materials science research international, vol.8, no.3 pp

Materials Science Research International, Vol.8, No.3 pp. 141-150 (2002)

Review paper

Progress in the Instrumented Charpy Impact Test

Toshiro KOBAYASHI**Department of Production Systems Engineering, Toyohashi University of Technology,

Tempaku-cho, Toyohashi 441-8580, Japan

Abstract: It has passed one century since French G. Charpy proposed the Charpy impact test in 1901. Instrumented impact test recording load history during fracture has been attempted since 1920's. It has become possible to obtain various information from this method. However, succeeding development of fracture mechanics has made it also as a qualitative screening test. To obtain quantitative fracture toughