastic waves propagating on a solid surface. Scanning an optical heterodyne probe to measure surface transient displacements, ultrasonic waves in the frequency range up to 50MHz are detected and passed to a personal computer through a digital oscilloscope. The recorded signals are reconstructed to make 3-D displacement images at any propagation time.

This system can

produce a series of successive images as an animation of wave propagation. Using this visualization technique, we observed the scattering and diffraction of ultrasonic waves around various shapes of artificial defects, and examined its application to nondestructive inspection

KEYWORDS Ultrasonic wave, nondestructive inspection, laser, visualization, crack, defect

1. I N T R O D U C T I O N Ultrasonic waves in solid media have become an important subject in nondestructive evaluation. Techniques for visualizing wave propagation can be very useful with regard to both the educational and research aspect of studying various wave propagation phenomena C1). The photoelastic technique and the Schlieren technique have been used for the visualization of ultrasonic waves (2). However, these techniques are applicable only to transparent media. In order to visualize ultrasonic waves propagating on opaque media, we have developed a digital ultrasonic imaging system using an optical heterodyne interferometer. This visualization system can produce a series of successive images as an animation of wave propagation. Using this method, we observed the propagation of ultrasonic waves around various shapes of artificial defects such as cracks, notches and penetration holes. We have been confident that this technique is available for nondestructive inspection and materials characterization.

290

J. Takatsubo, M. Imade and S. Yamamoto

2. V I S U A L I Z A T I O N S Y S T E M O F U L T R A S O N I C W A V E P R O P A G A T I O N A new developed ultrasonic imaging system is shown in Fig. 1. In order to generate ultrasonic waves, we can use either pulse laser or ultrasonic transducer as a oscillator. As shown in Fig.2, the ultrasonic waves

generated

by

laser

stimulation

propagate as spherical waves, while the waves emitted from ultrasonic transducers propagate as plane waves. In this experiment, we used ultrasonic transducers, because

Fig.l Measuring system for visualization

plane waves are fundamental waves to

of ultrasonicwaves

understand wave scatterings. An ultrasonic probe

is

mounted

on

a

specimen

for

generating repeated ultrasonic pulses. The ultrasonic displacements on the surface of the specimen are detected by scanning a laser

optical

probe

(optical

heterodyne

interferometer) with a minimum pitch of 1 # m. The detected displacements are stored in

a

computer

measurement,

hard the

disk. stored

After

the

data

are

reconstructed as animation images of ultrasonic wave propagations. By use of this

Fig.2 Method for generation of ultrasonic

imaging system, we can measure nano-meter

waves.

order displacements in the frequency range up to 50MHz. 3. B A S I C T Y P E S O F VISUALIZATION IMAGES In our visualization system,

ultrasonic

waves at all scanning points

are

stored

as

digital data, so that we can offer a variety of visualization

images O)

Fig.3 Dimensions of the penetration holes on aluminum plates.

Visualization of Elastic Waves by Laser Ultrasonics

291

Fig.4 Specimen with cylindrical hole for the visualization of ultrasonic waves. as given below. The measurements for these

images

were

conducted

using

aluminum specimens having different shapes of artificial defects (penetration holes), as shown in Fig.3. Figure 4 shows the dimension of the specimens.

Fig.5 Propagation images of the ultrasonic waves propagating around various shapes of artificial defects.

PROPAGATION IMAGE This image shows the displacement map of the ultrasonic waves at any propagation times. Figure 5 shows the propagation images of the ultrasounds propagating around various shapes of artificial defects, from which we can easily understand the scattering of the ultrasonic waves around the defects. We can offer the animation images on a computer display as well. AMPLITUDE IMAGE Amplitude image shows the maximum amplitude map of the gated wave, which corresponds to C scan image in ultrasonic inspection. Figure 6 shows the amplitude images around different shapes of defects. From these images, we can observe the intensity

distributions

of

the

scattering

waves around the defects. ARRIVAL TIME IMAGE Arrival time image shows the arrival time map of the wave front. Figure 7 shows

Fig.6 Amplitude images of the ultrasonic waves propagating around various shapes of artificial defects.

J. Takatsubo, M. Imade and S. Yamamoto

292

Fig.7 Arrival time Images of the ultrasonic waves propagating around various shapes of artificial defects.

Fig.8 Velocity Image of the ultrasonic wave propagating around a cylindrical penetration.

the arrival time images around artificial defects. VELOCITY IMAGE Velocity image corresponds to B scan image in ultrasonic inspection. For example, a series of waveforms detected on the line A in the right top figure in Fig.8 is shown in the left figure. When these waveforms are drawn as a B scan image, we can obtain the velocity image around a cylindrical hole as shown in the right bottom figure. The gradients of the inclined lines in Fig.8 represent the sound velocities. The sharp slope means the low velocity. Thus, we can easily estimate the sound velocities. 4. V I S U A L I Z A T I O N O F U L T R A S O N I C WAVES A R O U N D V A R I O U S K I N D S OF ARTIFICIAL DEFECTS In this section, we introduce some examples of the visualization of ultrasonic waves propagating around various kinds of artificial defects. Figure 9 shows the propagation images of the ultrasonic waves propagating around both slit and fatigue crack. The depths of these defects are about 10mm, and the width of the slit is 0.3mm. An angle beam transducer was used for emitting ultrasonic waves as shown in the left figure in Fig.9. The nominal frequency of the transducer was 5MHz. The wave propagations around these defects are different with each other. In the case of slit, the ultrasonic waves cannot transmit the slit but creep around the slit. On the other hand, in the case of fatigue crack, the ultrasonic waves almost entirely transmit the crack.

Visualization of Elastic Waves by Laser Ultrasonics

293

Fig.9 Propagation images of the ultrasound propagating around a slit and fatigue crack

Fig.10 Setup for the visualization of

Fig.ll Visualized images of the ultrasonic

ultrasonic waves emitted from

waves emitted from cylindrical

cylindrical focusing transducer under

transducer under water.

water.

The next example is the visualization of ultrasonic waves emitted from a focusing transducer. Figure 10 shows the setup for the visualization of ultrasonic waves emitted from cylindrical focusing transducer under water. The amplitude image and arrival time image are shown in Fig.ll. From these images, we can easily understand the focusing configuration of the transducer. Figure 12 shows the propagation images of the ultrasounds propagating on the surface of both Si3N4/$25C

and Cu/S25C specimens. Cylindrical samples of both Si3N4 and $25C

were embedded in $25C carbon steels. From these images, we can predict the propagation behavior of the ultrasonic waves in metal matrix composites containing particulate dispersed particles. Propagation images of the ultrasonic waves propagating around blowholes of welding steel are shown in Fig.13. As this steel have a large ad small blowholes, the complicated

J. Takatsubo, M. Imade and S. Yamamoto

294

Fig.12 Propagation images of the ultrasound propagating on the surface of both Si3N4/$25C and Cu/S25C specimens.

Fig.13 Propagation images of the ultrasounds propagating around blowholes in a welding steel.

scattering images are observed around the defects. Figure 14 shows the propagation images of the ultrasounds propagating around circular grooves with different depths. An angle beam transducer was attached on the aluminum specimens, and the ultrasonic waves propagating on the side surface were visualized. We can observe the difference of sound velocity dispersion due to the change in groove height.

Visualization of Elastic Waves by_Laser Ultrasonics

295

Fig.14 Propagation images of the ultrasounds propagating around circular grooves with different depths.

5. C O N C L U S I O N S The results obtained from this work are summarized as follows: (1) Noncontact ultrasonic imaging system was developed for the visualization of ultrasonic waves propagating on opaque solid media. (2) This imaging system provides various kinds of visualization images such as propagation image, amplitude image, arrival time image and velocity image. (3) By use of this visualization method, we can produce a series of successive images as an animation of wave propagation. (4) This

technique

is

available

for

nondestructive

inspection

and

materials

characterization. REFERENCES (1) Rose J. L., Ultrasonic waves in solid media, Cambridge university press, 1999, p.396 (2) Thurston R. N. and Pierce A. D., Ultrasonic measurement method, Academic press, Inc., 1990, p.292 (3) Takatsubo J., Imade M., Fan Q. and Yamamoto S., Visualization of elastic waves by digital laser ultrasonics (in Japanese), Vol.65 No.639, 1999, pp4299-4304. (4) J. Takatsubo and

S. Yamamoto, Propagation mechanism of ultrasonic waves in

porous ceramics, JSME Int. J., Vol.39, No.2, 1996, 266-71. (5) Q. Fan, J. Takatsubo and S. Yamamoto, Quantitative characterization of advanced porous ceramics based on a probabilistic theory of ultrasonic wave propagation, J. Appl. Phys., Vol.86, No.7, 1999, 4023-28.

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

297

ULTRASONIC DETECTION USING P H O T O R E F R A C T I V E MULTI QUANTUM W E L L

H.HIRATSUKA, M.YAMAMOTO, H.IKEDA, S.TSUJIKURA and Y.OKAMOTO

Opto-Mechatoronics Dept., Advanced Technology Laboratory, KUBOTA Corporation 1-1 Hama 1-chome, Amagasaki City, Hyogo Japan Y.MATSUDA, H.NAKANO and S.NAGAI

National Research Laboratory of Metrology 1-1-4, Umezono, Tsukuba, Ibaraki, 305-8563, Japan

ABSTRACT Photorefractive multi quantum well, which will be applied to laser ultrasonic testing, has been developing. One of major problems of laser ultrasonic testing is the detection. To detect ultrasonic wave, a laser is illuminated to the object, but the reflected laser wavefront is distorted by scattering on rough surface or fluctuation of atmosphere. Photorefractive multi quantum well has ability to compensate the distorted wavefront with very high responsibility. Two kinds of processing techniques of photorefractive multi quantum well is presented. One is the implantation of proton and the other is the growth at low temperature (around 300~ To evaluate the photorefractive effect of multi quantum well, two-wave mixing efficiency is measured, and maximum two-wave mixing efficiency of over 20 % is obtained. Then compensation of distorted wavefront using photorefractive multi quantum well is demonstrated. Furthermore ultrasound caused by a transducer or pulsed laser is detected using the proton-implanted multi quantum well. KEYWORDS Photorefractive effect, Multi quantum well, Wavefront compensation, Two-wave mixing.

INTRODUCTION Laser ultrasonic testing is attractive because both generation and detection of ultrasonic waves are done by laser without making contact with objects[ 1,2]. However one of major problems of laser ultrasonic testing is the detection. To detect ultrasonic wave, a laser is illuminated to the object, but the reflected laser wavefront is distorted by scattering on rough surface or fluctuation of atmosphere. To solve this problem, a number of laser ultrasonic receivers based on photorefractive two-wave mixing interferometer have been developed in recent years[3-5]. Photorefractive materials have ability to compensate the distorted wavefront. But the response time of most of photorefractive materials is too slow to work in a practical use. For example, when the object is moving or vibrating, the system which is using those photorefractive materials can not detect the ultrasound. On the other hand, because of large carrier mobilities, semiconductor materials are one candidate of photorefractive materials for a fast response application. Especially, multiple quantum well (MQW) which contains A1GaAs/GaAs2, or InGaAs/GaAs3 system of compound semiconductor, is expected as a photorefractive material with fast response time of/~s order.

298

H. Hiratsuka et al.

Photorefractive multiple quantum wells have been developing with fast response time[6-8]. To fabricate the photorefractive devices from the above materials, it needs to introduce deep level impurities into the MQW where the construction of internal electric field enough to occur the photorefractive effect exists. There are two ways to introduce the deep level impurities. One is a Low temperature MBE growth technique. Another isproton implantation after growth. In this report, condition of proton implantation for obtaining high resistance A1GaAs/GaAs MQW and photorefractive property for the transverse geometry in the plane of proton-implanted A1GaAs/GaAs MQW is investigated. And to evaluate the photorefractive effect of multi quantum well, two-wave mixing efficiency is measured. Then compensation of distorted wavefront using photorefractive multi quantum well is demonstrated. Furthermore ultrasound caused by a transducer or pulsed laser is detected using the proton-implanted multi quantum well. DEVELOPMENT OF PHOTOREFRACTIVE MQW Photorefractive multi quantum well (MQW) devices that work at practical response time with the low light levels returned from the rough surfaces has been developing. This device acts as adaptive optics to compensate wavefront distortions through the photorefractive effect derived from Quantum Franz-Keldysh effect. The devices used in our experiments were grown by molecular beam epitaxy on semi-insulating GaAs substrates. The active photorefractive layers consisted of a multiple quantum well layer composed of a 60 periods superlattice of 8 nm GaAs wells and 10.0 nm A10.3Ga0.7Asbarriers. In order to make MQW's semi-insulating and introduce photorefractive center, we studied two different kinds of methods. One is the growth at normal temperature (600~ followed by proton implantation and the other is the growth at low temperature ( - 300~ for MQW layers. The multi quantum well structure was grown on a AlAs sacrifice layer. The MQW film was removed from GaAs substrate using a lift-off technique. The lift-off MOW film has an area of 2x4 mm 2, and was bonded on glass slide by van der Waals bonding. Au-Ge coplanar contacts 2mm long with a gap of lmm were evaporated to apply transverse Franz-Keldysh geometry. Finished samples are shown in Fig.1 The absorption spectra and the electroabsorption ratio Ao~/o~as a function of the wavelength was measured. The electroabsorption ratio approaches 12 % at wavelength of around 845nm that corresponds to quantum confined exciton. We observed a large electro-absorption of the exciton that is derived from Quantum Franz-Keldysh effect. The changes in the absorption spectrum are accompanied by changes in the refractive index in the material through the Kramers-Kronig transformations. The fractional change in index An/n generally approaches 1%. These large changes in the absorption and index are critical ingredients for large photorefractive effect.

Fig. 1 Photograph of photorefractive MQW.

Fig.2 Schematic of the two-wave mixing experiment.

Photorefractive Multi Quantum Well

299

TWO-WAVE MIXING EFFICIENCY OF PHOTOREFRACT1VE MQW We performed two-wave mixing (TWM) and measured TWM efficiency to evaluate the photorefractive effect for some MQW devices by proton implantation and low temperature growth. The experimental geometry is shown in Fig.2. The measurements are performed by monitoring the intensity modulation of one transmitted beam as the other beam is mechanically chopped. External-cavity tunable diode laser is used as an optical source in the range 835 nm and 855 nm. A dc electric field of "~20 kV/cm is applied.

Proton implantation We investigated the optimum conditions of the dose amount of proton at 150 keV. The wavelength dependence of TWM efficiency is shown Fig.3. Because of the excitonic enhancement of the photorefractive effect, large TWM efficiency is obtained around the excitonic resonant wavelength. There is an optimum dose amount of proton. In our experiment, the sample at a dose of 1012/cm2 shows best TWM efficiency. The maximum TWM efficiency reaches up to 0.25

0.2

o'f "~176

0.15

*~

=,, 0.15 ._ 0.1

0.1 0.05

lao

A

4~

0.05 or)

0

I-I d,

-0.1

-0.15

1.00E+12 2.00E-1-12

0

~ -0.05 ~ -0.1

"=~ -0.05

-0.2

I

- - B - - 2~

\

-0.15

~'~'~

-0.2 835

840

845 wavelength

850

-0.25

855

835

840

(nm)

Fig.3 Two-wave mixing efficiency of proton implanted samples as a function of laser wavelength with the doze density of 5 • 1011 to 2 X 1012cm-2. 12% at 847nm. Both 5x1011/cm 2 and 2x 1012/cm2 dose make the performance lower. Proton implantation introduces defects in MQW and the defects work as photorefractive center. So it is assumed that a lot of defects cause a large photorefractive effect. But if proton is overdosed, MQW is destroyed and excitons disappear. Therefore no photorefractive effects occurs. The wavelength dependence of TWM efficiency with the doze of 1012/cm2 for some doze angles is shown Fig.4. The sample at a dose angle of 2 o shows best TWM efficiency. The maximum TWM efficiency reaches over 20 % at 848 nm. This is probably occurred by the fact that the profile of doze ion depends on the doze angle.

845 wavelength (nm)

850

855

Fig.4 Two-wave mixing efficiency of proton implanted samples as a function of laser wavelength with the implantation angle of 0 to 6 ~ . 0.1 0.08

e-

V/lll=5

0.00

o

v,,,,-,0

l [ !

0.0,

e-

x

0

-0.02

~

L

i

~'-

~> -o. 04

-0.06 -0.08 -0.1

8 5

840

845 wave length

(nm)

850

855

Fig.5 Two-wave mixing efficiency of lowtemperature-grown samples as a function of laser wavelength with the V/III flux ratio of 5 to 30.

300

H. Hiratsuka et al.

Low Temperature Growth The optimum conditions of low-temperature-growth are also examined. Photorefractive MQWs with various V/Ill flux ratio at 300~ were prepared. TWM efficiency is shown in Fig.5. All the TWM efficiencies are poor relative to that of proton implantation experiments. In case of MBE growth at standard temperature, V/Ill flux ratio is usually maintained around 30 for creating a stoichiometry of materials. In case of low-temperature-growth, excess As are precipitated and those work as photorefractive center. Then a lot of As precipitates are assumed to be good for photorefractive effect. According to the decrease of V/Ill flux ratio, As precipitates are also assumed to decrease and cause poor photorefractive effect. However, the results obtained are contradictory. The less As precipitates produced, the more photorefractive effects occurred. In the experiment, over-excess As may destroy the superlattice and no quantum effect may exist.

ULTRASONIC DETECTION WITH WAVEFRONT COMPENSATION The experimental setup is shown in fig.6. Si substrate in which the ultrasonic vibration of the frequency of 0.85 MHz is generated by the transducer is used as the object. The MQW by proton implantation is used as the photorefractive device. The polarized beam-splitter and wave plates equalize the signal beam intensity to reference one. After two-wave mixing, the signal beam is detected by S i photo detector and amplified. The signal is observed on the digital oscilloscope with 64 times averaging. Detected signals for plane wave and wavefront distorted by a flame of lighter are compared. First, fig.7 shows the ultrasonic signal detected by Michelson interferometer. In comparison between fig.7(a) and (b), when the wavefront of signal beam is distorted, the amplitude of detected signal is reduced. Next, fig.8 shows the 4

Fig.6 Schematic diagram of ultrasonic detection based on two-wave-mixing.

....................................................~ .................................................................................................................

4

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

> 2 E

v

.i.a

0

E t~ --

-1

tu)

03- 2

"

~3 -4

L]

i

05 - 2

|

. . . . . . .

............................................................................................................................................................................................. 4 ...............................................................................................................................................................................

~

-1 0

0

10 Time (/1 s)

20

-1 0

0

10 Time (/1 s)

(a) (b) Fig.7 Ultrasonic signals detected by Michelson interferometer. (a) Detected signal for plane wave. (b) Detected signal for distorted wave.

20

Photorefractive

Multi Quantum

301

Well

ultrasonic signal detected by two-wave mixing using photo-refractive MQW. In comparison between fig.8(a) and (b), when the wavefront of signal beam is distorted by a flame the amplitude of detected signal is not reduced, so it is found that the wavefront compensation successfully works. Furthermore frequency distribution of ultrasonic signal detected by Michelson interferometer is shown in fig.9. The solid and dashed curves are for plane wave and distorted wave respectively. In the case that the wavefront of signal beam is distorted the constitute of ultrasonic signal to detect (0.85 MHz) is reduced, and the constitute of low frequency caused by fluctuation of atmosphere by a flame comes out. On the other hand, as shown in fig.10, frequency distribution of 4 ..............................................................................................................

4 ..........................................................................................................................................................................................................

.............. 1

"o9

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

:

...........

-~-1

i

................................. 9

1

"1o

i

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

i

...........

E

E

.~_ oo - 2

.

.

.

.

.

.

.

.

.

.

.

.

.

.~_ co - 2

.

-3

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

-3

-4

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

0

10

'

4 i.....................................................................................................................................................................................................

20

-10

0

Time ( # s)

10

20

Time (/.t s)

(a) (b) Fig.8 Ultrasonic signals detected by two-wave mixing using photorefractive MQW. (a) Detected signal for plane wave. (b) Detected signal for distorted wave. 0 .................................................................................................................................................................................. -10

-10

133 "o

~- - 2 0 E

E

x_

:

I

:

i

:

i

o -30

Q. r

o

:

I;

-20

~, -30 Q. ;',

~-40 o

-40

. . . .

':

,,

,

'

.

'

.t" r, .

i ~,~..

i

a.

n

-50

-50

-60

-60 0

0.5

1.5

2

Frequency (MHz)

Fig.9 Frequency distributions of ultrasonic signal detected by Michelson interferometer of received ultrasonic signals. The solid and dashed curves are for plane wave and distorted wave respectively.

0

,

r

i

i

0.5

1

1.5

2

Frequency (MHz)

Fig.10 Frequency distribution of ultrasonic signal detected by two-wave mixing using photorefractive MQW. The solid and dashed curves are for plane wave distorted wave respectively.

302

H. Hiratsuka et al.

detected signal is not change between these cases that the wavefront of signal beam is plane and distorted. Because of wavefront compensation, low frequency noise caused by a flame is removed. The signal peak (0.85 MHz) to maximum sidelobe ratio, i.e. signal to noise ratio, is over 10 dB. This value almost matches one by Michelson interferometer. So, it is found that the photorefractive MQW successfully works as the wavefront compensator. LASER ULTRASONIC DETECTION BY TWO-WAVE MIXING In the laser ultrasonic testing, not only detection but also generation of ultrasonic should be done. Then the detection of laser ultrasonic generated by pulsed laser is demonstrated. The experimental setup is shown in fig.ll. A single crystal of Si of 10 mm thick is used as work piece, in which the ultrasonic vibration is generated by the pulsed laser (Nd:YAG Laser, wavelength: 532nm, output power: 155mJ, pulse width: 6"-'7 ns, repetition frequency: 10Hz). Received signal by the system is shown in fig.12. Two peaks of the direct longitudinal wave and the reflected echo are observed. The time interval between these two peaks is 2.368/ts, so it is lead that the speed of sonic in Si is 8446 m/s. As compared with the characteristic value of Si (8433 m/s), the error is 15.35 /zm of thickness. Moreover about 1.5 nm of the surface displacement caused by the reflected echo is detected. CONCLUSION

Fig.ll Schematic diagram of laser ultrasonic detection based on two-wave mixing. 400

........................................................................................................................................................................................................................... I ! i ISample : A single crystal Siof 10mm thick I

300

.....

200 .~

i- The direct longitudinal wave

-i . . . . . . . . . .

/

. . . . i. . . . .

//i

.......

i. .

. . . . . . .

i

!.. The reflected ech . . . . . . .

100

. . . . . . . . .

^

' ~ - . - ~ I ~ i V ~v~/v~m/v~, v - " -

0 t~

-~ - 1 0 0

.

.

.

.,

,,,4

4

5

.

b3 - 2 0 0 -300 -400

.. 0

9 1

2

3 Time (/1 s )

Fig.12 Received signal using ultrasonic receiver based on two-wave mixin~ in photorefractive MQW.

Two kinds of processing techniques, the implantation of proton and the growth at low temperature, of photorefractive multi quantum well have been presented. To evaluate the photorefractive effect of multi quantum well, two-wave mixing efficiency is measured, and maximum two-wave mixing efficiency of over 20 % is obtained by the proton-implanted MOW. Then compensation of distorted wavefront using photorefractive MQW is demonstrated and signal to noise ratio of over 10dB has been obtained.. Furthermore ultrasound caused by a pulsed laser has detected by TWM using photorefractive MQW. This work was supported by the program "Advanced Photon Processing and Measurement Technologies", contracted to the R&D Institute for Photonics Engineering from New Energy and Industrial Technology Development Organization (NEDO) of Japan.

Photorefractive Multi Quantum Well

303

REFERENCES H. Nakano, Y. Matsuda, and S. Nagai, "Ultrasonic velocity measurements in molten materials with the use of laser-generated ultrasound", Meas. Sci. Technol. 9 (1998) 617621. H. Nakano, Y. Matsuda, and S. Nagai, "Ultrasound detection by using a confocal FabryPerot interferometer with phase-modulated light", Ultrasonics 37 (1999) 257-259. A. Blouin and J. P. Monchalin, "Detection of ultrasonic motion of a scattering surface by two-wave mixing in a photorefractive GaAs crystal", Appl. Phys. Lett. 65 (1994) 932-934. B.F.Pouet, R.K.Ing, S.Krishnaswarmy and D.Royer, "Heterodyne interferometer with twowave mixing in photorefractive crystals for ultrasound detection on rough surfaces", Appl.Phys.Lett.69 (1996) 3782 M.B.Klein, G.D.Bacher, A.G-Jespin, D,Wright and W.E.Moerner, "Homodyne detection of ultrasonic surface displasements using two-wave mixing in photorefractive polymers", Opt.commun. 162 (1999) 79-84. Q.Wang, R.M. Brubaker, and D.D.Nolte and MR. Melloch, "Photorefractive quantum wells: transverse Franz-Keldish geometry", J.Opt.Soc.Am.B/Vol9, 1626 (1992) S.Iwamoto, H.kageshima, T.Yuasa, M.Nishioka, T.Someya, Y.Arakawa, K.Fukutani, T.Shimura, and K.Kuroda, "Photorefractive four-wave mixing in InGaAs/GaAs multiple quantum wells", MOC'99 proceedings, 94 (1999). I. Lahiri, L.J.Pyrak-Nolte, D.D.Nolte,M.R.Melloch, R.A.Kruger, G.D.Bacher and M.B.Klein, "Laser-based ultrasound detection using photorefractive quantum wells", Appl. Phys. Lett. 73 (1998) 1041-1043.

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COMPACT AND ROBUST INSPECTION SYSTEM FOR MICRO CRACKING DETECTION USING LASER-INDUCED SURFACE WAVES

M. OCHIAI, N. MUKAI, Y. SANO and

Power and Industrial Systems Research and Development Center, Toshiba Corporation 8, Shinsugita-cho, Isogo-ku, Yokohama 235-8523, Japan H. NAKANO

Material Measurement Section, National Research Laboratory of Metrology 1-1-4, Umezono, Tsukuba 305-0045, Japan

ABSTRACT Laser-induced surface acoustic wave (SAW) inspection based on a modified confocal Fabry-Perot interferometer (CFPI) is proposed. In most cases of the inspection for in-service structure and apparatus, it is necessary to make the inspection device compact and to perform robust measurement in the case of various shapes and surface conditions of the materials to be inspected. For these purposes, a prototype system using the CFPI as a detector of SAWs and a compact probe (~20mm*25mm) using a multi-mode optical fiber (core diameter: 400~tm) is developed. Moreover, the irradiation of laser pulses for generating surface waves and the bias voltage applied to the photo-detector are synchronized with the scanning of the distance between the mirrors in the CFPI in order to obtain good signal-to-noise ratio under any circumstances. As a result of the modifications, the system is applicable to the inspection of a simulated in-service structure. Twodimensional visualization results of artificial slits having depths of 0.5mm to 1.5mm machined on a welded joint and typical waveforms detecting stress corrosion cracking (SCC) are presented. KEYWORDS laser ultrasonics, surface acoustic wave, non-destructive evaluation, ultrasonic testing, confocal Fabry-Perot interferometer, optical fiber, reactor internal inspection, stress corrosion cracking

INTRODUCTION Regarding aged nuclear power plants, confirmation of the material degradation of irreplaceable structures, such as reactor pressure vessels and reactor internals, has become a matter of great concern. In particular, in order to estimate initial cracking behaviors and to establish cracking growth predictions, detection of micro cracking having a depth of lmm or less is important. Moreover, since the materials to be inspected are installed in narrow spaces in nuclear reactors and have non-uniform surface roughness, it is necessary that the inspection method allow not only remote detection of small cracking but also robust inspection of various surface conditions with a compact probe. The laser ultrasonic technique [1] has been developed during the last two decades and it has been shown that small slits having a depth of the order of 100~tm are detectable with laser-induced surface acoustic wave (SAW) [2,3,4]. It has also been demonstrated that laser ultrasonic testing using confocal Fabry-Perot interferometer (CFPI) is capable of remote inspection of a rough surface [5,6].

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In this paper, results of two major developments based on these state-of-the-art technologies are discussed. Firstly, the basic design and operation of the developed detection system is described. The system consists of a compact probe using optical fiber to realize inspection in a narrow space and a modified CFPI to improve signal-to-noise ratio under various conditions. Secondly, performance of the developed system confirmed through some feasibility tests is introduced. Twodimensionally visualized results of artificial slits and typical waveforms detecting actual cracking are presented.

DESIGN AND OPERATION OF A PROTOTYPE SETUP

Modified CFPI A conventional CFPI consists of two spherical mirrors having a high reflectivity arranged with their concave surface facing each other and separated by a distance, h, nearly equal to their curvature radius, as shown in Fig.l(a). The optical arrangement well known as an "etalon" has a sharp spectral response, as illustrated in Fig.l(b). The procedure of measuring small ultrasonic vibration using the CFPI is as follows: 1) A single-mode laser beam is irradiated onto (a) mirror mirror a vibrating surface. / 7 . . . ................ ~ ,o a p~o,o-~o,o~,o, incident laser-be;m--~ ~.............. - ' - - ~ ......... 2) The original frequency, fo (=c/A,o, c: light < h > PZD F-~edbackSignal velocity, A,0: light wavelength), of the reflected or scattered laser beam from the (b) surface is modulated to fo +--fa due to the Transmission l[ ~ ('~'~ Doppler effect. ,o+,~--~--~- ......... l:-,:--,-~,::-~-!~:~. 3) The frequency-shifted laser beam .............t ........................... ,:............. ~o-~......V_...._V...... _ ~_........,-~,/-.i-,.-. illuminates the CFPI maintained at point A in Fig.l(b). . [ 4) The output of the CFPI, which contains the Frequency of light signal, I, proportional to the velocity of the fo'fdfofo+fd surface as illustrated in Fig.l(b), is detected by a photo-detector. Fig.1 (a) Basic arrangement of the CFPI and In addition, to stabilize the separation of the (b) an example of a spectral response curve CFPI strictly at point A in Figl(b), an active control system based on the feedback control using a piezoelectric displacer (PZD) is incorporated. In some cases, however, two issues arise regarding this optical system. One is that it is difficult to stabilize the separation, h, under severe vibration circumstances. The other is that the inevitable continuous D.C. light not only increases thermal noise at the photo-detector but also decreases sensitivity of the photo-detector by decreasing its bias voltage. To deal with these issues we introduce some modifications into the conventional CFPI. Referring to Fig.2, an exaggerated timing chart, the procedure of measuring small ultrasonic vibration using the modified CFPI is explained as follows: 1) The piezoelectric displacer sinusoidally scans a spherical mirror a certain distance longer than one wavelength of the light, A0, as illustrated in Fig.2(a). 2) According to the separation, h, change, the reference signal, iref, of the modified CFPI behaves as illustrated in Fig.2(b). 3) A trigger signal is transmitted at the output signal level corresponding to the point A in Fig.l(b). 4) Emission of the laser pulse for generating ultrasonic wave is synchronized to the trigger signal as illustrated in Fig.2(c). 5) The bias voltage, liB, is applied to the photo-detector intermittently and it is also synchronized to the trigger signal with a delay of z'D and a duration of rM as illustrated in Fig.2(d). Owing to these modifications, the stable measurement of the ultrasonic signal, isig in Fig.2(e), can be performed without any control of the separation even under severe vibration circumstances.

307

Micro Crack Detection Using Laser Ultrasonics

(a) . . . . .

(b)

l! *t

........................... --I~i! ~ - -~ii

~

T_rg..__L_.e.y__e. .............................. .!_ ~ii ~'t '

(c) i! ~-t

n

(d)

i

n

iE.t ii

"t

Fig.2 Timing chart showing the procedure of the modified CFPI Because the scanned mirror passes though the best position for the vibration measurement at least twice per cycle and the generation and detection of ultrasonic waves are completely synchronized to the moment. Moreover, when the constant D.C. light, Io, is introduced to the photo-detector, the bias voltage, VB, is automatically deduced to prevent the over current. The modified CFPI allows application of the proper bias voltage because the measuring duration, z'M,is much shorter than the repetition rate, z'n, of the pulsed laser emissions. Therefore, the sensitivity of the photo-detector is improved by applying the proper bias voltage.

Compact probe Remote and non-contacting measurement of ultrasonic waves is one of the most obvious advantages of the laser ultrasonic technique. However, in some cases of reactor internals inspection, it is more important to make the inspection probe compact using an optical fiber because there is little space around materials to be inspected and few routes to propagate the laser beams. Therefore, we designed and assembled a compact probe using an optical fiber experimentally. The probe shown in Fig.3 has a diameter of 20mm and a length of 25mm. The designed lift-off of the probe is 30mm. The probing laser beam is delivered to the probe TriggerSignal

Fig.3 Appearance of the compact probe

Q-switched Nd: YAG

I

Scanning Stage Optical flat (0.9m x 0.6m) Trigger ....~enerator_.

cw

PIeD v~__----~ Supply [._.u~.p...!..y.j Test Piece

APD Filtel PZT Displacer

isig t Amplifier Filter- I I oscillat~

rSeiga~dler_]

Fig.4 Schematic diagram of the prototype setup

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M. Ochiai et al.

via a multi-mode optical fiber having a core diameter of 4001xm and a length of 60m and focused onto the surface. The reflected light from the surface is led to the CFPI through the same optical path. A schematic diagram of the prototype setup is shown in Fig.4.

TEST RESULTS Visualization of artificial slits The prototype setup is used to investigate basic 1.5mr mm Crack performance of laser-induced SAW testing with artificial slits. Simulated in-service material is a Detection Las~ welded stainless steel plate and it contains three machined slits having a width of 0.1mm and depths Compact Prot of 0.5mm, 1.0mm and 1.5mm, respectively. When a laser pulse from a Q-switched Nd: YAG laser Laser Pulses S t a i n l e s s S t e e l Plate having a wavelength of 1,064nm and an energy of Incident SAW 3 [ Echo sisnals about 15mJ/pulse is launched onto the surface at one side of a slit, a SAW is excited due to ablative t[ ~ slit depth: 1.5mm ............. !t........t~..... Jt........................ interaction between laser and material. The excited ~r : slit depth: 1.Omm SAW and its echo signal reflected from the slit are detected as micro surface vibration using the slit depth: 0.5ram modified CFPI with a continuous wave Nd: YAG 0 ~ .............. "'-'"~" r:-....... !v .......... , . . . . . . . , laser source having a wavelength of 532nm and a power of about l W. The probing laser beam delivered with the multi-mode optical fiber is 0 5 10 15 20 25 30 focused onto the surface at the same side of the slits Time (~tsec) with intervals of 5mm-10mm from the slits and of Fig.5 Typical waveforms of the echo about 30mm from the point of SAW excitation. signal from artificial slits having The signal from the interferometer is processed with an analogue filter having a frequency-band of 1-20MHz and is converted to a digital waveform consisting of 5,000 data points using a digital oscilloscope with 250Msample/sec. Each waveform is obtained with signal averaging over 4 shots to reduce noise. Trigger signal synchronized with the laser irradiation is also fed from the Q-switched laser source in order to identify the accurate time of SAW generation. Typical waveforms detected with the developed system are shown in Fig.5. The first wave appearing at about l l~tsec is the incident SAW traveling from the sound source to the detection point. When there is a cracking on the material, other waves generated due to SAW/slit interaction are observed. While the inspection point on the material is scanned, the signal amplitude at each position is converted into colors and displayed. A visualized result for the slits is shown in Fig.6.

(a) Appearance (b) Test result Fig.6 Comparison of (a) actual appearance of the test piece and (b) a result of slits visualization

309

Micro Crack Detection Using Laser Ultrasonics

Detection of stress corrosion cracking Stress corrosion cracking (SCC) is one of the most difficult flaws to detect because most SCC has a very narrow aperture of the order of 10~tm and a complicated and branching shape. To confirm the inspection performance to the SCC, feasibility experiments are performed on type 304 stainless steel plates having real SCC. Figure 7 shows an appearance, an enlargement and a result of liquid penetration testing of typical tested material. The SCC Fig.7 (a) Appearance, (b) enlargement and (c) result of has a width of about 20-100~tm penetration testing of the test piece and a depth of 1-3mm. The interval from the irradiation *---Incident S A W >. point of laser pulses (sound " ~ 0.5 E c h o signal source) to the detection point is 2 about 30mm and from the detection point to the SCC is ,--' -0.5 about 2-7mm. -1 i i , i i Typical waveforms detected with 8 10 12 14 16 18 20 the developed system are shown 1 ~" ~---Incident S A W in Fig.8. The first waves on each 0.5 ......................................................................... figure are the incident SAWs. E c h o signal ,,.., The second ones are identified as the echo from the SCC though an 0 estimation using the time-offlight (TOF) method.

.

i. ii t

-1

8

DISCUSSION

10

12

14

16

18 20 Time(gsec)

Fig.8 Echo signal detected at different positions on SCC

During the feasibility experiments, the developed system shows a sufficient stability without the feedback control even under vibration circumstances. As the previous studies described, the artificial slits having a depth of 0.5mm or less are detectable with the conventional CFPI. However, the signal-to-noise ratio is considerably improved with the modified CFPI because the modification enables application of the desired bias voltage to the photo-detector during the measurement. These results suggest the possibility not only of detecting cracking having smaller depth or more complicated shape, but also of performing robust inspection in the case of various surface conditions by using the modified CFPI. The designed lift-off of the compact probe is 30mm; however, it depends on the surface conditions, the probe diameter and the incident angle of probing beams. In some cases, a lift-off of 100mm or more would be possible using a similar probe. On the other hand, the delivery of laser pulses for generating SAW is not discussed in this paper, because we have already completed the development of 200mJ/pulse laser pulses delivery system for nuclear reactors (Laser Peening System [7,8]). Moreover, a fiber delivery technique of 100mJ, 5nsec laser pulses using a beam homogenizer has also been developed [9,10]. An example application of this study is to perform reactor internals inspection by mounting the developed probe on the delivery systems as shown in Fig.9.

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Fig.9 An example of application of the developed system

CONCLUSION In order to apply laser ultrasonic testing to the inspection of nuclear reactors, we proposed a detection system using a compact probe with multi-mode optical fiber and a modified CFPI. It is found that the developed system enables well-stabilized measurements to be obtained even under severe circumstances and a sufficient signal-to-noise ratio on rough and non-uniform surface conditions. Micro slits having 0.5mm depth machined on welded metal and some actual cracking, SCC, are observable with the developed technique.

REFERENCES

1. 2. 3. 4. 5. 6. 7. 8. 9. 10.

Scruby, C. B. and Drain, L. E. (1990). Laser Ultrasonics: Techniques and Applications. Adam Hilger, Bristol. Cooper, J. A., Crosbie, R. A., Dewhaust, R. J., Mckie, A. D. W. and Palmer, S. B. (1986). IEEE Trans. UFFC, UFFC-33, 462 Cooper, J., A., Dewhaust, R., J. and Palmer, S., B. (1986). Phil Trans. R. Soc. Lond., A320, 319 Shan, Q. and Dewhaust, R., J. (1993).Appl. Phys. Lett., 62, 2649 Monchalin, J. P. (1985). Appl. Phys. Letters, 47, 14 Monchalin, J. P. and Heon, R. (1986). Materials Evaluation, 44, 1231 Sano, Y., Kimura, M., Mukai, N., Yoda, M., Obata, M. and Ogisu, T. (1999) Proc. Int. Forum AHPLA '99, SPIE3888, Osaka, 33 Sano, Y. et al. (2000). Proc. of 8th International Conference on Nuclear Engineering, ICONE-8441, Baltimore Schmidt-Uhlig, T., Karlitschek, P., Yoda, M., Sano, Y. and Marowsky, G. (2000). Eur. Phys. J. AP, 9, 235 Yoda, M., Sano, Y., Mukai, N., Schmidt-Uhlig, T. and Marowsky, G. (2000). The Review of Laser Engineering, 28, 309 [in Japanese]

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

311

DETECTION OF DEFECTS IN MICRO-MACHINE ELEMENTS BY USING ACOUSTIC WAVES GENERATED BY PHASE VELOCITY SCANNING OF LASER INTERFERENCE FRINGES

H. Sago l, S. Matsumoto l, S. Nakano 1, H. Ogiso 2'1, H. Cho 3 and K. Yamanaka 3

1Mechanical Engineering Laboratory, 1-2 Namiki Tsukuba, Japan 2National Institute for Advanced Interdisciplinary Research, 1-1-4 Higashi Tsukuba, Japan 3Graduate School of Engineering, Tohoku University, Sendai, Japan

ABSTRACT High aspect-ratio micro structuring is a key process for making micro electromechanical systems (MEMS). Many micro fabrication methods are being developed (e.g., anisotropic etching processes and LIGA processes). Among them, the inductively coupled plasma (ICP) source etching is especially attractive because it can make high aspect-ratio micro structures at high speeds and because it requires comparatively small facilities. However, in such micro structures it is difficult to nondestructively evaluate the quality of high-aspect-ratio etched bottoms. Therefore, we developed a technique for detecting defects on high aspect-ratio micro structures in the Si wafers machined by using the ICP source etching. By using acoustic waves generated by phase velocity scanning of laser interference fringes, we nondestructively detected 10 t.tm order high defects located on the bottom of narrow and deep grooves from the other side of the Si wafer, which could not be detected by using optical techniques nor the scanning electron microscope. KEYWORDS laser-generated ultrasound, nondestructive evaluation, MEMS, subsurface defect, phase velocity scanning method, surface acoustic wave, Lamb wave, deep reactive ion etching.

INTRODUCTION For the fabrication of micro electromechanical systems (MEMS) on Si wafers, semiconductor manufacturing technology is mainly used, such as photolithography. To make MEMS devices, high aspect ratios are required. Recently, a deep reactive ion etching (DRIE) technology was developed that uses inductively coupled plasma (ICP).[1] DRIE technology alternately performs etching of Si using a high density plasma and deposition of side wall passivation. It is attractive because it achieves high aspect-ratio micro structures at high speeds by using small facilities. However, if either a small particle or a thin film is accidentally deposited on the bottom of the etched structure, projections or protrusions will be left there. Although this is a serious problem for making narrow and deep grooves, it is difficult to nondestructively detect small defects at the bottom of high-aspect ratio structures. To overcome this difficulty, we developed a sensitive technique based on the phase velocity scanning (PVS) method [2-4], which we developed before, to analyze the bottom of high-aspect etched grooves from the back side of the S i surfaces.

312

H. Sato et al.

RESULT AND DISCUSSION We made narrow grooves by using DRIE (Fig.l, Table 1) under Si (100) surface. The total etching time was approximately 4 hours. A tone burst of the surface acoustic wave (SAW), which propagates along the [110] (x) direction, was generated by the PVS system using the scanning interference fringes (SIF) formed by the intersection of the laser beams.[3,4] For the laser beam, we used the second-harmonic wave of a Q-switched Nd:YAG laser (9mJ, 70ns-pulse-width). The velocity of the SAW was 5083 m/s and the frequency was 70 MHz. To detect the generated and scattered waves, we used an optical knife-edge method with an Ar + laser (100 mW). The SAW generated by the SIF was detected by using an optical probe (signal V1), and was incident on the bottom of grooves (Fig.2). Then, SAW scattered by the groove was detected by using the same optical probe (signal V2). A video signal [5] was obtained by using a detector with a 20 MHz band-pass filter (Fig.3). The RF signal and the video signal were averaged 20 times. The distance between the SIF and the probe was approximately 6.2mm, and the distance between the probe and the probe-side edge of the bottom of the grooves was approximately 2 mm. In Fig.3, the signal V1 is the generated SAW, V2 is the SAW scattered by the probe-side edge of the bottom of groove B. Furthermore, two signals (V3, V4) were detected, which can not be scattered by edges of the bottom of groove B. Unidentified two signals were also detected from groove A (Fig. 4). To find the sources of scattered acoustic waves (V3, V4), we cut the grooves along the x direction, and examined the cross section by using SEM. Then, we found a 17-~mhigh projection approximately 5 mm away from the edge of groove A(Fig.5), and two 13-~m-high projections 1.3 mm and 1.6 mm away from the edge of groove B. Because we detected two signals V3 and V4 from groove A, but detected only one projection, we postulated that two wave modes existed on the groove, such as Lamb waves, and we estimated the velocities of V3 and V4 to be 8.3 x 103 rn/s and 5.6 x 103 m/s. By the same analysis for groove B, the velocities of V3 and V4 became 5.6 x 103 m/s and 2.6 x 103 m/s. For comparison, we calculated group velocities of anisotropic Lamb waves for the A0 mode and for the So mode and experimental velocities were found to be close to them. Then we conclude that V3 and V4 are scattered waves by a single projection. CONCLUSION We developed a technique for detecting defects on high aspect ratio microstructures, which machined by deep reactive ion etching, in Si wafers. Using acoustic waves generated by phase velocity scanning of laser interference fringes, we nondestructively detected 13-I.tm-high defects located on the bottom of narrow and deep grooves from the other side of the Si wafer.

Detection of Defects in Micro-l~lachine Elem~ents Table 1. Dimensions of grooves width length (rv-direction t (x-direction I Groove A 100 ~m 3.9 cm Groove B 100 grn 1 cm Groove C 100 I.tm 1 mm

313

thickness of

Si 32 ktm 48 l.tm 43 btm

Fig. 1. Schematic of a groove made on a Si (100) surface by using deep reactive ion etching..

Fig.2 Schematic of the experimental system for phase velocity scanning system of laser interference fringes, projection and mode-converted SAW.

Fig. 3 RF signal of a 70MHz generated wave (V 1) and scattered waves (V2, V3 and V4) from groove B.

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Fig. 4 B-mode image of video signals obtained by using 70 MHz SAW.

Fig. 5 SEM image of a cross section of groove A.

REFERENCES

1. Laemer, F., Schilp, A., Funk, K. and Offenberg, M., "Bosch Deep Silicon Etching: Improving Uniformity and Etch Rate For Advanced MEMS Applications", Proc. MEMS '99, Orlando, USA, January 1999, p.211. 2. Yamanaka, K., Nagata, Y. and Koda, T. (1991) AppL Phys. Lett. 58, 1591. 3. Nishino, H., Tsukahara, Y., Nagata, Y., Koda, T. and Yamanaka, K. (1993) Appl. Phys. Lett. 62, 2036. 4. Yamanaka, K., Kolosov, O., Nagata, Y., Koda, T., Nishino, H. and Tsukahara, Y. (1993) J. Appl. Phys. 74,6511. 5. Sato, H., Cho, H., Nishino, H., Ogiso, H. and Yamanaka, K. (1996) Jpn. J. Appl. Phys. 35, 3066.

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315

LOW FREQUENCY LASER ULTRASOUND U N D E R 100 K H z Kazushi Yamanaka, Hideo Cho, Teruyuki Nakazawa and Tomokazu Watanabe Department of Materials Processing, Graduate School of Engineering, Tohoku University, Aramaki, Aoba 02, Sendai, 980-8579 Japan E-mail [email protected]

ABSTRACT For nondestructive testing of large structures such as factory plants, buildings, tunnels and bridges, the vibration analysis in the frequency range of 1 kHz to 100 kHz is very useful, though automated and remote testing method has not been realized.

In this work, we

propose a laser ultrasonic method for this purpose, which has high efficiency in the low frequency range.

In the laser ultrasonics, the short pulse width (8-30 ns) of Q-switched

lasers is not optimum for the low frequency vibration. However, if we use normal pulse lasers, we suffer from the too long pulse width (>100/z s) and the spike-like relaxation oscillation. Then, we developed a new long pulse laser, a normal pulse YAG laser using a controlled discharge current, which achieves both relatively short pulse width and suppressed relaxation oscillation. The pulse had a rise time of 15 # s and a width of 50 /z s and a large energy of 2.5 J using a single laser rod. The advantage of this laser for nondestructive evaluation using low frequency vibration is theoretically analyzed and experimentally verified on bricks, steel joints and concrete plates. KEY WORDS Laser ultrasound, Normal pulse, Q-switched, Vibration, Ablation, Low frequency

1. INTRODUCTION Nondestructive testing of large structures such as factory plants, buildings, tunnels and bridges is very important, and rapid and reproducible testing is required. Since the ultrasonic testing using contact transducers and impact echo method using hammers [1] require access of a human operator, or a mechanical scanner, it is difficult to fulfil those requirements.

For this purpose, the laser ultrasonics [2,3] is most promising, because of its

non-contact nature.

In the laser ultrasonics, a Q-switched laser with a pulse width of 5 -

100 ns is usually employed to excite ultrasound, which is processed both in time domain and in frequency domain [3,4].

However, ultrasound in the frequency range of 1 - 100 MHz,

conveniently excited with a Q-switched laser is not optimum for testing large structures, because of the large attenuation.

In this respect, the normal pulse lasers with longer pulse

316

K. Y a m a n a k a et aL

width might be useful. However, if we use ordinary normal pulse lasers, the pulse width (>100/z s) is too long, the peak power is too small, and we are often disturbed by the relaxation oscillation characterized by many short pulses of tens of ns widths. Then, we try to optimize the excitation efficiency of low frequency vibration (1-100 kHz) by using a specially designed long pulse laser. The pulse width of the developed laser is around 50 tz s, which is much longer than that of Q-switched lasers, but shorter than common normal pulse lasers (longer than 100 tt s) and is not disturbed by the relaxation oscillation.

The advantage of nondestructive evaluation using this laser for low frequency

vibration, which we call laser impact (LI) method here, is theoretically analyzed and experimentally verified on bricks, steel joints and concrete plates. 2. PRINCIPLE 2-1 Physical Quantities To describe the principle of LI method, we need to review relevant physical quantities [3].

If a laser is switched on instantaneously at time zero and the absorption takes

place only at the surface, the surface temperature at time t is T ( t ) = T~ +--~--

(1),

where T~ is the initial temperature, F 0 is the absorbed laser power density assumed uniform across the irradiated area K is the thermal conductivity and tr is the thermal diffusivity (= K / ( p C ) ), where p is density and C is specific heat capacity of the object [3]. The temperature increases with square root time dependence and reaches the vaporization ~rKpC )2 temperature T v at time t v = 4Fo2 (T v - T~ , if we neglect the effect of melting.

this occurs only when the pulse width 9 is longer than t v .

Note that

Thus the ablation threshold for

absorbed laser power density is given by ~r = t v , which shows

&

=T(r~ -~)

(2).

It should be noted that the ablation threshold depends not only the material properties but also on the pulse width "r. A few examples of these quantities for focussed and unfocused long pulse lasers used in the LI method (A1 and A2) and for focussed and unfocused Q-switched lasers used in the laser ultrasonics (B1 and B2) are given in Table I. In the calculation, thermal conductivity K = 50 (W/m K), density p =7900 Kg/m3, specific heat capacity C--480

L o w Frequency Laser Ultrasound

317

J/(Kg K), evaporation temperature T V =3030 K and initial temperature T~=293 K were used, which are typical values in a mild steel.

To evaluate the absorbed power density, the optical

absorption coefficient was assumed to be 0.1, although it may become much larger in the ablation mode [3]. In the next sections, characteristics of the LI method and the ordinary laser ultrasonics are compared both in the thermoelastic mode and ablation mode. "l~pe of Laser beam

Beam Pulse Pulse diameter Energy w i d t h d , mm Q, J "t', It s (AI) Un-focussed 8 2.5 50 Long pulse (A2) Focussed 0.5 2.5 50 Long pulse (B 1)Un-focussed 4 0.1 0.01 Q-switched (B2) Focussed 0.5 0.1 0.01 Q-switched Table I. Type and operating condition for

Laser Power I 0, MW 0.05

Ablation Absorbed Vaporising Power density threshold time F 0 MW/cm2 Fth MW/cm2 t v , I t s 0.47 1.12• 106 0.0099

0.05

2.55

0.47

1.72

10

7.94

33.4

0.174

10

510

33.4

4.2X 10.5

typical pulse lasers

2-2 Thermoelastic Mode In cases (A1) and (B1), the absorbed power density is less than the ablation threshold, and hence they are in the thermoelastic mode.

However, the margin for the Q-

switched laser (B 1) is not large since the ablation threshold Fth (33.4 MW/cm2) is only 5 times that of the absorbed power density F o (7.94 MW/cm2). If we consider the melting, neglected here, a surface damage might take place.

In contrast, the margin for the unfocussed long

pulse laser (A1) is large enough, since F o /Fth <0.03, and perfectly non-destructive testing is realized.

Nevertheless, the total pulse energy Q of the long pulse laser is 25 times that of

the Q-switched laser.

Therefore, a significant advantage of the long pulse laser over the Q-

switched laser is expected for low frequency vibration excitation. This advantage will be quantitatively evaluated.

Assuming that the waveform of laser is

a half cycle of a sinusoidal wave with a half width of "r, it is expressed as 2xt I (t) = I 0 sin ~ 3z

3~" for 0 < t < D , 2

I (t) = 0

otherwise

where I 0 is the peak power.

(3),

The magnitude of frequency component of I(t) is obtained

by Fourier transform of eq. (3) as,

318

K. Yamanaka et al.

1 (3ft.) 2 cos

(4).

ft"

Note that the spectral component is proportional to the total pulse energy Q = 3 10t. at zero

20

I

'

,

'

'

'

'

' ' ' !___ long 50/z s 50 KW Q-sw. 10 ns 10MW

n't

~

-o

0

.

_

long 1 0 / z s 5 0 K W

______~

~

-

9-20

O (D. i,..

"

....

-40

-60

0

40

80

Frequency f, KHz Fig. 1 Spectral component of laser power in a long pulse and a Q-switched lasers. frequency (DC).

It shows the importance of pulse energy rather than the peak power for

efficient excitation of low frequency vibration. The amplitude decreases as the frequency increases and takes the first zero at the frequency f = lit.. It shows that the effective bandwidth is approximately given by the inverse of pulse width. In Fig. 1, the amplitude spectra of the lasers listed in Table I are plotted.

The solid

line is for the long pulse laser (Q =2.5 J and t. =50 ,u s which approximately gives I 0---0.05 MW) and the broken line is for the Q-switched laser B (Q=0.1 J and t. =10 ns which approximately gives I 0= 10 MW).

It is seen that the frequency component of the long pulse

laser in the frequency range less than 20 KHz is 20 dB larger than that of the Q-switched laser. Considering that the margin for the ablation threshold is much larger in the long pulse laser than in the Q-switched laser, it is a significant advantage as experimentally verified in section 3-2. Though the frequency component of the long pulse laser in the frequency range between 40 KHz and 80 KHz is less than that of the Q-switched laser, it can be enhanced by truncation of the pulse to t. = 10 /t s. Thus the frequency component becomes larger than that of the Q-switched laser as shown by the dotted line C.

Since the power density is not so

Low Frequency Laser Ultrasound

319

large, acoustooptic modulators can be safely used for this truncation. 2-3.

Ablation Mode Usually, the ablation mode, which damages a small area of the surface, is not

desirable for testing industrial components.

However, the ablation is sometimes acceptable

for large structures of concrete and bricks where the surface has already been damaged to some extent. focussed.

The long pulse laser can be used also for the ablation mode, when the beam is

In this case, it should be noted that the ablation threshold for the long pulse laser

A1 and A2 (0.47 MW/cm 2) in Table I is less than that for the Q-switched laser B 1 and B2 (33.4 MW/cm 2 ). Such a difference is due to the increase of pulse width "t" (0.01 It s to 50ls s), as shown by eq. (2).

In the long pulse laser, the vaporization of object surface

continues for a long time ('r -t v = 48.3 It s) and produces a long pulse of reaction force.

It

excites a low frequency vibration, much larger than in the thermoelastic mode. 3.

EXPERIMENTS

3-1

The Apparatus We developed a long pulse laser, a normal pulse YAG laser using a slow rise time fast

discharge circuit that achieves relatively short pulse width, without severe relaxation oscillation.

The waveform of the developed long pulse laser shown in Fig. 2 had a rise time

of 15 # s and a width of 50 /z s. laser head at 1064 nm.

A large pulse energy of 2.5 J was realized using a single

The low frequency vibration of objects is measured using a laser

Crack/ '

'

I

'

'

'

I

.

.

.

Long Pulse Laser 2.5J s o . s

7.

(50kW peak) ~- -

..~

....." , , ' ~

~/

iiI, +'~"+~0~r "~ \ / ~ . x , , , , , . ~ ' " ..... ,," LOW " I , ~" ~ .,~..~,'"' Frequency Interferomete~/------~ ,.,.,.,,~_, +.~"'~J Vibration ,

.i,.+

t-" .._1

Delamination

.

,

:>u/,c

++~

b

- -

o. .',~+~'`+ ,0'".jY' ..+.~ 0,~ +,

i

0 ,,+

. . . .

,+0 . . . .

Time t, /zs

Fig. 2 Waveform of long pulse laser

,v+.v vv+j^[,,.Spectrum) A

twaveformJ

Fig. 3

Vibration Mode

Schematic illustration of measurement system

K. Yamanaka et al.

320

-

I

E 50 ::t

. . . . i . . . . (a) Q-switch (Sns,532nm) 0.12 J

i

~1.

~

'

I

(a)Q-s'witch (8n;,532nm)'0.12 J 4

_

(b) Long pulse (50/zs,1064nm) 1.25 J

_

-50 I

i

|

i

i

o

1'0

,

'

Time t, ms

'

,

0

2'0

Fig. 4 Measured waveform on a brick by (a) Q-switched laser Co) a long pulse laser

5

10

15

Frequency f, kHz

Fig. 5 Power spectra on a brick by

(a) Q-switched laser (b) a long pulse laser

interferometer operated at the velocity decoding mode with a low pass filter of 30 KHz.

The

measured waveform is analyzed in the frequency domain and a vibration modal analysis is applied, as shown in Fig.3. 3-2

Comparison of Q-switched and Long Pulse Laser Figure 4 shows the vibration signal of a 210 X 100 X 30 mm brick.

When a 0.12 J

pulse of a Q-switched YAG laser (532nm, 8ns) was irradiated onto the specimen, the vibration signal was small and not clear, as shown in (a). It is because of the ambient noise and a large ultrasonic attenuation in the brick. When a 1.25 J pulse of the long pulse laser was irradiated, a much clearer signal was observed, as shown in (b).

The power spectra in Fig. 5

show a 20 dB enhancement of signal-to-noise ratio in the 5 to 10 kHz range by using the long pulse laser (b) over the Q-switched laser (a).

This is consistent with the fact that the pulse

energy of the long pulse laser was about 10 times that of the Q-switched laser (see Fig. 1). Figure 6 demonstrates the vibration modal analysis of a stainless steel (SUS304) Tshape joint (40 mm diameter and 75 mm length) using the long pulse laser. The vibration excited by a single shot of a 1.2 J pulse was measured without averaging at the sampling points shown in (a).

The amplitude and phase of a 17.06 kHz spectral peak were used to

produce the vibration mode image shown in (b).

Though the vibration was large (100 /z

m/s), there was no damage of the steel surface, due to the low peak power (24 kW).

In

contrast, when a 0.12 J pulse of the Q-switched laser was used, the steel surface was damaged due to the high peak power (15 MW), though the vibration is not so strong.

Low Frequency Laser Ultrasound 3-3

321

Application to Concrete To test the applicability to concrete, we irradiated a 1.2 J pulse on a 100 X 100 X 5

mm colored tile bonded onto a 300 X 300 X 60 mm concrete plate, and observed clear thermoelastic vibration spectrum as shown in Fig. 7 (a). When the concrete plate size was changed to 250 X 250X 50 mm, the power spectrum was changed to (b), confirming the vibration of the concrete plate.

However, when the colored tile was removed, the magnitude

of vibration was significantly reduced.

Even in this case, when we focus the laser beam, the

ablation produced a clear spectrum, as shown in (c).

It showed identical peaks to (b) below

5 kHz, whereas different peaks above 5 kHz, showing different vibration mode. The resonance frequency of deflection vibration of a free rectangular column can be calculated with a correction factor for the rotational inertia and shear deformation.

We have

proved that for measured and calculated resonance frequencies of concrete column, the agreement is quite good for the first three modes, showing a potential accuracy of the present method, which will be published elsewhere [5].

(a)

Sampling Points

(b) Vibration mode at 17.06 KHz

Fig. 6 Vibration modal analysis of a steel joint 4.

DISCUSSIONS In the focussed Q-switched laser, the extremely high power density (e.g. 500

MW/cm 2 in case B2 of Table I) not only vaporizes but also heat the vapor to a very high temperature and produces plasma.

Since the plasma absorbs the laser energy, relatively

small fraction of laser energy actually reach the object surface [3]. Though the recoil pressure due to the blow off is very high and has a long tail of 10/~ s duration, the peak pressure does not continue so long.

Therefore the efficiency for low frequency vibration

excitation may not be as high as the long pulse laser.

322

K. Yamanaka et al.

The very high peak power of Q-switched laser also requires special care for preventing damage of optics and for safety operation.

It is not easy to construct and operate a Q . . . .

'

I

'

'

I

'

I

Thermoelastic xlO

(a)

40

(b)

L ~

Thermoelastic

L ~

xlO

5

(c) Abl~tion

aS 20 I,...

o 13_

9

0 0

.

5

10

15

Frequency f, kHz (a) 300 •

mm concrete plate with a 100• 100x5 mm tile on the surface

(b) 250 x250•

mm concrete plate with a 100x 100•

(c) 250 x250•

mm concrete plate

mm tile on the surface

Fig. 7 Power spectra of vibration of concrete slabs with and without a tile

switched laser with a pulse energy as large as 2.5 J, for this reason.

In this respect, the long

pulse laser may be advantageous also in the ablation mode of generating low frequency vibration. In the present work we employed the wavelength of 1064 nm. However, this may not be optimum and other wavelength such as 10.6 /2 m of the CO2 laser may be more useful. Further study on physics of interaction between high power laser and object is needed to find optimum condition for excitation of low frequency vibration.

5.

CONCLUSION We developed a long pulse laser, a normal pulse YAG laser using a controlled

discharge current, which achieves both relatively short pulse width as a normal pulse laser, and suppressed relaxation oscillation.

The pulse had a rise time of 15 # s and a width of 50

# s and a large energy of 2.5 J using a single laser head.

The advantage of this laser for

nondestructive evaluation using low frequency vibration was theoretically analyzed and

Low Frequency Laser Ultrasound experimentally verified on bricks, steel joints and concrete plates.

323 The above study proved

that the laser impact method using a long pulse laser can be applied to nondestructive evaluation of large structures. A part of this work was supported by the Proposal-Based New Industry Creative Type Technology R&D Promotion Program from the New Energy and Industrial Technology Development Organization (NEDO) Japan. References 1) M.J. Sansalone and W. B. Street, "Impact-Echo", Bullbrier Press, Ithaca, New York, 1997. 2) J.-P. Monchalin, IEEE Trans. UFFC, 33, 485, 1986. 3) C.B. Scruby and L. E. Drain, "Laser Ultrasonics - Techniques and Applications-" (1990) Adam Hilger, Bristol, Philadelphia and New York. 4) P. Cielo, X. Maldague, G. Rousset and C. K. Jen, Materials Evaluation, 43, 1111, 1985. 5) K. Yamanaka, H. Cho, S. Ishikawa, T. Nakazawa and T. Watanabe, to be submitted.

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MAGNETICS AND OPTICS

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

327

Q U A N T I T A T I V E N O N D E S T R U C T I V E E V A L U A T I O N OF A S U R F A C E CRACK BY A R E M O T E M A G N E T O - O P T I C A L I N S P E C T I O N SYSTEM Jinyi Lee 1, Hiroshi Kato l, Tetsuo Shoji 2 and Dorian Minkov2 1Department of Mechanical Engineering, Saitama University, 255 Shimo-okubo, Urawa, Saitama 338-8570, Japan 2 Fracture Research Institute, Tohoku University

ABSTRACT A new remote sensing system using a magneto-optical method has been developed for inspection of defects which can be applied for on-line inspection where the accessibility is limited. Remote sensing of defects is achieved by a combination of the Faraday effect, optical permeability of domain walls and diffraction of a laser beam by domain walls. This paper describes a novel remote magneto-optical nondestructive inspection system (hereafter refer to as RMO system) and the evaluation method for detecting a crack. This new system makes it possible to carry out remote and high speed inspection of cracks from the intensity of the reflected light and to estimate the shape of a crack more effectively than by already existing methods. Advantages of this nondestructive evaluation method by a RMO system are demonstrated using nine types of slit on plate specimens. KEYWORDS Nondestructive inspection, Magneto-optical method, Faraday effect, Magnetic domain, Remote sensing.

INTRODUCTION The use of a laser as a light source allows remote sensing of a defect by detecting two brighter areas corresponding to a crack and a linear dark line. The method is based on the three characteristics of the magneto-optical sensor (MO sensor) ; Faraday rotation of the magnetic domains, optical permeability of domain walls and diffraction of the laser rays by a domain wall. When linearly polarized light is transmitted through a magneto-optical film(MO film), the degree of rotation of the plane of polarization of the light is characterized by an angle of Faraday rotation. The output optical intensity which corresponds to a particular point of the MO film depends on the intensity of the external magnetic field at this point and the setting of the analyzer. The MO sensor contains a MO film ((GdBi)3(FeA1)5012) grown on a substrate of (GdCa)3(MgZrGa)5012 and a vaporized aluminum film reflection mirror The Faraday rotation is almost doubled by this reflection type optical system. For horizontal magnetization of the specimen surface, two kinds of large magnetic domains occur on both sides of a crack due to the detection of the vertical component of the leakage magnetic flux (LMF) by the MO sensor. This leads to a lower density of the domain walls in the crack area compared with the un-

328

J. L e e et al.

cracked area. As the intensity of the incident or the reflected light is affected when it passes through a domain wall, the optical permeability changes and the image data is brighter in the area of the crack. Since laser rays are diffracted by a domain wall which exist on the center of a crack, dark and bright areas appear in turn at the image data of the RMO system. The width of a magnetic domain wall is usually less than l~tm, but the width of the dark area of the image at the center of a crack is more than 100~tm, due to diffraction of the laser rays [ 1-3]. This paper describes a novel remote magneto-optical nondestructive inspection system (hereafter refer to as RMO system) and the evaluation method for detecting a crack. This new system makes it possible to carry out remote and high speed inspection of cracks from the intensity of the reflected light and to estimate the shape of a crack more effectively than by already existing methods. Advantages of this nondestructive evaluation method by a RMO system are demonstrated using nine types of slit on plate specimens.

RMO SYSTEM The novel RMO system consists of the main system, the separator, and the processor as shown in Fig. 1. The expanded polarized laser beam is transmitted through a half mirror to the separator and reflected by the MO sensor positioned on the specimen. Information on the crack is transcribed onto the MO sensor by the applied magnetic field. This information can be analyzed using the laser ray after it is reflected first by the A1 film at the base of the MO sensor, then by the half mirror, has passed through the fixed analyzer and finally reached the main system. Two laser beams, generated by the half mirror in the main system, are transmitted to a digital camera and an optical sensor after passing through a lense. The output-l, image data from digital camera, is recorded, observed and processed by a TV monitor, a video recorder and a computer. The output-2 from the optical sensor is saved and processed by an optical power meter and a computer.

Fig. 1 RMO system

QNDE of Surface Crack by RMO System

329

1~75Ic=10 .......iii~--:~_~ (unit:ram)

|

(A),

.

.

.

.

.

.

,~=o.7, a,=l

.

.

.

.

.

.

.

(D)

we=0.7, & = l

(B)I :~=0.7,a~2 i (El ~ .

(C)l

.

.

.

.

.

.

.

.

wc=0.7~dc=3 .

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

iT

---~

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

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.

.

.

.

(G) . . . . . . . . . . . ( H),i,,~fo.s, a~=l.s&3 ..................... , I

t=~ wc=0.7~d c = 3

.

Fig. 2

(')

iwc=0.5~ &=l&3

!

Shape of slits

A steel plate of JIS SS400 was machined into plate specimens 100mm wide x 200mm long x 5mm thick. In this paper, we investigated the following nine slits with different cross sections, as shown in Fig. 2: rectangle with different height (A, B, C), isosceles triangle with different height and width (D, E, F, G), stepped bottom (H, I). Slits were introduced by mechanical machining (A, B, C) and electric dischage machining (D, E, F, G, H, I).

EXPERIMENTAL RESULTS An MO sensor has been used in this paper which the width of the magnetic domain, the thickness of the MO film, the magnetic field for saturation and the Faraday rotation parameter are about 20~tm, 151xm, 4kA/m and 22000deg/cm, respectively. Processed image data (output-l) are shown in Fig. 3 which were obtained by the RMO system with normal magnetization (4000 turns and 250mm distance between the poles), i.e. parallel to the surface of a specimen and perpendicular to the slit. A narrow dark segment was observed at the center of the crack and bright areas appeared on each side of it. A length of the dark narrow segment between the two bright areas was equal to a length of the crack when the magnetization current was 0.2A. 'Y' shape dark areas were observed at tips of the crack when the magnetization current was 0.6A, and the center of this 'Y' shape coincides with the tip of the crack. Thus, the crack length can be obtained by using output-1 of the RMO system. The distance between the two 'Y' shapes was constant when the magnetization current was increased with a specimen of Type C. In the case of isosceles triangle, Type G, the two bright areas were appeared at the center of crack when the magnetization current was small (0.1A), and extended to tips of the crack when the magnetization current was increased (0.2A and 0.6A). As shown in Type H and I, two bright areas were appeared on the deep side of the crack in the case of stepped bottom at the small magnetization current (0.1A), and extended to the another tip of the crack with increasing current (0.2A and 0.6A). Therefore it was found the shape of a crack is obtained from the image data of the RMO system. Additionally, as the magnitude of the bright area is proportional to the intensity of the reflected light, the RMO system provides an efficient method to establish a relationship between the size of the bright area and the depth of crack by using output-2. The relationship between the magnetization current and the normalized intensity of output -2 is shown in Fig. 4. Output-2 increases with increasing magnitude of magnetization current. Correspondingly, the intensity of reflected light, output-2, is related to the depth of the crack.

J. Lee et al.

330

Fig. 3

Results of inspection of slit with image data

Fig. 4 Results of inspection by use of nomarlized intensity of reflecting light

ANALYTICAL APPROACH In the dipole model [4-6], the vertical component of LMF, Hz, of a rectangle shape of crack is expressed as:

Iz(x,r,

rc, -,f x

zo+u+ (Zo +

+

/ 2y +

--(1) 4~rkt a-i~/2-y

X -- Wc / 2) 2 + y2 + (Z~ + U)2

where lc, wc, dc are the length, the width, and the depth of the crack, m and/z are magnetic charge per unit area and the permeability of the specimen, for which the latter is assumed to be constant in this paper, while (x, y, z=zo) and (0,0,0) are the position of detection and the center of crack respectively. Complex bottom shapes were simplified by dividing the slit length lc into narrow rectangle shapes as shown in Fig. 5 andHz was approximated as follows : n

H z ( x , y , Zo) = ~-'H~ ( x , y , Zo) i=1

---(2)

QNDE of Surface Crack by RMO System

331

where H~(x,y, zo) means the vertical component of LMF of a [i th] rectangle shape of slit. Fig. 6 show 2-D contour lines which were projected over the Hz=0.3 by use of equation (2) and specimen (C), (G), (H) and (I) for 2Wx=20mm, 2Wy=20mm, , Z0=0.01mm, ml4~rlz=l, 2 and 6. It shows good agreements between the simulated result and image data in Fig. 3. The two bright areas of image data occur when Hz(x,y, zo) > Hs, where Hs is the saturated magnetic field of the MO sensor, as shown in Fig. 7. Therefore, the bright area (Abright) is expressed as:

Abright- f~nz(x,Y,

Zo)l-ns~lxdy

---(3)

where IHz(x,y, zo)l is used to express the density of domain walls.

I

Z

(c) q

~, -y ,-,. X~

7.

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

' 7 ................|

(I)

"1

r

iiii1~

'~

,~,

(I)

,,

"

i=l

(H) IN)

(G)

i=l, I .............

i

2

r

i=l, 2 .'.-'6, 7

.._

r ,,

Fig. 5

Simplified complex bottom shape of the slits

Fig. 6 Simulated results for comparing with image data by use of simplified shape of slits

The integral equation (3) is inconvenient for practical use. Correspondingly, the following approximation:

A bright '~ 4.._~ x2, i - Xl,i

---(4)

9

i=0

F/

was used for easy and fast calculations, where xl,t and x2,i are solutions of a following equation (5) within the range 0 < xt,i < x2,i < Wx :

nz (x, WY . i,Zo )l - Hs : 0

---(5)

332

J. Lee et al.

where n is a number of lines which divide equally Wy while the size of the MO sensor is 2Wx by 2Wy. When x2,i is greater than Wx, x2,i can be put equal to Wx. Calculated results of the bright area with increasing magnitude of m / ( 4 ~ ) are shown in Fig. 8. Calculation details will be discussed in the conference. 500

(A) -C]- (B) -A-- (C) - e - (D) " - ~ (E) --Akr--(F) --.x-- (G) ~ - (H) -•

04

............ ......."

/

_ (

\;

:"

,j

/

E E

.....~....- --- 't

t~

~ 250

<

-g , mL

0

0

Fig.7 Theoritical consideration of the two bright areas by use of the vertical componet of the LMF

2

4

6

m/(4mz) Fig. 8 Calculated bright area with increasing magnitude of m/(4~t)

CONCLUSIONS A novel remote magneto-optical of this system is that image simultaneously. The shape of a crack can be calculated from the

nondestructive inspection system has been developed. The feature data and leakage magnetic flux information can be gained crack can be obtained from the image data and the depth of the intensity of the reflected light.

REFERENCES

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

Lee, J. and Shoji, T. (1999) JSNDI 4, 231. Lee, J., Shoji, T., Minkov D. A., Ishihara, M. (1998) JSME 619, 825. Kudo, K. and Uehara, T.(1995). The basic Optics, Kendai industrials publisher, Japan Mukae, S., Katob, M., Nishio, K. (1988) JSND111, 885 Minkov, D. A., Lee, J. and Shoji, T. (2000) Mat. Eval. 5, 661 Minkov, D. A., Lee, J. and Shoji, T., (1999) JMMM No.659

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

333

M A G N E T I C A N D O P T I C A L N O N D E S T R U C T I V E EVALUATIONS. F O R IRON-BASED MATERIALS

K. YAMADA, K. YAMAGUCHIt, S. TOYOOKA and Y. ISOBE*

Graduate School of Saitama University, Urawa, Saitama 338-8570, Japan *Nuclear Fuel Ind. Ltd., Sennangun, Osaka 590-0481, Japan

ABSTRACT The multiple diagnostic tools have been developed for iron-based structural materials of A533B (97%-Fe, ferromagnetic) and SUS304 (18%Cr, 8%Ni, paramagnetic) in an applied stress range up to 700 MPa or in a residual strain range up to 20 %. These methods were composed of (1) an Ar-laser speckle pattern interferometry for surface deformation measurements down to sub-micrometer, (2) a leakage flux sensor of a GaAs Hall element for observations of inhomogeneous magnetic leakage flux distributions after an uniform polarization, (3) a pick-up coil mounted in-between magnetic two yokes, for residual strain estimation by detecting magnetic Barkhausen noises generated during magnetic field sweep up to the saturation and (4) magneto-resistance observations in pulsed fields enhanced by the Martensitic Transformation. Very sensitive detections of residual strains were attained as small as 0 . 1 % by magnetic Barkhausen noise observation and 0.4% by leakage flux sensing, respectively for A533B. We found magneto-resistance effect of 0.03% for heavily strained SUS304 samples, in comparison with that of 0.4% in pure Fe. KEYWORDS NDE, Hall sensors, leakage flux, A533B, magneto-resistance, SUS304 INTRODUCTION The life managements of power plants of any kinds are of great importance nowadays for the securities of our modem societies and for material recurrences without fatal accidents. In this study, we aimed at the early detection of the material degradations far before the destructions of power plants. For these purposes, the optical diagnosis of the surface deformations and the magnetic diagnosis of the material degradations for iron-based material were investigated. Detections of surface deformations were performed by using Ar-laser speckle pattern interferometry (ALSPI) for structural material as accurately as 260 nm, even for rough surfaces as those are in the factory. In principle, the magnetic properties of iron-based materials are very sensitive to the lattice imperfections so that the degradations should show the close relationship with the macroscopic magnetic susceptibility, remanence and coersive force. Therefore, the magnetic diagnoses are composed of the detection of Magnetic Barkhausen Noise (MBN) during magnetic field sweep to the saturation magnetization, the magnetic leakage flux (LF) observations after uniform polarization of the sample of A533B (Fe97,Mn,Ni) and Magneto-Resistance (MR) for SUS304 (Fe7a,Crls.Nis). EXPERIMENTALS Very precise sample surface deformations in 3-dimentional directions were observed by Ar-laser Speckle Pattern Interferometry [1,2,3] with an accuracy of about 260 nm, which is valid for t Present address: Dept. Education, Fukushima University, Fukushima 960-1296, Japan

334

K. Yamada et al.

perpendicularly incidental light angles and the laser light wavelength (At-laser, L=515nm). ALSPI ensured the residual strains through the surface deformations together with the lattice constant changes observed by X-ray diffractions for the direct evidences. Fig.1 shows (a) the typical patterns obtained by CCD camera by ALSPI for thickness deformation in elongating direction of the sample of A533B, together with (b) the pattern of magnetic Leakage Flux (LF) distribution for comparison., which will be soon explained just below.

Fig. 1 The comparison between ALSPI and LF distribution patterns for the same sample. (a) ALSPI difference pattern of the surface deformation in the elongating direction. (b) Magnetic LF distribution pattern observed after a polarization of the same sample. Note here that one fringe distance corresponds to 260 nm in Fig. 1 (a) at the sample surface in the elongating direction and the Ltiders band locates at the skew shoulder "L" in (b) where the incoming and outgoing fluxes are drawn in B&W, respectively. Further, the gray colored band "G" [4] in Fig.l(a) at the end of the Liiders band represents the cumulated multiple fringes much larger than the distance of 260 nm. The experiments of magnetic LF observation were performed by using a small semiconductor Hall element of GaAs (1201xmxl201xm) with a lift-off distance of 250Inn after a uniform polarization of samples in 8000A/m (lkOe) as already shown in Fig. l(b) for example. Fig. 2 shows the leakage flux distribution in samples of A533B as a function of thickness deformations (a) with residual strains and (b) without residual strain, respectively. Where, the sample without residual strain was prepared by annealing in 1100~ 2hrs after locally mechanical thinning of 8, 60 and 50 %, respectively. The LF distribution showed little flux gradient fluctuation for a sample of a thickness deformation of 8% without strain. Whereas, a sample of a thickness deformation of 0.8% caused by a tensile stress of 480 MPa with residual strain inside showed much larger flux fluctuation than those for the 8% sample as easily seen by their fwst derivatives (c3Oz/o~x) in Fig.2. Here the coordinate system is chosen as x for the sample elongation direction depicted in the figure and z for the perpendicular direction to the surface.

Fig.2 The magnetic LF distribution for A533B samples (a) with and (b) without strains, respectively Note: The thickness deformations without strains were intentionally made for comparisons.

335

Magnetic Nondestructive Evaluation

We performed the experiments of MBN observations [5] for A533B by using a pick-up coil located in between the magnetic Po41es during field increases up to 2400 A/m (30Oe) in a constant sweeping speed of 3.6x10 A/ms. The MBN storage was performed after several hysteresys loop circulations to eliminate the minor loop characteristics before the data acquisition. MBN were observed in a very slow to obtain separate spikes of dM/dt signals. In this sweep rate, the MBN trace gave the largest fractal 2-D index of 1.6+0.5 for a strained sample [6,7].

(

[ No.t P,,,~

k J !

i..............

h

|,

II

-

--

e

|,

d c

8 e d c

b

b

l

i 0

It

0

2 H

II~Ao'ml

IH

lkAttnl

2

Fig.3 MBN spectra as a function of magnetic fields for different positions along A533B samples of No. 1 (640Mpa) and No4 (480MPa), respectively Notel: the positions "a"-"k" are arranged from the top to the end of the sample, and "g" and "f" locate almost the center of the sample and "a" and "k", the ends. Note2: the applied field direction (N-S) was parallel to x / / P (applied tensor stress). Fig. 3 shows the MBN as a function of magnetic fields up to 2400A/m for A533B at different positions of"a"-"k" along x for differently strained No. 1 and No.4 samples. As evidently shown in these figures, the MBN noises disappear for a heavily strained No. 1-sample at the position "g" and "f". On the other hands, slightly strained sample of No.4-sample, the differences of MBN spectra among positions are rather small. MBN traces were analyzed by an indicator of Average Field for Discontinuous Magnetizations (), which will be discussed later. The transversal and longitudinal MR's were observed for SUS304 with a very stable sample current (I) of 100mA/mm "~in pulsed fields (H) of 1 s rise up to 3 kOe and 2 s down to zero. Both of positive longitudinal and negative transversal MR's showed the same order of 0.03%, respectively for a sample with 8 % residual strain as shown in Fig. 4. It increased in a tensile stress range up to P=680MPa (e=8%). However, MR was found very small effects even in high fields of 8000A/m for the samples of small residual strain less than 1% [7].

0.4

" '

0.3

-

~ 02

~

'

0.04

' .............

A ,IAA

AIAA A

'

U,03

o TI'ImL

I!

oo

|

,

9

,.

9 L. (SS0.~,)

A

o r,

(sso~,,)i

T.. (e~a)

0,01

oo

,,

,

AAdk

0.02

ALous.

0.1

....

SUSSO4[ R.T.

T. (OMPs)

II

t

-0.2

(a)

0

le

4.01

-ql.1 ,

I

200

,

I

400

,

Magnetic held l-qOe]

I

-w

600 Co)

el ,

I

200

,

'

400

Ma~edlc lieu l[[Oe]

Fig. 4 Magnetoresistance as a function of magnetic field (a) a pure Fe(~=0%) sample, (b) SUS304 samples (e=8%), respectively.

I 6oo

K. Yamada et al.

336

The magnitude of MR in SUS304 was 10 times smaller than that of pure Fe, which was performed in an independent experiment. It must be noted here that the MR for as-annealed sample was smaller than detectable level. Further, the demagnetizing field must be considered for transverse MR due to elongating sample shape, which will be discussed later. The longitudinal MR was found positive and transverse MR, negative both for Fe and SUS304. Theses positive and negative MR must be corrected for the effective magnetic field, due to the different demagnetizing field for the longitudinal and transversal directions of the sample. DISCUSSIONS

We examine in this section on the availability of LF, MBN spectra and MR measurements as NDE technology. At first, we observed magnetic LF distribution for A533B samples. In principle, the distribution of LF from inhomogeneously polarized sample with thickness deformation are expressed by the Gauss's law as

(R-r)

dlvB(R) = ~gradp[r(r)]. ~ d V

+ H~p[r(r)]. grad[ (R. -r), 3]dV , 4z}R-r[

gradp[r(r)] = ~ gradr(r)

(1)

(2)

aT

where, B(R), p(r(r)) stand for the magnetic flux density and the local remanence at R and r with the lattice defect density r(r), respectively. The second part of (1) stands for the contribution of the leakage flux by the thickness deformations. The local remanence p(r(r)) decrease monotonically with r(r) as dp/dr <0 in (2) [8]. We show the dependence of magnetization on the lattice defects in the last part of the discussion by simulation. For A533B samples, the LF gradient fluctuations with distance is caused dominantly by the local remanence change rather than the geometrical change as thickness deformations. Secondly, in the MBN experiments, the amplitudes are not stable enough for NDE by different touching manners of pick-up coil to sample surfaces at each position because of various curvatures of structural shapes and because of various surface roughness with stains Therefore, we propose an indicator of Average Field for Discontinuous Magnetization (AFDM) as denoted derived by the observed quantity MBN voltage V as H Max

BMax

H(t)V(R,O,h,t)dt = ,Ou,= . . . .

~HdBm o BMax

I V(R,O,h,t)dt 0

(3)

f dBm 0

Here, H(O and V(R, O,,h,O denote applied magnetic field and observed BHN voltage in the pick-up coil at a position R, angle 0 and a lift-off distance h. It is important to note here that the magnetic flux density Bent denotes the internal flux for the material, instead of V(R, 0, h) obtained by leakage flux change outside, which is a plausible assumption because of the observed leakage flux are connected by dB~./dt=CL dBout t/dt = -V(~ O,,h,t) both for the nominator and the denominator in (3) [6,7]. By this reason, the quantity of becomes rather insensitive against MBN amplitudes due to the common amplitude CL for the nominator and denominator in (3). Equation (3) is now easily replaced with the physical quantities as [9,10,11,12] g

1 J ~(x~j -x,j). 6//,

=H ~ --~ -_

xDJ

(4)

Magnetic Nondestructive Evaluation

J j=Ok=O

337

XDj

Here Axe. denotes the k'th magnetic domain shift at time t=t~.. In this interpretation of the indicator, <1to> represents the average field for normalized domain jump distances to saturation. Fig. 5(a) shows and (b) residual stresses in x and y directions as a function of residual strains. Experimentally, we obtain the linear relationship between and a For A533B material as

(6)

[A/m] = 91x104o [kg/mm 2] ,

Note here that this linear relationship is valid in a strain range less than 2%. Thus, we showed a possibility of the visual indication of degradation by leakage flux gradient available both for A533B caused by the smaller remanence and for SUS material by martensitic transformation. Further available indicator for material of degradation by is useful for industrial use.

2;8

,..:"",,.____./ )

27

It

"

26

"

" ' ~ 1 .~.

,w ~'"

9

~"I"

r 2 5 It;

9

0

ca)

'

2

I

,I

4

!

~

'

6

'

8

,

R~dual s'train s [%1

I

10

0

(b)

2

4

R~dud ~

6

8

10

s [%1

Fig. 5 Comparison between and the residual strain (~x, try), respectively We examine here MR in SUS304 material to obtain the relationship between the amplitude of MR and the volume ratio of austcnite to martensite as a function of residual strain or stress. We adopt a physical model [12] that the austenite material has a resistivity constantpa (7, t ) and that martensite, PM(Z, E ,M(H)) at a ambient temperature T, residual strain E and sample magnetization M(H). For the austenite and martensite phases, the specific resistivity must be different so that PM(Z, E )=/=PM(Z, E ,0). As a first approximation, resistivity changes with magnetization in martensitic parts are short-circuited by the remaining austenite. Therefore, the total resistivity R for the sample is expressed as

Ra = cpA (T,6)

(7)

1-v P~ =

cpu(T,6,M) V

,

(p~ = m 0 +ap~)

(8)

Here, v denotes the volume ratio of the transformation.(v---0 means whole sample is of austenite). PMOand @M denote the resistivity of mart~nsite in zero magnetization and the deviation from PMO in H>0 with a condition for the real case of 5PM/PM0.<< 1. The total resistance of the sample composed of RAand RM is approximated by their parallel connection as

338

K. Yamada et al. cpA(PMo +8 PM )

R=

(9)

(PA "- PMO - ~ PAl )V + PMO + ~ PM

The magnetoresistance change in a field H is now expressed by the resistance R/R(H=O) as b~(H) ~ v R(H = O)

~

8 Pu(,) PMO

~

(10)

Here, applied magnetic fields in an anisotropic sample shape must be corrected by the demagnetizing field. For this purpose, anisotropic magnetizations of the sample were measured for as-annealed and a strained sample in longitudinal and transversal directions against sample elongating direction. Fig. 6(a) shows the magnetization of an as-annealed sample as a function of applied magnetic field in parallel and perpendicular directions (M/n~o~Mj_~), respectively and Fig.6(b), the magnetizations for a strained SUS304 sample. As are easily noticed, M/n~on~is larger than M• only due to the larger demagnetizing fields caused by the anisotropic sample shape. Fig.6(c) shows MR(H*) as a function of H* derived by the effective fields H*=0.33H, by which corrections, MILo~g(H*)well coincided with M#t~(H*) (closed circles in Fig. 6(a, b)).

"'~ . . . .

I~ ,.,

. . .

am

~ oo ~ 7

.~m aooo ,~[

. . . . . . . . . . .

~

noo

mn

J,~, :m~sensJH[Oel

~,o

~

.iooo--inn

~

~

... IJliJ '"

.......

mu

~ooo ]m

//~ mqMucm~ul~]

i

,,v,,

i

'.........

I//M/ll.~nl; A ,I~

A

~" U!

ad

o ~j

[IM//Lons . | iner cozrecuon m

. . . .

q

.

!

.

-

Malln~lr I~ld lt[Oe]

(,)

(b)

-106

(r

Fig. 6 Magnetization of as-annealed and strained sample of SUS304(H), and MR in strained SUS304 sample. We observed the surface deformations of SUS304 samples by an electron microscope to observe skew slip bands, which might cause the magnetic anisotropy. Fig. 7 shows the surface state of an as-annealed sample and a heavily strained sample. The surface states are isotropic observed for an as-annealed sample (a, b) in comparison with those for a strained sample (c, d) with 640MPa, where slight anisotropies are found.

Fig. 7 Surface observations of as-annealed and heavily strained SUS304 samples by an electron microscope [ 13].

Magnetic Nondestructive Evaluation

339

Finally we show the decreasing manner of magnetization with dislocation density increase by using the Monte-Carlo method [14]. The spin system of 303=27000 cells with and without a half plane spin voids locating at the cube center to its fight side in this simulation. The change of whole spins from the initial state was calculated by Metropolis method using Heisenberg Hamiltonian H as below. H =-.~.JijSiSj -B~.,S i

(11)

,

i

t,J

where Jo denotes the exchange energy acting on the nearest neighbor sites between i and j, Si and Sj. represent spin vectors at these sites respectively, and B, the applied magnetic field. Because of the long computation time, i,j in Eq. (11) run only over the first nearest neighbor spins with a constant exchange energy J in the former simulations [14]. Now, we consider not only the nearest neighbors but also the fimher sites with different Jo, because metals under consideration such as iron contains itinerant electrons. Here we adopt the RKKY model for J0 [ 15,16,17] as

Isin 2Krr,j - 2Kpr~jcos 2Krr~j)

(,,,;r,,)'

. . . . . . .

.

.

.

.

.

.

.

.

.

ox (-

(12)

Here Ke denotes Fermi wave number and L, the mean free path. These parameters were set Kr=1 and k=3 in the lattice constant unit. Here, KF was tentatively set at 6F=10eV with the free electron mass me. 10

I x

tO

O.B

x

E

'li

'

•

~

~

'

x w ~ o u t Oislocation 9 with dislocxllon

•

N

"o

I

!

x

.

0.4

02-

~

o'.5

!

!

~o

~-~

!

2n

i

Im

2.~

m

aO

Tp..rnr~.ralzjrn tart} zjnlft

Fig. 8 The simulation of temperature dependence of magnetization with dislocations Fig. 8 shows simulated results of the temperature dependence of the magnetization in B=O in Eq.(ll). The simulations were performed with an initial condition of the random spin arrangements corresponding to the paramagnetic state above in the Curie temperature. A series of magnetization processes was simulated with the same initial condition in low temperatures. As shown in Fig.8, the magnetizations without any dislocations well obey the Curie-Weiss law. Calculated magnetization with the planar edge dislocations was smaller than those without dislocations in the lower temperature range. The magnetization reductions were well simulated by magnetic domain formations around the edge dislocations. The difficulties in the Monte Carlo simulation for metal magnetism must be discussed here. One of them is, for example, the anti-commutative law of Fermions. In this paper, we neglected the quantum statistical behavior of the spin dusters because the spin cluster number n together with the total spin angular momenta n h/2 are fluctuated by the thermal energy in high temperatures around RT. The lattice type of the spin sites must also be considered precisely, because the magnetism is completely different for different lattice formations. However, the simulations coincide well with real magnetic behaviors of metals at least in RT, or these simulations might be valid only for the system of the classical magnetic moments. The development of these simulations for metal magnetism might provide the analytical methods for NDE in the future.

340

K. Yamada et al.

CONCLUSIONS In this study we successfully evaluated the residual strains of structural materials as a function of applied stresses with a high resolutions down to 0.26 nm by ALSPI, with a high sensitivity down to 0 . 1 % strain by MBN for A533B sample. By using magnetic LF sensor, any spatial inhomogeneity of residual strains, or more strictly speaking, distributions of local remanences can be detected for ferromagnetic material of A533B. The spatially detectable range at the sample surface, however, was found very short as 50 ~m of order. For such purposes of the internal inspection, MR is a suitable method because of the deep flow of the current. REFERENCES

1. J.W. Goodman, Statistical Properties of Laser Speckle Patterns, Springer Verlag, Berlin 1993 2. Suprapedi and S. Toyooka,~Optical Review, 4, 1997,284-287 3. K. Yamada et al. Proc. 32~ ISATA ( Materials for Energy-Efficient Vehicles) pp 573-580, June 1999,Vienna, Austria 4. T. Kolmo et.al. Proc. 29th Strnp. On Stress-strain Measurement and Strength Evaluation, (1998)pp915. D.C. Jiles, Review of magnetic methods for NDE, NDT Int. 21, 5(1988)pp311-319 6. K. Yamada et.al. Nonlinear Electromagnetic System, ed. by Kose and J. Sievert, IOS Press, 1998, pp153-156 7. K. Yamada and T. Saitoh, J. Mag. Magn. Mater. 104,1991,pp341-342 8. S. Ishige, Master thesis, Saitama Univ., March 2000 9. H. Kronmuler, Candian J. Phys. Vo145(1967) 10. S. Shoji, Doctor Thesis, Saitama Univ., March 1999 11. K. Yamada et al, Proc. IWAM '97, Nagasaki, 1997,pp114-119 12. K. Yarnada, Y. Isobe et al., J. Magn. Soc. Jpn., 23, 1999,pp718-720 13. K. Yamada et al., in press, Proc. Workshop on Magnetism and Lattice Imperfections, Hanamaki, Iwate, April, 2000 14. K. Yamaguchi et.al, in press, Proc. Workshop on Magnetism and Lattice Imperfections 15. T. Kasuya, Prog. Theo. Phys. 16(1956) 45 16. K. Yoshida, Phys. Rev. 106(1957)893 17. K. Yamaguchiet al, to be published by Proc. EMMA2000, May, Kiev 2000

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

341

EVALUATION OF POST WELD HEAT TREATMENT TEMPERATURE OF LOW ALLOY STEEL WELD JOINTS USING AC MAGNETIC METHOD

M. SHIWA. Y. HORII, R. KUME, H. YONEYAMA and A. YAMAGUCHE

Tsurumi R&D Center, Japan Power Engineering & Inspection Corporation, 14-1 Benten-cho, Tsurumi-ku, Yokohama, Kanagawa, Japan, 230-0044

ABSTRACT 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-l.0Mo) under several PWHT temperature conditions were investigated. It is found that 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. 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. Also influence of the surface condition on the testing result was investigated. Two types of surface treatment, a chemical polishing (CP) and mechanical polishing (MP), were used in the test. Measurement after the CP surface processing gives more reliable results. KEYWORDS AC magnetic method, PWHT temperature, Normalized 3rd harmonic component intensity

INTRODUCTION Post weld heat treatment (PWHT) is implemented mainly to reduce the residual stress by weld and to improve the mechanical and metallurgical properties of weld heat affected zone (HAZ) and weld metal. The properties of weldments after PWHT are governed by the PWHT temperature strongly. Therefore it is important to know actual applied PWHT temperature after heat treatment. In this paper, an alternating current (AC) magnetic testing method is proposed to evaluate the PWHT temperature of weld metal of low alloy steel, nondestructively. As magnetic properties are sensitive to microstructure [1], stress [2] and heating temperature [3] of ferromagnetic material, magnetic property measurements have recently been investigated as NDE method to detect creep damage [4]. The results of this paper show that the PWHT temperature can be evaluated by analyzing AC magnetic testing signals.

EXPERJMENTAL PROCEDUER

P WHT specimens Two kinds of low alloy steels (base metals), 1.25Cr-0.5Mo and 2.25Cr-l.0Mo, were prepared for the tests. Chemical compositions and mechanical properties of the specimens are given in Table 1 and Table 2, respectively. The plate specimens used in this investigation were obtained from MIG

342

M. Shiwa et al.

Table 1. Chemical compositions of Cr-Mo steels (mass%) Specimen C Si Mn P 1.25Cr-0.5Mo 2.25Cr-l.0Mo

0.14 0.07

0.64 0.26

0.56 0.47

0.010 0.09

Table 2. Mechanical properties of Cr-Mo steels (RT) Yielding stress Tensile strength Specimen (0.2%/MPa) (MPa~ 1.25 Cr-0.5Mo 354 487 2.25Cr-l.0Mo 316 441

S

Cr

Mo

0.001 0.04

2.32 2.21

0.49 0.98

Elongation (%) 29 39

welded joints of low alloy steel of 20 mm in thickness, 320 mm in length and 130 mm in width. The PWHT was carried out using an electric furnace heating. The specimens were first heated up at a rate of 5 K/min, and then the specimens were continuously heated for one hour at the holding temperature within the range of 853K and 1063K. Before magnetic testing, two types of surface treatment, a chemical polishing (CP) processing (acid solution wiping), and mechanical polishing (MP), were used to remove the surface scale of specimens. A C magnetic measurement system

Figure 1 shows the block diagram of AC magnetic measurement system. The exciting and detecting probe consists of two coaxial coils with differential connection. A ferrite core of 2.5 mm diameter and 20 mm length was inserted in the coils. The specimens were magnetized by supplying sinusoidal voltage to the probe. Resonant frequency of the probe was about 180 kHz. The exciting and the detected voltages were recorded using a 2-channel waveform recorder. After testing, the frequency spectrum of the recorded signals was analyzed.

NF: P64

Probe coil

JTT: DCM120

Detected voltage

2ch input 5MHz sampling 10bit resolution

xl

Isolation amplifier NF: 1915

Ferril oltage

-~2MHz 20Vp-poutput

Waveform recorder

Function generator Specimen Fig. 1. Block diagram of AC magnetic measurement system.

Evaluation of PWHT Using AC Magnetic Method RESULTS Basic properties of P WHT specimens Relationship between Fracture Appeared Transition Temperature (FATT) and PWHT temperature of weld metal (WM), HAZ and base metal (BM) of the 1.25Cr-0.5Mo steel specimens are shown in Fig. 2. The FATT of WM rapidly recovered at PWHT temperature of 953K.

340

343

+WM o- HAZ --m--BM 1.25Cr-0.SMo Iv.SS-------__.._

320 [.., 300 .<

o. S-S ..

--..

S-ss--~ ~ . . . .

2so

",

l

"...

....Q

Figure 3 shows microstructures of the weld 260 As-welded o metals of the 1.25Cr-0.5Mo steel specimens, 200 400 600 800 1000 as-welded (a), PWHT temperature at 913K PWHT temperature / K (b) and 1013K (c). It can be seen that the grain sizes of as-welded and PWHT Fig. 2. Relationship between Fracture Appeared specimens are nearly the same. On the other Transition Temperature (FATT) and PWHT hand, volume of carbide precipitation of the temperature of weld metal (WM), HAZ and PWHT specimens is higher than that of as- base metal (BM) of the 1.25Cr-0.5Mo steel welded specimen. The size of carbide specimens. precipitation in the sub grain of the PWHT specimens increased with the PWHT temperature. oo

Fig. 3. Microstructure of the weld metals of 1.25Cr-0.5Mo steel specimens: (a) for as-welded; (b) for PWHT at 913K; (c) for PWHT at 1013K.

Test signals obtained from P WHT specimens Figure 4 shows the relationship between the detected peak voltage and the PWHT temperature of 2.25Cr-l.0Mo weld metal to which mechanical polishing (MP) and chemical polishing (CP) surfaces treatment was applied, respectively. It is shown that the detected peak voltages change with the PWHT temperature with similar tendency. However the value of detected peak voltage of them was not stable during the measurement.

+ Detected Voltage(M" P) .-o-- Detected Voltage(CP)

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

6.5

'

. . . . .

" '2." 25Cr-1. oMo'! O

> 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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o 4.0 _J

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0.1

I

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0.2

0.3

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

O

0

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0

0 03 0 r

o IE

~

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0

......

3.5

I

I

I

I

4.0

4.5

5.0

5.5

Half Crack Length

6.0 x 10-3

m

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

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

10.0

8.0

Z v'

20.0 x 10-3

15.0

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rn

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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 <100> 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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o

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.m

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0

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0MPa 20MPa 40MPa 60MPa 80MPa 100MPa 150MPa 200MPa

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i

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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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co

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Plastic strain , %

(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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.

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Fig. 8. Master curves after full annealing.

EVALUATION OF RESIDUAL STRESS

Experimental Procedure The results of previous section intimated that the residual compressive stress in the elongation direction exists at the surface of the specimen. To verify the prediction, we evaluated the residual stresses in the elongated specimens by using some conventional methods. We machined specimens from annealed plate. To obtain the master curves for compressive stress, the compressive specimen, whose dimensions are 39mm x 39mm x 1 lmm, was also prepared. The

Table 2. Chemical compositions and mechanical properties. Chemical compositions (mass%)

Yield

Tensile

stress

strength

C

Si

Mn

P

S

(MPa)

(MPa)

0.16

0.35

1.41

0.015

0.002

380

529

362

T. Yamasaki and M. Hirao

chemical compositions and mechanical properties are listed in Table 2. Before and after elongation, master curves were acquired. Then we sliced the specimens to release the residual stresses. The released strains were measured and were converted to the residual stresses. X-ray method and magnetic anisotropy method were also executed. Figure 9 shows the master curves. Besides the maximum magnetostriction, the minimum values of transverse magnetostriction are plotted. When magnetic field is perpendicularto the stress, the minimum value will not appear for large compressive stress. This, together with Fig. 5, also denotes the existence of residual compressive stress. Figure 10 represents the master curves of specimen HR-4 after 3% elongation. Both pairs of master curves traverse each other at 150MPa tension. The residual stress is evaluated to be

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(b) Minimum magnetostriction. Fig. 9. Master curves.

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(a) Maximum magnetostriction. (b) Minimum magnetostriction. Fig. 10. Master curves of specimen HR-4 after 3% elongation.

Magnetostriction of Plastically Deformed Steel

363

150MPa compression. For the other specimens, the residual stress evaluation was done in the same manner. The results of residual stress measurements are listed in Table 3. Since X-ray method indicated transversely isotropic compressive stress of about 200MPa before and after the deformation, the changes in stress indication are listed. The stress relief method revealed that the maximum magnetostriction method could evaluate the residual stress in the elongation direction, while it showed that the compressive stress of comparable magnitude also existed. The magnetic anisotropy supported the results of the magnetostriction method, assuming the residual stress was uniaxial.

Table 3. Results of residual stress evaluation. Specimen

HR-3

HR-4

HR-5

HR-6

HR-7

Elongation

1.6%

3.0%

1.5 %

2.9%

1.5%

Stress

0

1

Magnetostriction

- 115

X-ray

16

0 2

0 1

0 2

-150

0 1

O Z

-110

Magnetic

-90

O Z

(~ 1

-143

-115

-140

-110

O 2

21

anisotropy Stress relief

O 1

-60

-151

-109

-158

-120

-91

-85

-98

-106

CONCLUSION We investigated the effect of plastic deformation on the master curves of maximum magnetostriction method. The obtained results are as follows. The master curves were greatly shifted by the plastic deformation, which can be explained assuming that the residual compressive stress was introduced at the surface of the specimen. Compared with the results of the other methods, magnetic anisotropy method supported the above prediction. REFERENCES 1. 2.

Jiles, D.C. (1988) NDT Intl. 21-5, 311. Yamasaki, T. and Hirao, M. (1998)Nondestr. Charact. Mater. VIII, 665.

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

365

VISUAL CHARACTERIZATION O F W E A R IN LARGE CALIBER WEAPONS D o m i n i c k Salafia Metrology and Simulation Division Material Test Center U. S. Army Yuma Proving Ground Yuma, Arizona 85365

Abstract: As part of the Army Test and Evaluation Command (ATEC) the Metrology and Simulation Division at the U.S. Army Yuma Proving Ground (USAYPG) has the mission to measure and record the detrimental effects of firing conventional and experimental munitions on large caliber cannon tubes. The primary objective is to ensure that the weapon to be fired will safely meet mission requirements for the quantity and energy of the munitions under live fire testing. One aspect of this mission is to conduct physical measurements on rifled and smooth bore cannon tubes. The measured value is compared to the acceptable tolerance; from this the disposition of the weapon is then determined. In the past physical measurements were taken with a "star gage". The star gage is used to measure wear on rifled cannon tubes. This device measures the wear of the reference "land and groove" at the zero and 90-degree positions respectively. Although this method offers a high degree of precision, faults that exist at other positions are not recorded. As a result of this limitation, wear phenomena such as erosion or build up could not be quantified. Recent developments in barrel measurement instrumentation have expanded the analytical capabilities of the status quo. Instead of two data points, 2000 data points may be measured from 0 to 360 degrees. From this data, tooth profiles for rifled barrels and wear in smooth bore cannon tubes may be displayed and threedimensional models may be developed. Different algorithms have been developed to display collected data from a variety of perspectives. The resulting types of perspectives and the visual characterization of different types of wear will be examined in the body of this paper.

INTRODUCTION In order to maximize safety, and to accurately assess the performance o f conventional and experimental muntions, regulation dictates that physical measurements be performed to gage the wear induced. Inspections normally occur at specific distances with respect to the "origin of rifling" for the weapon being tested. The origin of rifling (OR) is usually indicated as d=0. Subsequent measurements down the bore are positive. Figure l illustrates the results of such a measurement. A three-dimensional model may be constructed by adjusting for the twist o f the rifling and piecing the data together linearly using a Delaunay triangulation method. Figure 2 illustrates this methodology. Although the raw data is useful for generation o f computer models, the amount o f data is cumbersome to process and not easily interpreted by the untrained eye. By reducing the data to 48 land and groove radii, the data processing is faster and produces clearer results. The data is reduced by averaging the radial values with respect to lands and grooves. Using this method, the wear effects between firing missions may be quantifiably visualized. Groove radial analysis, land radial analysis, land/groove height radial analysis, percent wear analysis and a comparative analysis showing the cumulative effect o f firing an additional 496 rounds through a 155mm cannon tube will be illustrated and discussed. The standard statistical analysis may be applied to the data from each measurement position. This discussion will focus

366

D. Salafia

on three dimensional analysis techniques and how this method may be used to gage the increase in wear from mission to mission.

FIGURE 2

As indicated in Figure 2., the protuberance on the first groove represents a loss of material. Figure 3 is a density (top view) plot of Figure 2. From this perspective, the chrome loss shows as a darker line. Loss of land material (tooth wear) is indicated by a lighter shade (or color).

FIGURE 3

Visual Characterization of Wear in Weapons

367

RADIAL ANALYSIS

Three Dimensional Groove Radial Analysis The reduced data may be represented as a surface plot (FIG 4). The top view is shown in fig 5

i~~,

4,

FIGURE 5

Since most of the wear is on the rifling (lands), the groove radius is consistent with the manufacturing tolerance of the weapon. This is a useful technique for acceptance of new cannon tubes.

Three Dimensional Land Radial Analysis Figures 5 and 6 illustrate wear on the lands. Unless the person viewing the data is familiar with measurement tolerances, the degree of wear measured is not clearly evident.

FIGURE 6

FIGURE 7

D. Salafia

368

Three Dimensional Land/Groove Height Analysis This analysis is the most indicative of wear with respect to the weapon under test. The relation of land to groove heights is plotted as a three dimensional surface with the corresponding top view Figures 8 and 9. Note that the height is approximately zero at the 12 0'clock position indicating a significant tooth loss.

P E R C E N T W E A R ANALYSIS In order to compare and to quantify the degree of wear induced from mission to mission on a particular cannon tube, the land to grove height is normalized with respect to the nominal land to groove height. This value is determined by the model, caliber and manufacturing process. The value used in this discussion was obtained from the inspection history of the cannon tube. Normalization maps the data into a 0 to 100 percent scale. Figures 10 and 11 show the wear pattern at 2060 rounds and 2556 rounds respectively.

FIGURE 10

FIGURE 11

Visual Characterization of Wear in Weapons

369

The above illustrations show the effect of firing 496 rounds. Notice that the wear is progressing up the cannon tube. Since most of the wear occurs at 0 degrees (12' o' clock) the wear pattern may be rotated 180 degrees and plotted with respect to o'clock position (Figures 12 and 13).

FIGURE 12

F I G U R E 13

The wear incurred by firing 496 rounds is the difference of Figures 7 and 8. is portrayed in figure 14.

FIGURE 14

370

D. Salafia

DISCUSSION As illustrated above the analyst may now easily quantify wear pattems as a function of the number and/or type of munitions fired. Although the entire gun tube may be mapped in by the same methodology, this the author feels that the methodology is best exemplified by concentrating in the area where most of the wear will occur. The above analysis was performed with a commercially available software package. The analyst may use all of the potential of the software package to expand the scope of the analysis or to easily automate the analysis and explore new methodologies. Since the illustrations in this paper are confined to grayscale, some of the impact that color has to offer is somewhat diminished. Albeit, even with this limitation, the reader may notice distinct wear patterns. The comparative analysis in Figure 14 readily illustrates and quantifies the location and severities of wear incurred by the additional firing of 496 rounds. By starting with new cannon tubes as a baseline, cross-referencing the type and quantity of munitions tested, and incorporating visual inspection (borescope) one can conduct a long-term controlled experiment to generate a computer wear model. CONCLUSION The advent of measurement technology in the arena of munitions testing has provided a means to visually characterize and gage the effect of testing on the weapon. An historical database is planned such that a computer model may one day simulate the effect of virtual firing missions. The above methodology, as well as the standard two-dimensional analyses may be tied into the Wear Analysis Database currently in use at the U.S. Army Yuma Proving Ground. In closing, it is the opinion of the author that this technology has the potential to serve as a valuable tool in the U.S. Army's initiative to enhance overall operations, minimize developmental costs, and ensure safer operations by the use of computer generated models.

REFERENCES

1. Salafia Dominick, DeLeon Norberto and Outlaw James F. , "Automation of Wear Analysis for Large Caliber Weapons". AIP Conference Proceeding #497, pp346-351, June 1999 2. Davis, Terry L. Lead NDT Inspector, USAYPG, verbal consultation 3. Sandwell, David T. "Biharmonic Spline Interpolation of GEOS-3 and SEASAT Altimeter Data", Geophysical Research Letters, 2, 139-142, 1987. 4. Watson, David E., Contouring: A Guide to the Analysis and Display of Spatial Data, Tarrytown, NY: Pergamon (Elsevier Science Inc.): 1992.

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

371

STUDY OF VISIBLE AND INFRARED METHODS OF BURIED DEFECTS UNDER ATMOSPHERIC ENVIRONMENT BY MEANS OF THERMAL IMAGE TECHNIQUES

Y. OKAMOTO, A. KAMOI and T. KONDO University of East Asia, 2-1, Ichinomiya-gakuencho, Shimonoseki City, Yamaguchi-ken, Japan V.P.VAVILOV Tomsk Polytechnic University, 3, Savinykh Street, 28, Tomsk, Russia

ABSTRACT Visible and infrared systems are widely used to detect invisible buried defects of industrial structures and machineries using remote sensing NDT devices. Several imaging techniques are applied to analyze the location, dimension and its deterioration rate of the buried objects of engineering structures like piping, vessel, building, land mine, ancient heritage under atmospheric environment. In this paper, experimental and numerical detection characteristics of the buried models were estimated. It was verified that the detection limit of the model was largely influenced by temperature deviation above the object by the conduction heat flows. The test results and its heat transfer mechanism pertaining to the presented detection techniques are also summarized and analyzed basically and quantitatively. KEYWORDS Buried defect, remote sensing NDT device, visible and infrared method, heat balance equation, thermal image technique

INTRODUCTION Visible and infrared multi-spectrum instrumentation systems are widely used to detect invisible buried objects of industrial structures and machineries as remote sensing NDT devices [1] . Thermal image method was carried out to analyze the location and dimension of the objects numerically. Several imaging techniques of the buried objects like radar, magnetic flux and electric current tests are applied to detect the location, its dimension and deterioration rate of the buried objects of engineering structures like piping, vessel, building, land mine, ancient heritage under atmospheric environment. However, those methods cannot be applied under detached condition [2] . Generally, the underground-buried objects like the ancient remain, civil basement and land mine are occasionally laid down in the ground and they are invisible from the surface. A noble detached sensing method, so called the thermal imaging detection method, was conducted to visualize the abnormal temperature distribution of the ground surface above the buried objects using the

372

Y. Okamoto et al.

abnormal temperature distribution of the ground surface above the buried objects using the infrared radiometer [3J . In this paper, experimental and numerical detection characteristics of the buried models were preliminary estimated as a reverse problem. It was verified that the detection limit of the objects is largely influenced by temperature deviation above the object by the conduction heat flows along the surface of the ground soil. The test results and its heat transfer mechanism of rectangular and cylindrical models pertaining to the presented detection techniques are also summarized and analyzed basically and quantitatively.

TEST APARUTTUS The square test piece made of the polystyrene the thermal conductivity of which is approximately similar to the air object had been used to simulate the buffed model of the ancient tomb and civil basement. The cylindrical test piece made of the metal and plastic is equivalent to the mock-up model of the near-underground landmine. Buried model test pieces made of polystyrene and metal are settled in the ground soil composed of clay and sand grain of 0.1-1.0 mm in diameter. The test pieces are laid down in the ground to model for a practical application. The radiation temperature distribution of the test piece is measured using the infrared Hg-Cd-Te radiometer, the wave length of which is 8 to 13 u m and the thermograph of the IR is recorded in the data recorder. The radiation temperature distribution of the object is displayed on the CRT of the IR. The radiation temperature of the ground soil is consistent with the true one, because the value of the emissivity of the soil surface is nearly unity. Solar and artificial radiation heaters radiate the heat to the ground surface at the heat flux q to detect the buffed invisible object, as shown in Fig. 1. The experiment is performed with adding the thermal energy in course of daytime for solar heating. The artificial radiation heater radiates the heat to the ground about 1 hour and cools the ground surface after finishing the heating. Ambient temperature of Tg, solar heat flux and wind velocity is continuously measured in course of operation time.

heater with reflector radiat,on ~.tlng " " IR

IR "

solar heating ,

!-- lm--~

. grouncl surface /

I

/ .............. L

(a)

/ ~o$-/fi 3m -----_4 ~

metal

polystyrene

(b)

Fig.1 Surface heaters; (a) for artificial radiation heating; (b) for solar heating

Thermal Image Technique

373

NUMERICAL EQUATION Numerical calculation on the heat transfer mechanism of square and cylindrical models was carried out using finite element and differential methods. Figure 2 shows the calculation model of the buried object in the ground soil space [4~ . The buried rectangular and cylindrical defects, thermo-physical properties of which are differed from that of the surrounding space ABCD.

~

I

injected heat flux ~ t ~ ~

A

" B

E ~ H

C

I environment air buried

soil space

~-

'T~..

~- X

[

D

X0

y

Fig. 2 Calculation model of the buried object Solar and artificial radiation heaters radiate the heat to the surface AB at the heat flux q, as already shown in Fig.1. The injected heat q is transferred to the soil space by the heat conduction and environment space by heat convection and conduction. Temperatures of the space Tl(x, y, r ) and rectangular defect Tz(x, y, r ) are expressed by heat balanced equations.

a~(0 2Tl/ 0 X2+ 0 2T,/ 0 y 2) __ G3Tl/ 3 r a 2 ( 0 2T2/ 0 x 2 +

0 2T2/ 0

y 2) =

0 T2/ 0 z

(1) (2)

where a~ and a2 are thermal diffusivity, x is horizontal axis, y is vertical axis and r is time. Temperatures of the axial symmetrical space and cylindrical defect Tl(r, y, r ) and T2(r, y, r ) are shown in the followings. al(0Tl/ 0 r + l / r . 0 r , / Or+ 02Tt/ 0y2) = C3Ti/ G 3r (3)

a2(OT2/ 0r+l/r" 0T2/ Or+ 02T2/ 0 / ) = 0T2/ c3 r

(4)

where r is radius. Boundary conditions for each surface and wall are shown in at AB surface;-a. ~0 Tz/c3y+(c~r C~r)(Ti- Tg) = q( r )

(5)

q( r ) is represented by 550 sin n( v- 8)/24 (kcal/hr m 2) for solar heating and constant for radiation heating. Boundary conditions are shown in at side and bottom of space BCD;

~ ~ c3T~/0 n = 0

(6)

374

Y. Okamoto et al.

at boundary surface of space and defect; ~, ~0 T1/0 n = X2 0 T2/0 n

(7)

where a c+a rare the convective and radiative heat transfer coefficient, T gis the temperature of gas, n is normal direction of the boundary BCD. The governing equations and boundary conditions are numerically solved using the control volume and finite element methods.

TEST RESULT EXPERIMENTAL RESULT

Figure 3 shows the thermograph of the soil surface that bury 3 polystyrene plates (a), (b) and (c) of 10, 20, 30 mm in width and 10 cm in depth. The soil temperature on the buried plate becomes larger than that without the plate by the solar heating, because the expansion heat flows in the soil around the plate.

Fig. 3 Thermograph of 3 polystyrene plates Figure 4 shows the thermograph of the buried aluminum cylinder by the halogen lamp heating. The temperature above the aluminum cylinder becomes smaller than that without the cylinder, because the contraction heat flows in the ground along the cylinder.

Fig. 4 Thermograph of aluminum cylinder

Thermal Image Technique

375

Figure 5 shows the photograph of the ground surface of the buried aluminum cylinder taken by a visible CCD camera. The left end of the inclined bar shows the location of center above the buried cylinder. The scanning line of the under the Fig. 5 shows the brightness distribution of the ground surface. Change of the brightness in the center position above the buried cylinder shows the existence of the buried object. However, the detection limit using the CCD camera is inferior to that of the infrared thermograph, as shown in Fig. 4.

Fig. 5 Photograph of the buried cylinder using CCD camera

THERMAL-IMA GE DETECTION ANAL YSIS Figure 6 shows the temperature trend of the ground surface above the buried styrene plate of 40 cm in width and 10 cm in depth by the solar heating. The temperature of surface above the plate Tc becomes larger than that without the buried object Ts in daytime and smaller than that in night. Therefore, the temperature difference ATc=Tc-Ts becomes positive in daytime and negative in night, because of expansion flow and contraction flow around the object respectively.

40 ,-., o

30

~

5.0 ,~Tc,,,~= 0.98 "C

4.0 ._. ,._) 3.0 ~

7,20

2.0

I00~ ' ----~"/,

~.__ d~',., , "'~ "-010

6:00

18!00 '

12':00 ' .

24 !00 '

: 00

time (o' clock) Fig. 6 Temperature trend of the ground surface above the buried plate

376

Y. Okamoto et al.

The temperature distribution of the buffed metal cylinder, as shown in Fig. 4 is numerically analyzed by solving equations (3) and (4). It is verified from the analysis that the numerical temperature trend solved by the equations is similar to the experimental temperature. Table 1 represents experimental and numerical results of the heat flow and generated temperature, which depend on difference of the thermal conductivity of the soil space ~ ~and buried object ;l 2. The heat flow pattern and generated temperature difference ATc=Tc-Ts are mainly determined by relative values of ~1 and ~.2 It is remarked that the heat flow and generated temperatures during heating process become reverse in case of cooling process. Table 1 Heat flow and generated temperature depending on thermal conductivity ;t, 1 and ~ 2

~, l; Thermal conductivity of soil space

~, 2; Thermal conductivity of buried object

Temperature deviation above the buried object by the solar heating method is generated by the conduction heat flows around the object. The temperature gradient above the buried object becomes non-uniform, because contraction and expansion flows are generated around the object.

CONCLUSION Thermo-physical property of the ground soil around the surface and buried object was analyzed experimentally and numerically using the thermal image technique. The generated abnormal temperature distribution in course of heating is numerically solved by the heat transfer equation in the soil surface. When the heat is injected to normal direction of the surface, the surface temperature distribution above the object becomes concave and convex. On the basis of described fact, we concluded that the thermal image method is able to analyze the detection limit of the buried object in the soil space.

REFERENCES 1. Okamoto Y., Kamoi A. and Ishii T., (1998), Quantitative Infrared Testing, Eurotherm, 12, pl 9-20 2. Okamoto Y.,Inagaki T.,(1995),Remote-Sensing Thermal Image Method, Corona publisher 3. Okamoto Y., Agu H., T.Inagaki T., (1997), SPIE Thermosense, 3056, p33-41 4. Ingaki T. and Okamoto Y., (1995), Int. Conference of Physical Protection, Tokyo, p 1-8

MICROSTRUCTURE

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

379

CONTACTLESS MEASUREMENT OF INDUCTION-HARDENING DEPTH

BY AN AXIAL-SHEAR-WAVE EMAT M. HIRAO, H. OGI, and Y. MINAMI

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

ABSTRACT We describe the method of electromagnetic acoustic resonance to detect the resonance frequency shift of the axial-shear wave propagating around the circumference of steel cylinders, which is caused by the induction-hardening treatment.

Using the magnetostrictive mechanism, we

design an electromagnetic acoustic transducer (EMAT) to generate and receive a high-frequency axial-shear wave and to carry out a contactless resonance spectroscopy measurement.

Mea-

sured spectral response allows reconstructing the surface layer profile in conjunction with the linear perturbation scheme and the single-layer modeling for the effect of material's radial inhomogeneity.

Inferred case-depth favorably compares with the destructive measurement of

hardness profile within 0. l mm error. KEYWORDS: Resonant frequency, axial-shear-wave EMAT, steel cylinder, induction hardening, linear perturbation scheme

INTRODUCTION Induction hardening is a standard surface modification method for steel products such as shafts and gears.

Metallurgical phase transformation produces a case of hardened metal and a com-

pressive residual stress there, both of which protect from wearing and suppress the nucleation and growth of fatigue cracks during the service. the case depth evaluation. it is destructive.

There are currently two testing techniques for

One is the direct measurement of hardness on the cross section, but

The other detects the ultrasonic backscattering signals from the grain structure

change in the steels, which is less laborious and nondestructive [1,2].

Because the core metal

has larger grains, the core-to-case interface can be thus detected in an A-scope display.

How-

M. Hirao, H. Ogi and Y. Minami

380

ever, it requires the parts to be immersed in a water tank for acoustic coupling, making the on-line inspection inapplicable.

Another shortcoming is that the dominant front-surface echo

masks the signal from the shallow interface and this technique is unavailable to a thin layer. Seeking a noncontact ultrasonic method for evaluating modified surface layers, this study proposes a new method of using the resonance spectroscopy with the axial-shear wave.

It is an

elastic wave, which travels along the circumference of a cylindrical surface with the axial polarization [3-6].

Resonance ultrasound spectroscopy is combined with an electromagnetic

acoustic transducer (EMAT) to measure the shift of the resonance frequency caused by the induction hardening of steel cylinders.

Metallurgical modification accompanies the changes in

the mass density and elastic moduli.

Their changes produce the radial variation of the shear

wave velocity, which affects the spectrum pattern of the axial-shear wave resonance.

We

analyze the resonance frequency shi~ with the linear perturbation scheme to reconstruct the surface-layer profile.

Capability of incorporating an EMAT in a contactless resonance spectro-

scopy measurement enables developing a less consuming, more accurate, and robust inspection technique for steel cylinders.

Inferred case-depth agrees with the result of destructive hardness

measurement within 0. l mm error.

PERTURBATION SCHEME FOR AXIAL-SHEAR-WAVE RESONANCE An axial shear wave is a pure shear wave trapped near the free surfaces of cy!indrical rods (and pipes) and is available using EMATs.

In the cylindrical coordinate system (r, O, z), the axial

shear wave carries only the axial displacement, w, and travels in the circumferential direction. Displacement w is independent of z, that is, w=w(r, 0). -

~

The governing equation is +

w=0,

(1)

r

where p is the mass density and tx the shear modulus.

For homogeneous p and Ix, a harmonic

solution w= AJn(kr)exp[i(o3t-nO)] is possible, when the resonance condition J'(kR)=O or

nJn(kR)-kRJn+l(kR)=O, is satisfied. Here, k=colV~is the wavenumber with the shear wave velocity V~ = ~/p/p.

(2)

The

stress-free boundary condition at the outer surface r=R for radius R of cylinder is implied. Index n defines the periodicity around the circumference and equals the number of turns of the meander-line coil used in the EMAT.

Geometry of the experiment will be shown later.

Axial-Shear-WaveEMATfor Case-DepthMeasurement There are an infinite number of resonance frequencies that satisfy Eq.(2) for each n. them by fm(')using another integer m to identify the overtones. to the surface area for the fundamental mode of m=l. with n.

381 We denote

The oscillation is concentrated

The degree of concentration increases

The fundamental mode provides a measure of the material modification in the surface

area and the higher modes detect the inner region.

This unique nature of axial-shear wave is the

key for characterizing the radial inhomogeneity of material properties by systematically selecting the penetration into the material [3-6].

Fig. 1. Analytical model for the axial-shear wave resonance in case-hardened cylinder. For interpreting the measured shill of resonance frequencies and inferring the case depth, we consider two cylinders of different p and Ix, but subjected to the same n, R, and boundary condition.

Their density and elastic modulus, (p0, go) and (91, Ix1), may change in an axisymmetric

way (Fig.l).

The axial-shear wave accompanies the displacements w0 and Wl in them, which

are the solutions ofL(w0; p0, go)=0 and L(Wl;pl, Ix0=0. We derive the resonance frequency shill following the perturbation theory due to Auld [7], which is based on the reciprocal theorem of linear elasticity and the first-order approximation in terms of small deviations.

Calculating the identical equation

i

po,m)w,aS- t(w,, p,,u,)woaS= o,

(3)

with Eq.(1), partial-integrating, and using the stress-free boundary condition at r=R, we have

Aco

I~r{-(APlj~(kr)+(~Po3[J'-~2(kr)+J'+12(kr)]}dr =

COo,,

\ Po )

-

(4)

2I~ rJ:
in which A~Aoo is the normalized difference between the resonance frequencies observed in the

M. Hirao, H. Ogi and Y. Minami

382 two cylinders.

We assumed that p0 and ~ are homogeneous and the material perturbations

Ap(r) and A~t(r) from them are small enough as is the case in the surface hardening treatments. We neglected their higher order terms.

This involves making w0--Wl in the integrand, and then

substituting wo.-d,,(kr). We infer the case depth from the measured Aco/co0with a supplementary modeling for the radial inhomogeneity.

In case of induction hardening of steels, the martensitic transformation

saturates within a surface layer and then a step-function model is appropriate, where the matrix (core) is overlaid with a layer of hardened metal (case) of a uniform thickness h and a constant perturbation, (Ap, Alx). With the above model, Eq.(4) simplifies to the form Ac~ --~ooI

,,m

= A (h / R ;n, m)(AP 1 + B(h/R;n,m)( A'u 1 \ po J k, po J

(5)

We see that Aco/cooof individual resonance mode depends on h/R as well as n and m in the functions A and B.

Values for Ap/po and AB/~ should be calibrated for the test steel.

800

,

A

,

,

'

,

'

,

,

,

,

,

--" hDEs=0.70mmI ---O-- hDES=2.62ram !

600

m 400 O

~

2oo

,

1

I I

I I

I I ,,

0

I

2

I I

,

-

,

I

4

,

I

,

6

I

,

8

Distance from surface (mm) Fig. 2. Vickers hardness distribution measured by indentation at 0.2-kgf load.

RESONANCE SPECTROSCOPY MEASUREMENTS We prepared six specimens, 30mm diameter and 200mm long, from a drawn carbon-steel rod (C: 0.47 mass%).

Their middle parts of 100mm long were given a series of induction hardening

(not tempered)to introduce the surface layers, whose thicknesses range from 0.70mm to 3.06mm. Figure 2 exemplifies the nearly stepwise distribution of Vickers hardness.

Axial-Shear-Wave EMATfor Case-Depth Measurement

383

solenoid coil /~

meander-line coil

ILJJJU superheterodyne spectrometer Fig. 3. Magnetostrictively coupling EMAT for axial-shear-wave resonance. Figure 3 is a schematic of the present EMAT devised to couple to the axial shear wave using the magnetostrictive effect of ferromagnetic metals [4]. The EMAT consists of two coils.

The

specimens were inserted in the solenoid coil, which supplied a bias field around 0.03 T along the length.

We used the meander-line coil of 20mm wide, which was printed onto a flexible poly-

mer film at a fixed spacing of 6=2rd~n -~l.29mm with n=73.

It was loosely wrapped around the

specimen's center and the ends were fastened with each other, providing a noncontact acoustic coupling.

The meander-line coil was driven by rf bursts, typically 100Vpp and 150gsec duration,

to induce the alternating field in the circumferential direction on the specimen surface, which was superimposed on the static axial field. Then, the surface elements beneath the coil under-

,-:,.

'

'

11

!

<

'

'

L!

'1

i!

o~/

li

.................. ~ ........ ~l/ --~-- ~ i_ ~ ~ ~l ~ ~ 'L ~l| ~ ~

It i

"

I

2.6

2.8

3

3.2

3.4

3.6

Frequency (MHz) Fig. 4. Resonance spectra measured before and after the heat treatment.

M. Hirao, H. Ogi and Y. Minami

384

went shear deformation responding to the obliquely oriented total fields, which generated the shear waves of the axial polarization.

Reversed magnetostrictive mechanism altered the mag-

netic field in the specimen surface, which was picked up by the same coil to form the received signals.

Swept-frequency operation, typically taking ten seconds, measured the resonance spec-

trum of each specimen; the details of superheterodyne analog circuitry and operations appear in references 8-10. In Fig.4, the spectra are compared before and after the induction hardening.

Agreement with

the resonance condition, Eq.(2), was confirmed using the untreated specimens.

We observe the

negative shift for all the modes, reflecting a decrease of shear wave velocity in the surface layer. Going into details, a large shift occurs for the fundamental mode (re=l) and less shifts for the higher modes for the thin surface layer. detected modes.

For the thick layer, nearly equal shifts appear in all the

This contrast illustrates the essential feature of the axial-shear wave resonance

for measuring the radial gradient of elastic properties.

We fit a Lorentzian function to the

spectral data around the maximum amplitude to obtain the resonance frequency from the center axis.

The shift is normalized with the original measurements and is summarized in Fig.5 up to

the sixth mode.

As h increases, Aco/co0of all the modes monotonously approach the saturation

value of-1.1%.

The decrease starts to occur first with the fundamental mode, the second mode

follows, then the third mode, and so on.

When the penetration is entirely within the surface

layer, (p, It) of the hardened steel determines the resonance frequency and the core metal has nothing to do with the observed resonance.

When the penetration is moderate, encompassing

o -0.5

$ <1 -1

0

1

2

3

hDE S ( m m ) Fig. 5. Resonance frequency shift of up to the sixth mode observed in case-hardened steel cylinders.

Axial-Shear-Wave EMATfor Case-Depth Measurement

385

both core and case, the resonance frequency is given by p and Ix of the two slightly different metals.

Equation (5) provides the weight between the two media as a function of h/R and the

amplitude pattern shown in Fig. 1.

CASE-DEPTH EVALUATION The Archimedes method gave Ap/p0 =-0.71% with a coupon specimen cut from the thickest surface layer.

We then calibrated AB/Bo=-2.88% using Eq.(2) along with this and A03/o~0=-1.1%

of the fundamental mode.

The martensitic transformation diminished both p and ~ with more

influence to IX, causing the decrease of axial-shear-wave resonance frequencies.

The remaining

unknown variable is h, which was determined so as to best satisfy Eq.(5) with the measured Ao~/o~o. Figure 6 compares the best fitting h against the case depth determined through the Vickers hardness measurement referring to Hv=400.

We observe an excellent agreement between them.

Mode selection via n and m is important to deduce the correct answer.

We have selected the

resonance modes, which exhibited the transitional shift for each specimen; m=l for h=0.70mm and m=2 for h=l.48mm, for example.

Otherwise, the resonance frequency shift led to inaccur-

ate value or was unable to yield an answer.

'

I

'

I

'

I

3

1

0

, ,

0

1

2

3

hDES (mm) Fig. 6. Comparison between the case depth measured with the acoustic resonance and from the Vickers hardness.

386

M. Hirao, H. Ogi and Y. Minami

We also used another meander-line coil of n=104 and 8=0.42mm, which covered less than a half of the circumference. Despite the decreased signal strength, this coil was capable of measuring the resonance spectra and resulting in essentially the same evaluation of the case depth, which is included in Fig.6.

CONCLUSION We have presented and discussed the theoretical analysis for the shill of the axial-shear wave resonance frequency caused by the small axisymmetric change of elastic properties. A linear perturbation scheme was demonstrated to be useful by the noncontact spectroscopy measurements with the induction-hardened steel cylinders.

Inferred case depth favorably compared

with the destructive measurements of Vickers hardness. The advantage over the conventional immersion technique includes the contactless measurement and the capability of evaluating the thin surface layers. Moreover, this ultrasonic method applies to other surface modification such as carburizing, nitriding, etc., because these treatments also cause small fractional changes of elastic properties and the basic formula (4) is valid. These aspects will facilitate the installation of this quick, accurate, and robust measuring method in actual production environment. It should be noted at the same time that the proposed EMAT technique requires a cylindrical surface and the axisymmetric material change. For an asymmetric layer, the frequency change will give a circumferential average. REFERENCES

1. Good, M.S., Metals/A4aterials Technology Series #8408-003, ASM Metals Congress 1984 (American Society for Metals). 2. Fujisawa, K. and Nakanishi, A. (1989) Nondestr. Test. Comm., 4, 131. 3. Johnson , W., Auld, B.A., and Alers, G.A., Review of Progress in QNDE, Vol. 13, edited by D.O Thompson and D.E.Chimenti (Plenum, New York, 1994), p.1603. 4. Ogi, H., Hirao, M., and Minoura, K., Review of Progress in QNDE, Vol.15, edited by D.O Thompson and D.E.Chimenti (Plenum, New York, 1996), p.1939. 5. Ogi, H., Hirao, M., and Minoura, K. (1997) J.Appl.Phys., 81, 3577. 6. Ogi, H., Hamaguchi, T., and Hirao, M. (2000)Metall. andMater. Trans., 31A, 1121. 7. Auld, B.A., Acoustic Fields and Waves in Solids, Vol.2 (Wiley-Interscience, New York, 1973). 8. Hirao, M., Ogi, H., and Fukuoka, H., Rev. Sci. lnstrum., 64, 3198 (1993). 9. Petersen, G.L., Chick, B.B., Fortunko, C.M., and Hirao, M. (1994) Rev. Sci. Instrum., 65, 192. 10. Hirao, M. and Ogi, H. (1997) Ultrasonics, 35, 413.

Nondestructive Characterizationof MaterialsX Green et al. (Eds) 9 2001 ElsevierScienceLtd. All rights reserved.

387

T E M P E R E M B R I T T L E M E N T OF D U P L E X STAINLESS STEEL AND ITS N O N - D E S T R U C T I V E E V A L U A T I O N

M.Katoh*, K.Nishio*, T.Yamaguchi*, Y.Matsumoto*, H.Ikeda** *) Kyushu Institute of Technology 1-1, Sensui-cho, Tobata, Kitakyushu, 804-8550, Japan E-mail: [email protected] **) Kagoshima Technical College 1460-1, Hayato-machi, Aira-gun, Kagoshima-ken, 899-5102, Japan E-mail :hikeda@ kagoshima-ct.ac.jp

INTRODUCTION Duplex stainless steel has been widely used for materials for a primary cooling system of atomic power plant and equipments for petrochemical plants. It has good tensile strength, weldability, corrosion resistivity and toughness. On the other hand, it has the following bad points. That is, when it is held for a long time at temperatures lower than 500 ~ both its proof strength and hardness tend to increase and its toughness tends to decrease [1]. These have been attributed phenomena to spinodal decomposition [ 1]. In this paper, we examine the phenomenon of both base metal and weld metal of duplex stainless steel aged at 450 ~ at which spinodal decomposition will be accelerated. First, we confirm the fluctuation of chromium during aging by simulating the spinodal decomposition of Fe-Cr binary system using the Monte Carlo method. We make clear that the hardness change during aging is more significant in ferrite than in austenite and the toughness decreases with the aging. In order to reduce the harmfulness of ferrite, we perform plasma welding using argon plus nitrogen gas. Since nitrogen is an austenite forming element, we can increase the ratio of austenite in the weld metal and obtain the better toughness during aging. The process during aging can be evaluated by the change of the longitudinal wave velocity : the lower the wave velocity, the higher the toughness. MATERIALS USED AND EXPERIMENTAL PROCEDURE The material used in this study is duplex stainless steel SUS329J4L (thickness:4 mrfi), which contains 0.016 C, 0.49 Si, 0.80 Mn, 0.015 P, 0.001 S, 7.11 Ni, 25.37 Cr, 3.03 Mo, 0.13 N and Fe (balance) in mass percent. It consists of both austenite and ferrite, whose area ratio is about one to one as shown in Figure 1. Welding was performed with plasma arc (welding current:250 A, welding speed:150 mm/min) in a chamber. The atmosphere in the chamber were changed as follows: (1) argon, (2) nitrogen gas being injected from the side of a welding torch in argon atmosphere (referred to as argon + nitrogen injected), and (3) mixed gas of 70% argon + 30% nitrogen. Fig.1 A microstructure of duplex After aging the base metal and weld metal stainless steel used

388

M. Katoh et aL

for 1, 10, 50 and 100 h at 450 ~ microstructures were observed using an optical microscope, the distribution of elements was analyzed using EPMA, and Vickers hardness and impact test were performed. Figure 2 shows the shape and dimensions of an impact test specimen (thickness:4 mm). For the weld metal a notch was machined in the center of the weld metal. Moreover, the change in wave velocity

45 ~

! I

55

Fig.2

during the aging was measured using the B1-B2 method.

Shape and dimensions of an impact test specimen

SIMULATION OF SPINODAL DECOMPOSITION We simulated the process during aging using the Monte Carlo method, in which we considered the motion of atoms as a probability process and determined the arrangement of atoms so that the total energy of the system was minimized. The procedure of the simulation is as follows: (1) construct body-centered cubic lattices of Fe-Cr binary alloy; (2)arrange Fe atoms at each lattice point and arbitrarily change Fe atoms corresponding to chromium concentration of 10, 20 and 30 atomic percent into Cr atoms; (3) select one Cr atom at random and select one atom from 8 nearest neighboring atoms; (4) change the positions of these two atoms; (5) obtain the energy change in this process and decide whether these atom positions are adopted or not by comparing the probability W-

exp( - AE / kT)

(1)

1 + exp( - AE / kT) with a random number, where k:Boltzmann constant (1.38066x10-23 J/K), AE:energy change (J) and T:absolute temperature (K). The steps from (3) to (5) were considered to be unit operating time and we performed this process 1.0x105, 3.0x105, 5.0x105 and 10x105 times. Since we considered body-centeredcubic unit cells of 20x20x20, the number of atoms was 16,000 and we used a periodic boundary condition. The temperature was taken to be 300, 400 and 450 ~ Figure 3 (a) and (b) show examples of the change of Cr-Cr atomic pairs during the operating time by two dimensional representation when Cr percent is 20% and the temperature is 450 ~ for the operating time of 0 and 10x105, respectively. Chromium atoms which are initially arranged at random in iron atoms tend to fluctuate when the operating time increases. Figure 4 shows a quantitative example of the results obtained, in which the relation between the ratio of Cr-Cr atomic pairs and the operating time for the case of chromium percent being 20% is shown. The ratio of Cr-Cr atomic pairs to the nearest atomic pairs increases as the operating time increases. This phenomenon is significant when the aging temperature is 450 ~ and the chromium percent is 20%. Thus, we could confirm the fluctuation of chromium during the aging. According to the spinodal decomposition the ferrite phase separates into the 0~ phase with high concentration of chromium and the 0~' phase with low concentration of chromium.

NDE of Duplex Stainless Steel during Aging

Fig.3

389

Change of Cr-Cr atomic pairs during the operating time

INFLUENCE OF AGING AT 450 *C ON BASE METAL o 3 0 0 " C []

~

both the o~ and T phases in Vickers hardness test

o

after aging for 100 h at 450 *C. The size in the oc

~

phase is smaller than that in the y phase after the aging, and we can observe the larger change in

~

6

hardness in the ~ phase than in the 7 phase because

~

4

the sizes of both the a and T phases are almost the same before the aging. Figure 6 shows the relation between Vickers hardness and aging time. Both the

~

2

(z and T phases have the same Vickers hardness of 325 before the aging, and the hardness in both phases tend to increase with the increase in aging

Fig.4 Relation between the ratio of Cr-Cr atomic pairs and the operating time (Fe-20%Cr)

t~

Fig.5

lo

450"C

Figure 5 shows the change in indentations in

z~ 4 0 0 " C

8

,

I

,

I

,

I

,

I

t

2x10s 4x10 5 6xlO~ 8x10 s lOxlOS OPERATING TIME

Change in indentations in both the a and 3' phases in Vickers hardness test after aging for 100 h at 450 oC

390

M. Katoh et al. SIN)

,

I

'

I

'

I

'

I

'

'

I

'

I

'

I

'

9

I :

' o

Austenite 400

,~-~--~

9

~ 350 , , Aging temperature : 450"C ,

0

I:

20

,

I

40

,

i

60

,

I

,

,

80

AGING TIME, t Oh) Fig.6 Relation between Vickers hardness and aging time

100

0

I

,

I

,

1

,

I

20 40 60 80 AGING TIME, t (h)

,

100

Fig.7 Relation between Charpy impact value and aging time at 450~

time. Though the hardness in the T phase tends to saturate to about 400 after the aging time of about 10 h, the hardness in the 0~ phase continues to increase larger than 450 after the abrupt increase until the aging time of about 30 h. Figure 7 shows the relation between Charpy impact value and the aging time at 450 ~ The impact value of the base metal is initially 300 J/cm2 ; it decreases with the increase in aging time and has the value of about 60 J/cm2 after the aging of 100 h. Among the aging temperatures of 350, 400 and 450 ~ the decrease in Charpy impact value with the increase in aging time was the largest at 450 ~ Figure 8 (a) and (b) show SEM photographs of fracture surfaces before and after the aging for 100 h at 450 ~ respectively. Though we can observe dimple patterns before the aging, cleavage fracture surface is partly observed after the aging for 100 h. Microscopically no change was observed before the aging nor after the aging by an optical microscope.

Fig.8 SEM photographs of fracture surfaces after the impact test of the base metal

INFLUENCE OF NITROGEN ON EMBRITTLEMENT OF WELD METAL Since the change in the hardness during the aging is more significant in the o~ phase than

NDE of Duplex Stainless Steel during Aging

391

in the T phase, it is considered that the embrittlement due to the aging might be reduced by increasing the ratio of 7 phase in the weld metal. It is reported that the amount of ~5phase in austenitic stainless steel weld metal decreases when GTA welding is performed by shielding using mixed gas with argon and nitrogen gas [2]. Hence, we tried to decrease the amount of ferrite in the duplex stainless steel weld metal by mixing nitrogen gas. Welding was performed by plasma welding.

Fig.9 Microstructures near the fusion boundary of the weld metal Figure 9(a) and (b) show microstructures near the fusion boundary when argon and mixed gas of 70% argon + 30% nitrogen are used as the shielding gas, respectively. In Figure 9(a), the 0~ phase is coarsened and the 7 phase is observed along the grain boundaries of 0~ phase. The similar microstructure is observed when shielded with argon + nitrogen injected. The weld metal grows epitaxiaUy on the grains in the heat-affected-zone. The amount of r phase in the weld metal is about 60% in both cases. But quite a different microstructure is observed in Figure 9(b). The amount of V phase in HAZ increases more than that in Figure 9(a), and the amount of r phase in the weld metal is about 31%. The content of nitrogen in Figure 9 (a) and (b) is 0.116, and 0.304 mass%, respectively. Thus the increase in the amount of r phase in Figure 9(b) is due to the increase in nitrogen content, and nitrogen is an austenite forming element. Figure 10 shows the relation between Vickers hardness in three weld metals and aging time. Before the aging, the hardness is almost the same in the three weld metals and the base metal. When the weld metals are shielded with argon and argon + nitrogen injected, the hardness increases to 440 after aging for 100 h and this tendency is almost the same as the base metal. For the weld metal shielded with mixed gas of 70% argon + 30% nitrogen, however, the hardness is 345 after aging for 100 h, and this weld metal is softer than the other two. The change in the hardness is due to the difference of the amount of 7 phase in the weld

SS0 ~,

Aging temperatere : 450~ Ar + ~m ~ i n ~ ~ /

~4~

metal

~ 350 Ar 70%+ Nz gas 311%

250 0

20

48

60

80

100

AGING TIME, t (h) Fig.10 Relation between Vickers hardness in three weld metals and aging time

392

M. Katoh et al.

metals. This also affects the impact value as will 200 I I I I' Aging temperature : 450~ be mentined next. m O All" Figure 11 shows the relation between 150 Charpy impact value and aging time. The impact zx Ar + N~ injected values as welded are 165, 170 and 190 J/cm2 for rd 3O% the weld metals shielded with argon, argon + ~ 100 nitrogen injected, and mixed gas of 70% argon ;~ +30% nitrogen, respectively. Among the three ~ 5o weld metals, the one shielded with mixed gas of 70% argon + 30% nitrogen has the largest impact ~, I I I I, value for both the as welded and the one after the o 20 40 60 80 100 aging for 100 h, although the impact value tends to AGING TIME, t (h) decrease with the increase in the aging time. For Fig.11 Relation between Charpy the weld metals shielded with argon and argon + impact value and aging time nitrogen injected, the tendency of the decrease in the impact value with the increase in the aging time is similar. Figure 12(a) and (b) show SEM photographs of fracture surfaces of the weld metals shielded with argon and the mixed gas of 70% argon + 30% nitrogen after aging for 50 h, respectively. In Figure 12(b) dimple patterns are observed, while cleavage fracture surface is obseved in Figure 12(a).

Fig.12

SEM photographs of fracture surfaces after the impact test of the weld metal aged for 50 h at 450 ~

Figure 13(a) and (b) show microstructures for the as welded when shielded with argon and mixed gas of 70% argon + 30% nitrogen. In Figure 13(a), the T phase is observed along the grain boundary of matrix of the cz phase. A similar microstructure is observed when the as welded metal is shielded with argon + nitrogen injected. In Figure 13(b), however, both the tz and ), phases are uniformly distributed and chromium and nickel are enriched in the 0~ and ~, phases according to the analysis with EPMA, respectively. No significant change in microstructure is observed after aging for 100 h. CHANGE IN WAVE VELOCITY DURING AGING Figure 14 shows the relation between longitudinal wave velocity and aging time for the

NDE of Duplex Stainless Steel during Aging

393

Fig.13 Microstructures as welded base metal and the weld metals. In every case the wave velocity tends to increase abruptly in the initial stage of the aging and continues to increase

~ 5900 "~ ;~

gradually after that. The longitudinal wave velocity

;d $800 _

for the as welded state was 5,720 and 5,750 m/s for the weld metals shielded with argon and argon + nitrogen injected, respectively. On the other hand, the longitudinal wave velocity of the weld metal shielded with mixed gas of 70% argon + 30% nitrogen was 5610 m/s, which is lower than the other two weld metals during the aging. These changes in longitudinal wave velocity in three weld metals correspond with the change in hardness

~, $700 _ m > ,~ ..... I $60o

1

~

0

t

I

I

Ar + N2 gas injected

e

20

t

a

l

Ar 70%+ N2 gas 30% I

I

I

40

60

80

100

AGING TIME, t (h) Fig.14 Relation between wave velocity and aging time for base metal and weld metal

shown in Figure 10 and the amount of o~ phase in the weld metals. Wada et al. [3] analysed the change in wave velocity during the aging of duplex stainless steel and reported that the velocity increased during the aging due to the increase in the elastic constants of the o~ phase. The results obtained in this study corroborate the findings of Wada et al. Hence, it is concluded that the increase in longitudinal wave velocity during the aging is mainly due to the increase in hardness in the cz phase. According to the results of the present study, the longitudinal wave velocity tends to increase and the impact value tends to decrease with the increase in the aging time. This Shows that we can evaluate the temper embrittlement by using the change in longitudinal wave velocity. CONCLUSIONS (1) The fluctuation of chromium during the aging of the duplex stainless steel could be confirmed by using the Monte Carlo method. (2) The temper embrittlement of the base metal in the duplex stainless steel is more significant in the cz phase than in the ), phase.

394

M. Katoh et al.

(3) The temper embrittlement of the weld metal can be improved by increasing the amount of ~, phase through the process of using mixed gas of argon and nitrogen as the shielding gas. (4) The degree of the temper embrittlement can be evaluated by the change in longitudinal wave velocity : the higher the impact energy, the lower the wave velocity. ACKNOWLEDGENTS The authors would like to express their gratitude to Professor K. Yasuda of Polytechnical University for preparing weldments by plasma welding used for this study. REFERENCES 1. Y.Isobe, A.Kamimura, K.Aoki, F.Nakayasu, Characterization of Thermal Aging of Duplex Stainless Steel by SQUID, Proceeding of 13th Internatinal Conference on NDE in Nuclear and Pressure Vessel Industry, ASM International, (1995), 337-344 2. Japan Welding Society, Metallographic Arias of Steel, Special Alloy Steel and Non-ferrous Metal Welds, Kuroki Shuppan, (1984), p.265 3. T.Wada, T.Matsuzaka, Analysis of Ultrasonic Wave on Duplex Stainless Steel with Thermal Aging, Proceedings of JSNDI Fall Conference 1997, (1997), 101-104

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

MEASUREMENT OF ALUMINUM

395

O F T E X T U R E C O E F F I C I E N T W4oo ALLOYS BY RESONANCE EMATS

C.-S MAN*, X. FAN* and J.G. MORRIS t

*Department of Mathematics, and t Department of Chemical and Materials Engineering University of Kentucky, Lexington, KY 40506, U.S.A. K. KAWASHIMA

School of Engineering, Tokyo University of Technology 1404-1 Katakura, Hachioji, Tokyo 192-0982, Japan

ABSTRACT A recent study [1] on the textures in strip-cast AA3105 and AA3015 aluminum alloys shows that ultrasonic measurement of the texture coefficient W400 would provide a promising means for on-line monitoring of textures in these alloys. Ultrasonic measurement of W400 in aluminum alloys, however, has long been an outstanding research problem in nondestructive evaluation. A previous investigation by Man et al. [5] on strip-cast AA3105 and AA3015 alloys, where they used the ratio ~ of velocities of through-thickness longitudinal and shear waves to estimate the value of W400, indicates that thermomechanical processing of these alloys induces slight shifts in the isotropic limit of the velocity ratio ~ and, by inference, also in the isotropic elastic constants. Such shifts may lie at the root of the difficulties encountered in ultrasonic measurement of W400 in aluminum. In this paper we undertake a further investigation (using AA3015 samples) with two major modifications. First, with a view to applications in on-line monitoring of texture, the velocity ratio ~ of the samples were measured by a (noncontact) resonance EMAT system. Secondly, we tried to devise the thermomechanical processing histories of the samples to see if we could confirm a pattern, observed earlier by Man et al. [5], in the effects of thermomechanical processing on the shifts of the isotropic limit of ~. Our findings suggest that, with prior knowledge on the thermomechanical processing history of the aluminum sheet, on-line monitoring of W400 remains a promising possibility that merits further investigation. KEYWORDS Crystallographic texture, aluminum alloys, on-line monitoring, resonance EMATs.

C.-S. Man et al.

396 INTRODUCTION

A recent study [1] by Lit et al. on the textures in strip-cast AA3105 and AA3015 aluminum alloys shows that ultrasonic measurement of the texture coefficient W400 would provide a promising means for on-line monitoring of textures in these alloys. Ultrasonic measurement of W400 in aluminum alloys, however, has long been an outstanding research problem in nondestructive evaluation [2-4]. Recently Man et al. [5] investigated anew the feasibility of using the velocity ratio (cf., for example, Hirao et al. [6])

~

V33 _I)~+2#( (V31 + V32)/2 - - _ # -

-

I+4# 8v/2 2 , 1 + #(A + 2#) 35 ~ cvv4oo)

(1)

for the measurement of W400; here V33 denotes the speed of longitudinal wave propagating in the normal direction of a sample sheet; V31 and V32 are speeds of through-thickness shear waves with polarization in the rolling and transverse direction, respectively; A and # are the isotropic elastic constants, and c is the anisotropic factor. They examined the difficulties one would encounter if one would, in the context of previous attempts, use equation (1) for the evaluation of W400 in aluminum. In previous attempts, the constitutive equation in question was derived through the Voigt-Reuss-Hill averaging procedure, which delivers also explicit formulae that express the elastic constants A, #, and c of the polycrystal in terms of the second-order and third-order elastic constants of the reference cubic crystal. One major difficulty lies in the fact that these formulae with the available single-crystal elastic constants do not deliver sufficiently accurate estimates of the parameters A, p, and c of the polycrystal for equation (1) to be used for the determination of W400. Appealing to a phenomenological analysis [7] on the constitutive equations of weaklytextured polycrystalline materials, Man et al. pointed out that equation (1) remains valid, to within terms linear in the texture coefficients Wzmn, for all weakly-textured elastic materials if A, #, and c are interpreted as macroscopic, empirically-determined elastic constants pertaining to the particular orthorhombic aggregate of cubic crystallites in question. In the phenomenological theory, the constants A, # and c define the "elastic material", which may acquire different textures from various thermomechanical processing procedures. Differences in alloying elements and in microstructures which significantly affect the values of A, # and c lead to different "elastic materials" in this theory. Under this reinterpretation of equation (1), they suggested that we could exploit this equation for ultrasonic measurement of W400 as follows: Suppose we are interested only in a specific aluminum alloy and a class of thermomechanical processing sequences that it may undergo, under which this alloy can be taken, in all circumstances of concern, as the same "elastic material" with constant values of A, #, and c. To proceed, first we prepare a calibration line by measuring both ~ (ultrasonically) and W400 (by X-ray diffraction or other suitable methods) for samples of this alloy which have undergone various thermomechanical processing histories of the given class. We may estimate the slope of this line by micromechanical modelling, but we must determine the intercept ~/(A + 2#)/# experimentally. With this calibration line in hand, we may measure the velocity ratio by ultrasonics and then read off the corresponding W400. To test this idea, samples of strip-cast AA3105 and AA3015 hot bands were annealed and cold rolled under various conditions to simulate a broad range of texture conditions. The texture coefficient W400 of these samples was measured by X-ray diffraction and their

Resonance EMATS

397

values measured with piezoelectric transducers. The data points of the a versus W400 plot were found to scatter along two parallel lines with equation = 2 . 0 5 2 - 3.88W400, and

a = 2 . 0 3 8 - 3.88W400,

(2)

respectively, where the slope was estimated by using a self-consistent micromechanical model. Roughly speaking, the data points pertaining to samples which were sufficiently annealed and those which were sufficiently hardened by hot- or cold-rolling fall on the line with the larger and the smaller intercept, respectively. Apparently, the annealing and the rolling led to small changes in the isotropic elastic constants. In this paper we continue the investigation in [5] and (using AA3015 samples) carry it further by making two major modifications. First, with a view to applications in on-line monitoring of texture, the velocity ratio a of the samples were measured by a (noncontact) resonance EMAT system. Secondly, we tried to devise the thermomechanical processing histories of the samples to see if we could get more data points filling in the space between the two parallel lines.

EXPERIMENT The samples used in the present study were prepared from a commercial strip-cast AA3015 hot band that we received from Commonwealth Aluminum Corporation. The processing histories of the samples are listed in Table 1. To determine the texture coefficients of these samples, pole figures were determined by X-ray diffraction both at the surface and at the midplane of the samples. The texture coefficients Wl,n,~ (even l, up to 1 - 22) and the volume fractions of the main texture components were derived from the pole figures by ODF calculations. While the textures at the surface and at the midplane of the samples had notably different intensities, they showed the same trend in evolution under both coldrolling and annealing. The average of the surface and midplane W400 values is listed in the last column of Table 1. A resonance EMAT (or R-EMAT) system was used to measure the velocity ratio a. The early work on R-EMATs was done by one of the present authors [8-9]. By now R-EMATs are well known and are especially suited for measurement on thin metal sheets. The REMAT used in the present experiments was made from a SmCo permanent magnet and a race track shaped coil (see Fig. 1). The larger and the smaller diameter of the coil was 18 mm and 10 mm, respectively. The number of turns was 12. The rectangle portion of the coil, as shown by broken lines in Fig. 1, was used and the other portion was bent and shielded. When the R-EMAT is placed above an aluminum sheet, it produces a static magnetic flux B that has both the z-component Bz and the x-component Bx (see Fig. 1, where the chosen Cartesian coordinate system is shown). An eddy current I is generated inside the sample by high frequency coil currents. The eddy current I has only the ycomponent. The interaction between I and Bx and that between I and Bz create the Lorentz force Fz and force Fx normal and tangent to the sample surface, respectively. The forces Fz and Fx produce the through-thickness longitudinal wave and the shear wave polarized perpendicular to the straight coil wire lines, respectively. Each wave is detected by the same R-EMAT through the reverse physical processes. When the R-EMAT is placed above a metal sheet so that the coil wire lines are parallel to the rolling direction, which is a two-fold axis of rotational symmetry of the orthorhombic sheet, the longitudinal wave (L) and the shear wave (St) polarized perpendicular to the rolling direction are generated.

398

C.-S. Man et al.

Table 1. Experimental data on AA3015 samples. Sample A E I M B F J N C G K O D H L P 1 2 3 4

Processing History Velocity Ratio a as received hot-band 2.0536 annealed A (750~ x 0.5 hr) 2.0473 annealed A (750~ x 2 hr) 2.0323 annealed A (900~ x 2 hr) 2.0288 cold-rolled A (30% reduction) 2.0530 cold-rolled E (30% reduction) 2.0424 cold-rolled I (30% reduction) 2.0373 cold-rolled M (30% reduction) 2.0279 cold-rolled A (60% reduction) 2.0531 cold-rolled E (60% reduction) 2.0468 cold-rolled I (60% reduction) 2.0450 cold-rolled M (60% reduction) 2.0331 back-annealed C (650~ x 4 hr) 2.0510 back-annealed G (650~ x 4 hr) 2.0471 back-annealed K (650~ x 4 hr) 2.0403 back-annealed O (650~ x 4 hr) 2.0356 back-annealed (500~ x 4 hr) and 2.0445 then cold-rolled (60% reduction) J back-annealed (500~ x 4 hr) and 2.0396 then cold-rolled (60% reduction) N back-annealed (500~ x 4 hr) and 2.0393 then cold-rolled (40% reduction) O back-annealed (500~ x 4 hr) and 2.0481 then cold-rolled (60% reduction) F

W400 x 103 -1.79 0.53 2.38 7.71 -2.98 2.46 4.55 5.99 -2.88 -0.87 1.08 3.71 -2.15 -0.57 3.73 4.75 0.18 2.61 1.85 -1.26

The longitudinal wave (L) and the shear wave (Sr) polarized parallel to the rolling direction are generated after we rotate the R-EMAT by 90 degrees about the sheet normal. Some of the measured results are shown in Fig. 2 (sample B) and Fig. 3 (sample F). In these figures, the thick lines pertain to the St-wave and the L-wave, and the thin lines pertain to the Sr-wave and the L-wave generated and detected by the R-EMAT placed with its coil wire lines parallel, and perpendicular to the rolling direction, respectively. The labels Sr3 and Sr4 denote the third- and the fourth-order resonances of the St-wave, St3 and St4 the third- and the fourth-order resonances of the St-wave, and L2 the second-order resonance of the longitudinal wave. In Fig. 2, Stn and Srn (n = 3, 4) are separated by some frequency difference, which reflects the difference in the two shear-wave speeds for sample B. On the other hand, thick line L2 and thin line L2 appear at almost the same frequency as expected. In Fig. 3, Stn and Srn are close to each other in frequency. For a specific through-thickness wave mode, the frequency of the n-th order resonance peak is given by v

f(") = n.y~,

(3)

where d is the thickness of the sheet, and v is the phase velocity of the wave in question.

399

Resonance E M A T S Z

surfaCe

I~

~E~ c o i l

Fz

t o p v i ew

Fig. 1. Resonance EMAT. ~, 0 8 ~. _ ,~, , , , , , .> .i-., 0.6

,,I.,

, , , , , , , ' a)

~0.4

" C3

~

<

4

4.5

.> -o ez

S t4

r4

0"I

r

,-

<~-

5.5

5 b

L2 i

I

I

I

I

I

I

I

I

I

6

I

i

i

i

I

i

I

l

!

t

,

6.5

Frequency(MHz) Fig. 2. Resonance spectra of sample B.

~ 0 , 8

l l l i l l l l l l l l l l l l l l l

N 0.6

St3~

0.4

a) r3

0.2 E

< ~0.4

0

l l l l ' l , i l l l i l l l l l l

4

l l l l l l l l l l l l l l l l l l l

St4

E

<

!

4.5

i

i

i

,

,

i

,

r4

(b)

i

,

,

i,

5.5 6 Frequency (MHz) Fig. 3. Resonance spectra of sample F.

400

C.-S. Man et al. 2.07

|

|

|

i

!

|

2.06

2.05

--~x. 2.04

2.03

....<x:x x~

_

I

2.02

2.01 -zl

i

i

-2

I

0

i

2

i

4

6

i

8

10

W4oo x 10 3 Fig. 4. Data points plotting velocity ratio a versus texture coefficient W40o of AA3015 samples. Hence, if we use the m-th order and n-th order resonances for the longitudinal and the shear waves, respectively, the velocity ratio a can be determined from the formula

~ t~ ---- (Y31 +

n

fJ2~

y~)/2 = ~ " (f~7) + fJ;))/2

(4)

In the present work, the velocity ratio a was calculated using L2, St4 and Sr4 for all the samples. The data on a are listed in Table 1. From our experience on similar measurements made with the same R-EMAT system, we estimate that the relative error of the a values reported is of the order of 10 -5 . The type of R-EMATs shown in Fig. 1 provides data also for the evaluation of the texture coefficient W420 through the formula (see, for example, [6]) f~)-

f~)

(f~) + f~2))/2

=

V31- V32

16v~lr 2 = ~cW4~0. (V31 + V32)/2 35#

(5)

For the purpose of using equation (1) for measurement of W400, however, we need not distinguish the Stn and Srn resonance peaks. This allows the use of a round disk type coil, which can simultaneously generate and detect all L, Sr and St waves and is especially suitable for on-line measurements. The data points for all the twenty samples are depicted in Fig. 4. Our experimental findings are generally in agreement with those in the earlier investigation [5], except for a systematic

Resonance EMATS

401

offset between the present R-EMAT and the previous piezoelectric measurements: the equations of the two parallel lines are now = 2.056- 3.88W400, and

a = 2.043- 3.88W400,

(6)

respectively. Again, the experimental data indicate that thermomechanical processing led to small shifts in the quantity ~/(A + 2#)/# or, equivalently, in the isotropic Poisson's ratio. Indeed such shifts provide an additional explanation, other than the difficulty in picking the appropriate single-crystal elastic constants, why trying to use equation (1) and the (constant) Hill averages of A, # and c for the measurement of W400 in this aluminum alloy would be doomed to failure. DISCUSSION Let us begin our discussion with a caution. In the derivation of equation (1), the sample is taken to be homogeneous. Our X-ray measurements, however, showed that there was a rather strong inhomogeneity in texture through the thickness of all the samples. With the measurements we made, we are in no position to assess in what way the throughthickness texture inhomogeneity affected the experimental data, how we could take the texture inhomogeneity into account in our interpretation of the data, and whether taking the arithmetic mean of the W400 values measured at the surface and at the midplane of the samples would be adequate. In what follows we shall simply treat the samples as if they were homogeneous. Within experimental error, we can say that all the data points in the present study lie within the region bounded by the two parallel lines in Fig. 4. Unlike the previous study [5], where the data points scatter largely along two parallel lines, by using suitable thermomechanical processing histories we successfully produced some samples whose data points fall squarely in the interior of the region. As far as monitoring of texture by ultrasound is concerned, this must be taken as a negative finding. A glance at Fig. 4 suggests, for the samples of AA3015 alloy studied, that a value of a generally does not correspond to a unique W400, but to a range equivalent to a margin of error of +0.0017. The location of a data point, however, should be considered in conjunction with the processing history of the sample in question. In the earlier study [5], it was observed that the data points pertaining to sufficiently annealed samples are close to the upper line, while those of sufficiently hot- or cold-worked samples scatter along the lower line. Our present measurement results have corroborated this observation. Moreover, with few exceptions, the effect of thermomechanical processing on the intercept ~/(A + 2#)/# seems to manifest a general trend. If we imagine a straight line which passes through a data point and is parallel to the two drawn lines, annealing tends to move the point down the line and towards the right, i.e., moving it closer to (or at least not further away from) the upper line, whereas cold-rolling has the opposite effect of moving it up the line and towards the left. The exceptions, notably the location of data point I and the shifts from I to J and from I to K, might be caused by inhomogeneities in the sample I. Three specimens were cut from the sample, one for X-ray and ultrasound measurements (the results of which are summarized in the coordinates of data point I in Fig. 4), and two others for making samples J and K, respectively. The anomalies could be understood if the three specimens, which were nominally representatives of the same sample I, in fact had different textures. Of course, what we have offered here is just a speculation. To shed light on what really happened, new specimens have to be prepared

402

C.-S. Man et al.

under the same processing sequences as samples I, J, and K, and the X-ray and ultrasound measurements repeated on the new specimens. Gathering the present measurement results and those in the earlier study [5], we draw the following conclusion: Although a measured value of a generally does not correspond to a unique value of W400 in the aluminum alloys studied, the margin of error in predicting W400 from measurement of a will be substantially reduced if we have prior knowledge on the thermomechanical processing history of the sample in question. Indeed, for the prediction of W400 in AA3015 alloy, simply using a straight line which, in the sense of least squares, best fits the data points in Fig. 4 will carry a margin of error substantially less than =t=0.0017. Thus, in practice, on-line monitoring of texture in aluminum alloys by ultrasonic measurement of W400 remains a promising possibility that merits further investigation. Also, it would be intriguing to find out what changes in the microstructure of aluminum alloys would account for the shifts in isotropic elastic constants induced by thermomechanical processing. ACKN OWLED G EMENTS The work of Man, Fan, and Morris reported herein was supported in part by a DoD EPSCoR grant from AFOSR (No. F49620-98-1-0469). Commonwealth Aluminum Corporation was the industrial partner of the Aluminum Project under this grant. REFERENCES

1. Liu, Y., Cheng, X.-M., Morris, J.G., Man, C.-S. and Huang, M. (1999). In: Proceedings of the Twelfth International Conference on Textures of Materials, Vol. 1, pp. 481-486, Szpunar, J.A. (Ed). NRC Research Press, Ottawa. 2. Clark, A.V., Reno, R.C., Thompson, R.B., Smith, J.F., Blessing, G.V., Fields, R.J., Delsanto, P.P. and Mignogna, R.B. (1988). Ultrasonics 26, 189. 3. Thompson, R.B., Smith, J.F., Lee, S.S. and Johnson, G.C. (1989). Metall. Trans. A 20A, 2431. 4. Anderson, A.J., Thompson, R.B., Bolingbroke, R. and Root, J.H. (1996). Textures and Microstructures 26-27, 39. 5. Man, C.-S., Morris, J.G., Rehbein, D.K., Thompson, R.B., Liu, Y. and Fan, X. (2000). In: Review of Progress in Nondestructive Evaluation, Volume 19, pp. 16451652. American Institute of Physic, Melville, New York. 6. Hirao, M., Hara, N., Fukuoka, H. and Fujisawa, K. (1988). J. Acoust. Soc. Am. 84, 667. 7. Man, C.-S. (1998). Arch. Rational Mech. Anal. 143, 77. 8. Kawashima, K. (1990). J. Acoust. Soc. Am. 87, 681. 9. Kawashima, K. and Wright, O.B. (1992). J. Appl. Phys. 72, 4830.

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

403

CHANGE OF ULTRASONIC ATTENUATION AND M I C R O S T R U C T U R E E V O L U T I O N IN CREPT STAINLESS STEELS

T.OHTANI*, H.OGI**and M. HIRAO**

*Advanced Material Lab., Ebara Research Co. LTD., 4-2-1 Hon-Fujisawa, Fujisawa, kanagawa 251-8502, Japan ** Graduate School, Osaka University 1-3 Machikaneyama-cho, Toyonaka, Osaka, 560-8531, Japan

ABSTRACT We studied the ultrasonic attenuation change during creep tests on austenitic stainless steel (JIS-SUS 304). We applied tensile stresses to the materials at 973K and measured ultrasonic attenuation using electromagnetic acoustic resonance (EMAR). EMAR is a contactless resonant method with an EMAT and is free from extra energy losses in the attenuation measurement, resulting in the measurement of intrinsic attenuation in solids. We used shearwave EMAT, which transmits and receives shear wave in the thickness direction of the plate specimen, to obtain the ultrasonic velocity from the resonant frequency and the attenuation coefficient from the ringdown curve at the resonance. Attenuation exhibits much larger sensitivity to the damage accumulation than the velocity. Approaching to the rapture, it becomes ten times as large as the initial value, which is attributed to microstructural changes, especially, dislocation mobility. This is supported by the TEM observations for dislocation structure. This technique has a potential to assess the damage advance and to predict the creep life of metals.

KEYWORDS Creep damage, stainless steel, ultrasonic spectroscopy, EMAT, ultrasonic attenuation, noncontact evaluation, dislocation.

INTRODUCTION Many of fossil power plants, which were constructed during 1960's and 70's and exceed more than 100,000 working hours, have presently been operating while they have underwent progressive damage, mainly from creep, as the time proceeds[l, 2]. Furthermore, by shifting the base load of power from fossil power plants to nuclear power plants, they are faced to even more severe operating condition such as daily or weekly startup and shutdown in order to correspond to rapid change of the demand. As the consequence of this trend, the material's degradation is being accelerated. Therefore, a nondestructive technique is now highly required

404

T. Ohtani, H. Ogi and M. Hirao

for safely operating plants and predicting the remaining life. It is also important for the technique to be simple and quick operating to cope with the large number of measuring points. This paper describes the ultrasonic attenuation change with EMAR (electromagnetic acoustic resonance) and microstructual evolution during creep tests on austenitic stainless steel (JISSUS 304). EMAR is based on the electromagnetic acoustic resonance (EMAR)[3], which is a combination of the resonant technique and a non-contacting electromagnetic acoustic transducer (EMAT) [3]. Incorporation of EMATs in a resonant measurement contributes to improving the weak coupling efficiency of EMATs to a large extent. The attenuation measurement is inherently free from any energy losses associated with the contact transduction, resulting in the pure attenuation in a metal sample.

MATERIALS AND SHEAR-WAVE EMAT We obtained the creep specimens from hot-rolled commercial plates of an austenitic stainless steels (JIS-SUS304). The gauge section is 5 mm thick, 18 mm wide and 35 mm long (Fig.l). The rolling direction was parallel to the longitudinal direction of the specimen [4]. The mechanical properties are as follows; 0.2% proof stress is 249 MPa, ultimate tensile strength 637 MPa and elongation 72.8%~ We used a shear-wave EMAT of 10 x 10 mm 2 active area, which consists of an elongatedspiral coil and a pair of permanent magnets in opposite directions being normal to the specimen surface (Fig.2). The measurement principle of EMAR appears in [3].

130 oo

~I 15 ,3

Permanent

L,, 15~ C~

_L~__~_~

Eddy ~

~,,_..., / ~

Sample Surface

35

~..Rolling Directiql~

Fig.1. Specimen geometry.

Fig.2. Structure and mechanism of shearwave EMAT.

EXPERIMENT Creep tests were carried out at 973 K in air for various stresses between 100 and 130MPa. Alongside the samples, we attached the unstressed samples to investigate only the effect of thermal history. Interrupted and continuous tests were conducted. In interrupted test, we interrupted creep loading and furnace-cooled the samples. After measuring the attenuation coefficients, we restarted the creep test. We repeated this procedure for every 20, 30 or 100 hours until the rupture. In continuous test, we continued the creep test until the creep strain reached a target

Creep of Stainless Steels

405

value. We thus obtained a series of crept samples with various strains and measured their attenuation coefficients. To further study the microstructure evolution as creep progress, we observed the microstructure of specimens in continuous test with optical microscope (OM), scanning electron microscope (SEM) and transmission electron microscope (TEM). TEM foils were sliced from the midplane at the specimen center using oil-cooling electrodischarge wire to 0.3mm thick, and then carefully polished with SiC paper to 0.2mm thick. Electrolytic polishing in 10% phosphoric acid-ethanol solution induced small center holes. Transmission electron microscopy operating at 250kV was used to obtain the micrographs at thinnest parts around holes. These micrographs are taken in the computer with the scanner for further analysis.

RESULTS

Interrupted test Figure 3 displays the frequency dependence of the attenuation coefficient, c~, in interrupted test (973 K, 120 MPa). The time to rupture, t, was 283 hr. The polarization of the shear wave was parallel to the load. We observe that ot always increases with the resonant frequency in an accelerating manner in the course of creep, ot just before rupture is ten times larger than the initial values. The change only by heating is much less than those of crept samples. Figure 4 shows relationship between c~, relative velocity ratio AV/V0 (V0: initial velocity) of the l lth resonance mode, creep strain, and creep life fraction, t/tr. The time to rupture, t,, was 826 hr. From the start to t/L=0.4, ot remains nearly unchanged. For the period of 0.4 to 0.5, ot shows peak. After that, c~ jumps to the rupture. On the other hand, AV/V0 increases until t/t,=0.2, decreases until t/tr=0.5 and then increases to the rupture. The maximum increment of AV/Vo was around 2 %. As the evolution of ot and AV/V0 are not monotony, it is advisable to use both changes to evaluate the creep damage. Similar results were obtained at other modes and under other stresses.

Time [h]

0.06 7 ~ - + ] - ~ - - - - m - - - - - , - - - ~ * ) ,-iL O.0h ~ <~ Open:Crept sample 9 O 9 ::60h Solid:Reference on ]1-1 9:120h /v T:lS0hl o 0 0.04 [O ,1240hi --

- -

L

0.02 ~

ov~v

()

2(X)

0.08

4(X)

"~'---"w~-"" T

6(X)

A:AV/V, C) :Strain'

0.06

8(X)

[

r

I2

,

1

~,- 1o o~: ~ _

o 0.04

v ~v~ t ~ . _1t, I

va z ~ t ~ . q t ~

"~

1

,z, 2

,

,

....

,_~

3 4 Frequency [MHz]

5

6

Fig.3 Frequency dependence of the shear-wave of attenuation for the polarization in the stress direction (Interrupted test, 120MPa, 973K ).

0.00

5~

-I

0.02

o.oo/i~Itf

- 15

~

) - ' J

0,0

,

0.2

_

_.a__

0.4

,

0.6

_

,

0.8

-2 .0

0

t/t,

Fig.4 Attenuation and velocity evolutions of the 1 lth resonance mode and creep strain. The polarization is parallel to the stress direction (Interrupted test, 100MPa, 973K).

406

T. OhtanL 11. Ogi and M. Hirao

Continuous test

In continuous test, we prepared four and eleven specimens with different creep strains under 120 and 100MPa, respectively. In Figure 5, we plot the creep curve under these two stresses. Each plot shows the measurement of each specimen. The deviation in continuous test is larger than that in interrupted test (Fig.4). If the creep rupture strain and time of each sample are almost the same, they will be useful to predict the creep life. They, however, depends on samples and stresses. Therefore, they could not apply to the remaining life prediction. Then, we estimated the creep rupture time, tr, of each specimen from the creep curve. The estimation was made with the modified 0 function [5] and a rupture parameter, P,~ [6]. We plot the creep strains and t/tr under 120 and 100MPa in Fig.6. Creep strains could arrange on one curve. It shows the good correlation between the creep strain and expended life fraction, t/t~. Then, Figure 7 shows the relationship between the attenuation coefficients of the l lth resonance mode and the estimated life fraction. Like in interrupted test (Fig.5), From the start to t/tr=0.4, ct remains nearly unchanged. For the period of t/tr=0.4 to 0.5, ct shows peak. After that, ct jumps to the rupture. Figure 8 shows relationship between ct, relative velocity ratio, AV/Vo (V0: initial velocity) of the l lth resonance mode, creep strain and creeping time or creep life fraction t/t~ (973K,100MPa). The trend is similar to these in interrupted test (Fig.4). Therefore, by the observing microstructures of specimens in continuous test, we could find what influences the attenuation evolution as creep progresses.

DISCUSSION Possible factors contributing to the change in the attenuation coefficients as creep progresses are scattering from grain boundaries, precipitations and voids, and dislocation damping for a MHz frequency range. Grain scattering

We observed the change of average grain size as the creep deformation with the micrographs by OM. The average grain size increases monotonically as creep progress. The size near the rupture was two times larger than initial one. If the grains are much smaller than ultrasonic wave- length, the scattering from grain boundaries, as, is shown as follows [7]. a~ = Sd3 f 4

(1)

where S is the scattering factor, d the average grain size, and f frequency. We calculated cts for the change of grain size, where S=2.25x10 -~~~sec3/~m 3 [8] which is the value of carbon steels, whose elastic moduli are similar to those of stainless steel. The change of ots is calculated to be 2.72x10 3 9sec -~(f=3.5MHz) and much smaller than the measurement. Moreover, monotonous increase of grain size can not explain the change of attenuation coefficients like in Fig. 8.

Creep of Stainless Steels

~5~,

,

,

/

,

407

~..%o-~.'

,

./

I

12'OMPa ' lOOMPa]

9

10

"'2 5 O

"r, ~ j ] ..~ , ~ , ~

0

2

;o

,

4

;o

,

600 '

,

800 '

,

0

1 100

0.0

0.2

0.4

I

0.6

,

I

0.8

1.0

t/t

Time [h I

Fig.6 Creep strain versus the estimated life fraction (Continuous test, 120 and 100MPa, 973K)

Fig.5 Creep strain versus time (Continuous test, 120 and 100MPa, 973K).

'

0.10

i

'

i

'

i

'

i

IOOMPa ..............................i 13:5th mode i A :8thmode O "1lth mode ....... " . . . . . . . . . . I

0.08 ,.--, - ~ 0.06 '-~ 0.04

'

o // l

~.... /

0.02

.';Q

-

[] ..... ...

0.2

/

[[ ~'i .;~

t~

0.00 ~ 0.0

_

-

0.4

0.6

0.8

1.0

t/t

Fig.7 Relationship between the attenuation coefficients of the 5th, 8th and l lth resonance modes and the estimated life fraction (Continuous test, 100MPa, 973K).

0.10 0.08 ,',

0.06

~"

0.04

4

I.i/

15

i00 , a

I

I0: Strain'

/

V

2 10

o

~ 5~

0.02 0.00 0.0

0.2

0.4

t/t

0.6

0.8

-4 1.0

0

Fig.8 Attenuation and velocity evolutions of the 1 l th resonance mode and creep strain versus the estimated life fraction (Continuous test, 100MPa, 973K).

408

T. Ohtani, H. Ogi and M. Hirao

Scattering from precipitations and voids

From SEM micrographs, precipitations and voids are observed from t/tr=0.25 and 0.4, respectively. The number densities increase slightly to the rupture. Their average sizes are around 2~tm. Because their sizes (>2Ftm) were much smaller than the probing wavelength (>0.5mm), the scattering in the Rayleigh region does occur [7]; a .: nyl2 - Ca 2 (ka 4 ),(ka << 1)

(2)

where n denotes the volume density of the scatter, ~, is the scattering cross section, k the wave number and a the radius of the scatterer. C is constant depending on the wave mode, n, and the scatterer's type (i.e. inclusion or cavity). From SEM micrographs, we assume that the attenuation increases in a monotonous way. From Eq. (2), we calculated ~ts to be less than 1% of measurement. Therefore, these effects cause only negligible changes in c~ and the dislocation damping can solely explain the observed acoustic response. Dislocation damping

Dislocations vibrate responding to the ultrasonic stress with a phase lag because of viscosity and dissipate of the energy. This anelastic mechanism also lowers the ultrasonic velocities. The dislocation lines are pinned by point defects, precipitations and other dislocations. The pinning points act as nodes of the vibrations. From Granato and Lticke' s string model for the dislocation damping [9], we have in the lower frequency range a

=

A1AL4f 4 and AV/Vo = -AzAL 2

(3)

where A1, A2 are positive constants depending on in shear modulus, Poisson's ratio, specific damping constant, and Burger's vector. A is the dislocation density and L the dislocation length. According to this model, a is proportional to A and the fourth power of L of the effective dislocations, which are mobile and can vibrate with the ultrasonic stress. Note that not all the dislocations interact with ultrasonics. The frequency dependence of the attenuation coefficient in Fig.7 and the velocity drops in Fig.8 support this theoretical model. To a further study, we observed the evolution of dislocation structure by TEM. Figure 9 shows the TEM micrographs as creep progress. Figure 9(a) shows the dislocation structure before the creep, t/tr=0 and Fig. 9(b)~(d), at t/tr=0.38, 0.67 and 0.98. In the structure before the creep, dislocations distributed nearly uniform, although tangled dislocations and twins appeared. Figure 9(b) shows the microstucture at t/tr=0.38, when ct shows a peak (Fig.8). Many tangled dislocations were observed. The formation of sub-boundary started from this stage. In the subgrain, the density of pinned dislocations become high. In Fig. 9(c), ordered subgrains were observed. The dislocation density within subgrains reduced. In Fig. 9(d), which is near the rupture, the subgrains become fine and the density within subgrains increased again. From these TEM micrographs and some quantitative results [10-12], we assumed that the change of ot is related to the change of dislocation structure, especially, free dislocations in subgrains because the dislocations bound in subboundaries could not vibrate responding to the ultrasonic stress. The whole phenomena can be explained as follows. The dislocation density increased from the beginning of the creep as the creep deformation advanced. After this, dislocations were tangled and the formation of subgrains started. By the continuous

Creep of Stainless Steels

409

deformation, the majority of dislocations generated were spent in the formation of subboundaries. As a result, many subgrains developed and the dislocation density in subgrains reduced. Furthermore, by the acceleration of the creep deformation, dislocation multiplication overcame the consumption of the dislocation into the sub-boundaries and the free dislocations in sub-grains increased. Finally, the crack initiated at multiplied dislocations on subgrains, grain boundaries, precipitations and the rupture started.

Fig.9 TEM micrographs of crept specimens at t/tr=0, 0.38, 0.67, 0.98 (Continuous test, 100MPa, 973K).

CONCLLUSION Creep damage in austenitic stainless steel (SUS304) at 973K in air was evaluated through ultrasonic attenuation measured with the EMAR method. The attenuation extremely increased as the creep damage progresses, amounting to ten time larger than initial values, which is interpreted as resulting from microstructural changes, especially, dislocation mobility. This is supported by TEM observations. This technique has a potential to assess the damage. There is clear relationship between the attenuation coefficient and expended life fraction. The further work should include the quantitative investigation between the attenuation evolution and the dislocation structure.

410

T. Ohtani, H. Ogi and M. Hirao

REFERENCES

,

8. 9. 10. 11. 12.

Drinon, D.S., Liaw, P.K.and Rishel, M.K., (1992).ASME PVP, 239/MPC-33, 5. Viswanathan, R. (1989). Damage Mechanism and Life Assessment of High Temperature Components. ASME international. Hirao, M., Ogi, H. (1989). Ultrasonics 35, 413. Ohtani, T., Ogi, H. and Hirao, M., (1999). Rev. Prog. in QNDE, 18, 1847. Maruyama, K., Harada, C. and Oikawa, H., (1985). J.Soc. Mater., JPN., 34, 1289. Maruyama, K. and Oikawa, H., (1990). Trans. ASME.J. Pressure VesseLTechnol., 109, 142. Goebbels, K., (1980). Res. Tech. in NDT, 4, 87. Ogi, H., Hirao, M. and Honda, T., (1995). J. Acoust. Soc. Am., 98, 458. Granato, A. and Locke, K., (1956). J. Appl. Phys., 27,789. Barrett, C.R., Nix, W.D. and Sherby, O.D. (1966). Trans ASM., 59, 3. Orlova, A. and Cadek, J., (1973). Phil Mag., 28, 891. Hasegawa, T., Ikeuchi,Y. and Karashima,S., (1972). Metal Sci. J., 6, 78.

Nondestructive Characterizationof MaterialsX Green et al. (Eds) CO2001 ElsevierScienceLtd. All rights reserved.

411

NONDESTRUCTIVE EVALUATION OF A GREY CAST IRON USING THE ULTRASONIC MEASUREMENT H. MOCHIZUKI1, Z. YAJIMA2, Y. KISHI2 T. YOSHIDA3 and K. SHIMIZU1 1 : Sugiyama Co., Ltd. 2240 Iitomi, Nakatomi-machi, Minamikoma-gun, Yamanashi 409-3423, Japan 2 :AMS R&D Center, Kanazawa Institute of Technology 3-1 Yatsukaho, Matto, Ishikawa 924-0838, Japan 3 :Casting Technology Research Laboratory, Hitachi Metals Ltd. 11 Kinugaoka, Moka, Tochigi 321-43, Japan ABSTRACT Measurement of the X-ray diffraction profile and the ultrasonic velocity were carried out for annealed grey cast irons. The test material was prepared by high-frequency induction-melting method. The specimens machined from the ingot were annealed at 823 K, 873 K and 923 K for 1.8, 3.6, 7.2 and 36 ks, and were lightly electro-polished in order to remove some kind of thin oxide surface layer. The diffraction profile of ot-Fe 211 was observed by Cr-Kc~ radiations. Full width of half maximum intensity and integral width linearly decreased with increasing annealing temperature and time. Deference in the ultrasonic velocity, AV, which was determined from the value subtracting the ultrasonic velocity in the annealed alloy from that in the as-cast alloy, appeared to depend on the optical microstructures of the annealed alloys. For the alloy annealed at 823 K for 1.8 , 3.6, 7.2 and 36 ks alloy and the alloy annealed at 873 K for 1.8, 3.6 and 7.2 ks, which optical microstructures are equivalent to those of the as-cast alloy, A V slightly increase with increasing annealing time. While for the other alloys, in which secondary graphite and ferrite phases were observed, the former value decrease with increasing the later one.

KEYWORDS Grey cast iron, Ultrasonic velocity, X-ray diffraction, FWHM,Integral width, Lattice distortion

INTRODUCTION Small mass-produce parts made from the grey cast iron may be annealed because of the prevention of secular change and the improvement of machinability. However, the appropriate annealing condition is not made clear science optical microstructure of the annealed alloy is equivalent to that of the as-cast alloy. Although the ultrasonic velocity in the material is generally known to be affected by microstructure, residual stress, residual strain and other metallographical factors [ 1], there is no mention about a correlation between the ultrasonic velocity and the annealing condition for the grey cast iron. It is important to know

H. Mochizuki et al.

412

whether the measurement of the ultrasonic velocity have an effect on the estimation for the annealed cast iron, because the ultrasonic velocity in the materials can be easily measured. In this study, the ultrasonic velocity, X-ray profile, and mechanical properties of annealed grey cast irons are investigated, and then influence of annealing condition on the ultrasonic velocity are discussed on the basis of the experimental results.

EXPERIMENTAL PROCEDURE The grey cast iron used in the present work was made from steel scrap and cast iron returns, which were melted in a high-frequency induction smelting furnace. Following some inoculation treatments, the melt was casted in a sand-mold 25 mm diameter and 250 mm length. Its chemical composition was as follows: 3.36C, 2.27Si, 0.49Mn, 0.041P and 0.026S ( mass. % ). The as-cast alloy consisted of graphite and pearlite phases as show in Fig. 1. Table 1 shows mechanical properties and X-ray parameters of the as-cast alloy.

Fig. 1.

Optical microstructure of the as-cast alloy.

Table 1. Mechanical properties and X-ray parameters of the as-cast alloy. HRB 0"8, MPa Bo, deg. flo, deg. 93 274 1.1293 2.0820

HRB : Rockwell hardness o'8 : Tensile strength Bo: Full width of half maximum intensity of ~-Fe 211 diffraction /3o: Integral width ( ratio of integral intensity and peak intensity ) of a-Fe 211 diffraction

Three kinds of test specimens, which were the plate-shape specimen ( 50 mm ( length ) x 10 mm ( width ) x 5 mm ( thickness ) ) for X-ray measurement, the stick-shape one (10 mm ( active gauge length ) x 14 mm (diameter) for mechanical tests, and the rod-shape one ( 22 mm ( diameter ) x 50 mm ( length ) ) for the ultrasonic measurements, were machined from the

NDE of Grey Cast Iron

413

ingot. These specimens were then annealed at 823 K, 873 K and 923 K for 1.8, 3.6, 7.2, 10.8 and 36 ks ( which are named 823-T, 873-T and 923-T alloys, T : annealing time ), and were lightly electrical-polished in order to remove some kind of thin oxide surface layer. The ultrasonic tests were carried out by a direct contact method with a probe of longitudinal wave frequency of 5 MHz. Deference in the ultrasonic velocity, A V, which was determined from the value subtracting the ultrasonic velocity of the annealed alloy from that of the as-cast alloy, was estimated. Table 2 shows the condition of X-ray measurement. The X-ray diffraction profile of cz-Fe 211 was observed by Cr-Kcz radiations. Full width of half maximum intensity ( FWHM ), B, which is measured the range at a half peak value of the KCZl and Kct2 doublet, and integral width, fl, were measured in order to estimate the lattice distortion of the annealed alloys[2,3]. The value offlwas determined by following equation [3];

fl= IN

(1)

where Ip and IN are the maximum intensity and the integral intensity of diffraction profile, respectively.

Table 2. X-ray diffraction condition Parallel beam method Method

Fig. 2

Characteristic X-ray

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Optical microstructures of the annealed alloys.

414

Fig. 2

H. Mochizuki et al.

Continued.

RESULTS AND DISCUSSION Typical microstructures of the annealed alloys are shown in Fig. 2. Optical microstructures of the 823-T, 873-1.8, 873-3.6, and 873-7.2 alloys are equivalent to that of the as-cast alloy. For the 873-10.8, 873-36 and 923-T alloys, secondary graphite and ferrite grains were observed, and their volume fractions increased with increasing the annealing temperature and time. Tensile strength, 0"8, and Rockwell hardness value, HRB, of the annealed alloys were measured, and these results are shown in Fig. 3. o8 and HRB of the 823-T alloys slightly increase with increasing the annealing time, T, while the o'8 and HRB vs. T curves of the 873-T 100

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NDE of Grey Cast lron

415

and 923-T alloys differ from that of the 823-T alloys. For the 873-T alloys, o'8 and HRB slightly increase, and then rapidly decrease with increasing T. While for the 923-T alloys, O-B and HRB decrease with increasing T, and then these values converge to constants. Hence, the o-B and HRB vs. T curves of the annealed alloys are considered to reflect the optical microstructures of the alloys. Ratio of FWHM, B / Bo, and ratio of integral width, fl / flo ( Bo and flo : values from the as-cast alloy ), are plotted against annealing time, T, as shown in Fig. 4. B / Bo and fl / flo

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416

H. Mochizuki et al.

linearly decrease with increasing T, therefore, the lattice distortions in the alloys seem to be released by annealing. Deference in the ultrasonic velocity, AV, is plotted against annealing time T in Fig. 5. The A V value appears to depend on the optical microstructures of the annealed alloys. For the 823-T, 873-1.8, 873-3.6 and 873-7.2 alloys, which optical microstructures are equivalent to those of the as-cast alloy, AV slightly increases with increasing T. While for the other alloys, in which secondary graphite and ferrite phases were observed, the A V value decrease with increasing T. In conclusion, measurement of the ultrasonic velocity in the grey cat iron was found to detect the precipitation of secondary graphite and ferrite phase. Many investigators have reported on ultrasonic velocity in grey cast irons and nodular graphite cast irons, in order to clarify influences on microstructures of the cast irons [4, 5]. According to their reports, ultrasonic velocity in cast irons is influenced by shape of graphite, volume fraction of graphite and pore and so on, but there are no mention about the annealing condition for cast irons. Relationship between the ultrasonic velocity and the annealing condition is still under investigating.

CONCLUSION X-ray diffraction analysis and measurement of the ultrasonic velocity were carried out for the annealed grey cast irons. Ratio of FWHM and ratio of integral width, which were determined from the ot-Fe 211 X-ray diffraction profile, linearly decreased with increasing annealing temperature and time, therefore, the lattice distortions in the alloys seem to be released by the annealing-treatment. Deference in the ultrasonic velocity AV depended on the optical microstructures of the annealed alloys. For the 823-T, 873-1.8, 873-3.6 and 873-7.2 alloys, which the optical microstructures are equivalent to that of the as-cast alloy, AV slightly increase with increasing T. While for the other alloys, in which secondary graphite and ferrite phases were observed, A V decreased with increasing T. In conclusion, measurement of the ultrasonic velocity in the grey cat iron founds to detect the precipitation of secondary graphite and ferrite phase.

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. The Japanese Society for Non-Destructive Inspection (1993). In: Elements of Nondestructive Inspection, The Japanese society for Non-Destructive Inspection ( Ed ), The Japanese society for Non-Destructive Inspection, Tokyo, Japan. 2. Cullity, B. D. (1989). In: Elements of X-Ray Diffraction -Second Edition-, Agne Co. LTD, Tokyo, Japan. 3. Rigaku Corporation (1991). In: X-Ray Diffraction Handbook-Forth Edition-, Rigaku Corporation ( Ed ), Rigaku Corporation, Tokyo, Japan. 4. T. Abe, private latter. 5. Kusaka Rear-Metal Laboratory, Co. LTD., private letter.

AUTHOR INDEX

Ab6, H., 129 Amde, A. M., 167 Amelina, E., 279 Aoyama, H., 45, 75 Asai, M., 99 Balogh, J., 175 Biwa, S., 223 Blouin, A., 27 Cho, H., 135,311,315 Choquet, M., 27 Deb, S. K., 175 DeLeon, N. L., 85 Djordjevic, B. B., 279

Enoki, M., 193

Kamoi, A., 371 Karasawa, H., 121 Kato, H., 185, 327 Katoh, M., 387 Kawashima, K., 199, 395 Kim, K. C., 239 Kishi, N., 231 Kishi, T., 53, 193 Kishi, Y., 99, 411 Kishibe, D., 231 Kishimoto K., 143 Kitahara, M., 231 Kobayashi, K. F., 61 Koktavy, B., 273 Koktavy, P., 273 Komai, T., 185 Komura, I., 198 Kondo, K., 371 Konno, Y., 151 Koo, J. H., 53 Koyama, K., 113, 121 Kubo, T., 198 Kume, R., 341 Kurita, M., 105

Fan, X., 395 Fujita, H., 199 Green, R. E., Jr., 3, 91 Green, W. H., 91 Hirano, K., 349 Hirao, M., 69, 357, 379, 403 Hiratsuka, H., 297 Hirose, A., 61 Horii, Y., 341 Hoshikawa, H., 113, 121 Ihara, I., 205 fida, T., 257 Ikeda, H., 297, 387 Imade, M., 289 Inoue H., 143 Isobe, Y., 333 Jang, B. K., 53 Jhang, K. Y., 239 Jiang, G., 185 Ju, Y., 129

Ledbetter, H., 19, 69 Lee, J., 327 Lee, K. W., 349 Lemor, R., 175 l.~vesque, D., 27 Lokajicek, V., 273 Livingston, R. A., 167

Made, A.M., 167 Man, C.-S., 395 Manghnani, M. H., 175 Matsuda, Y., 297 Matsumoto, S., 311 Matsumoto, Y., 387 Mihara, T., 159 Minami, Y., 379 Minkov, D., 327 Mochizuki, M., 411 Monchalin, J.-P., 27 Moreau, A., 27 Morris, J. G., 395 Mukai, N., 305 Musashi, T., 159

418 Nagai, S., 297 Nagata, Y., 151 Naito, S., 151 Nakano, H., 297, 305 Nakano, S., 311 Nakazawa, T., 315 Nishio, K., 387 Ochiai, M., 305 Ogi, H., 69, 379, 403 Ogiso, H., 311 Ohno, N., 223 Ohtani, T., 403 Ohtsu, M., 257, 265 Okabe, Y., 247 Okamoto, Y., 297, 371 Outlaw, J. F., 85 Pavelka, J., 273 Prasad, M., 175 Rouch, L. L., 279 Saito, T., 239 Saka, M., 129 Salafia, D., 85, 365 Sano, Y., 305 Sanpei, S., 129 Sato, H., 311 Sawa, T., 205 Sedlakova, V., 273 Shchukin, E., 279 Shibuya T., 143 Shigeishi, M., 265 Shima, T., 199 Shimizu, K., 99, 411 Shimoike, G., 69 Shiwa, M., 341 Shoji, T., 327 Sikura, J., 273 Suzuki, T., 349

Author Index

Takahashi, T., 231 Takashima, K., 69 Takatsubo, J., 289 Takeda, N., 11, 45, 75, 247 Tanaka, K., 45, 75, 205 Taniyama, N., 113 Toyama, N., 53 Toyooka, S., 333 Tsuji, T., 213 Tsujikura, S., 297 Vavilov, V. P., 371 Vidensky, I., 279 Wang, Y., 175 Watanabe, M., 193 Watanabe, T., 315 Watanabe, Y., 223 Winter, J. M., Jr., 91 Yajima, Z., 99, 411 Yamada, H., 151 Yamada, K., 333 Yamaguche, A., 341 Yamaguchi, K., 333 Yamaguchi, T., 387 Yamamoto, M., 297 Yamamoto, S., 289 Yamanaka K., 135, 159, 213, 311, 315 Yamasaki, T., 357 Yamawaki, H., 239 Yoneyama, H., 341 Yoshida, T., 411 Yoshida, Y., 185 Zinin, P.V., 175

419

KEYWORD INDEX

Absorption, 19,223 AC magnetic method, 341 Acoustic emission, 53, 61, 75, 193, 257, 265, 273 Acoustic microscopy, 175 Acoustic reflection coefficient, 247 Aerospace structures, 11 Air-coupled transducers, 3 ALFRP, 75 Aluminium, 91 Aluminium alloy castings, 185 Aluminium alloys, 395 Attenuation, 19, 231 Austenitic stainless steel, 349 Autoradiography, 167 Axial-shear-wave EMAT, 379

Backscattering wave, 185 Bending stiffness, 45 Bridges, 167 Buried defect, 371

Cement-based material, 231 Ceramic coating, 193 CFRP, 53 Civil infrastructures, 11 Coating, 199 Coatings, 143 Coaxial line sensor, 129 Composite coating, 205 Composite material, 45 Composites, 11, 19,27 Concrete, 167 Concrete-core test, 257 Confocal Fabry-Perot interferometer, 305 Contact stiffness, 213 Continuous fiber-reinforced ceramic composites (CFCC), 175 Corrosion detection, 27 Crack, 273,289 Crack kinematics, 265 Cracks, 19 Creep damage, 403 Crystallographic texture, 395

Damage estimation, 257 Damage mechanics, 257, 265 Debonding, 61 Debye characteristic temperature, 19 Deconvolution, 159 Deep reactive ion etching, 311 Defect, 289 Delamination, 45 Differential surface tangential probe, 113 Dislocation, 213,403 Dispersion, 19

Eddy current testing, 113, 121 Elastic constant, 205 Elastic constants, 19 Elastic properties, 247 Elastic stiffnesses, 69 Electrically conductive polymers, 11 Electromagnetic acoustic transducer, 69 Electromagnetic acoustic transducers (EMAT), 3, 403 Electromagnetic emission, 273 Encapsulant resin, 129 Ettringite, 167 Experimental stress analysis, 105 Extensional mode, 53

Faraday effect, 327 Fatigue crack, 349 FBG sensor, 75 FDM, 239 Fiber/matrix interfacial reaction zone, 61 Flaw detection, 113 Fractography, 61 Fracture behavior, 61 FWHM, 411

Grain size, 27, 185 Granite, 273 Grey cast iron, 411

Hall sensors, 333 Hardness, 19 Heat balance equation, 371 HOPG, 213 Hydration of cement, 279

420 IC package, 129 Impact-echo, 167 Induction hardening, 379 In-situ damage monitoring, 349 Integral width, 411 Interferometers, 3 Interlaminar-toughened CFRP, 247 Internal damage, 75 Internal friction, 19 Inverse analysis, 193, 205

Lamb wave, 53, 143, 151,311 Laminated construction, 45 Laser interferometer, 193 Laser-based ultrasonics, 27 Laser-based ultrasound, 27 Laser-generated ultrasound, 311 Lasers, 3, 289 Laser ultrasonics, 27, 305 Laser-ultrasound, 27 Lattice distortion, 411 Leakage flux, 333 Leaky surface wave, 199 Lift-off noise, 121 Linear perturbation scheme, 379 Low density core, 91

Magnetic domain, 327 Magneto-optical method, 327 Magneto-resistance, 333 Magnetostriction, 357 Martensitic transformation, 99 Material degradation, 239 Materials characterization, 27 Measurable nonlinear parameter, 239 Measurement-theory, 19 MEMS, 311 Metal matrix composites, 61, 69 Microscopic damage, 11 Microstructure, 27 Microwave, 129 Moisture, 129 Moment tensor, 265 Multi quantum well, 297

Nano-indentation, 205 Nano-scale, 213 NDE, 333 Non-contact, 3 Non-contact evaluation, 403 Non-contact measurement, 69 Nondestructive damage analysis, 349

Keyword index Non-destructive evaluation, 305, 311 Nondestructive inspection, 289, 327 Nondestructive testing, 105, 185, 273 Nonlinear acoustic effect, 239 Non-linear processing, 151 Normalized 3rd harmonic component intensity, 341

On-line monitoring, 395 Optical fiber, 53, 305 Optical fiber sensor, 11 Optimization, 143

Paper, 27 Penny shaped cracks, 231 Phase velocity scanning method, 311 Photoelasticity, 159 Photorefractive effect, 297 Plastic deformation, 357 Plates, 143 Polymer composites, 27 Polymer-based composite, 223 Pore distribution, 91 Pore size, 91 Pore/metal composition, 91 Potassium, 167 Precise measurement, 135 Process control, 3 Process monitoring, 27 PVS method, 135 PWHT temperature, 341

Quantitative evaluation, 247 Quartz, 135

Raman spectroscopy, 175 Rate process, 257 Rayleigh-like waves, 143 Reactor internal inspection, 305 Reinforced plastics, 45 Remelted zone, 185 Remote sensing, 327 Remote sensing NDT device, 371 Residual stress, 279, 357 Residual stress measurement, 105 Resonance, 69 Resonance EMATS, 395 Resonance frequency, 213 Resonant frequency, 379 Reversible displacement, 213

Keyword index Scattering, 19,223 Scattering cross-section, 231 Shape-memory alloy, 11, 99 SiC/Ti-6A1-4V composite, 61 Signal phase, 121 Signal processing, 151 SMA, 53 Sound velocities, 19 Source location, 53 SQUID, 349 Stainless steel, 403 Steel cylinder, 379 Steels, 19, 27, 91,105, 357 Stratified sampling, 167 Stress corrosion cracking, 305 Stress measurement, 357 Structural health monitoring, 11 Subsurface defects, 213, 311 Surface acoustic waves, 135,205,305,311 Surface flaw depth, 121 Surface probe, 121 SUS304, 333 Synthetic Aperture Focusing Technique (SAFT), 27 Synthetic aperture, 27

Tensile deformation, 99 Tensile test, 61 Texture, 19 Theory of elasticity, 105 Thermal image technique, 371 Thickness, 185 Transmitted light technique, 45, 75 Two-wave mixing, 297 Ultra lightweight metal foams, 91 Ultrasonic atomic force microscopy, 213 Ultrasonic attenuation, 199, 223,279, 403 Ultrasonic inspection, 45 Ultrasonic micro-spectrometer, 247 Ultrasonic monitoring, 279 Ultrasonic nondestructive testing, 27 Ultrasonics, 3, 27, 143 Ultrasonic spectroscopy, 403 Ultrasonic testing, 151, 185,305 Ultrasonic velocity, 199, 411 Ultrasonic visualization, 159 Ultrasonic wave, 289 Uniform eddy current probe, 113

Velocity anisotropy, 135 Viscoelasticity, 223 Visible and infrared method, 371 Visualization, 289 Voids, 19

Wave dispersion, 143 Waveform analysis, 159 Wavefront compensation, 297 Wavelet transform, 143 Waves, 19 Weld inspection, 113

X-ray computed tomography, 91 X-ray diffraction, 105,279, 411 X-ray diffraction profile, 99

A533B, 333

t~' - martensite, 349

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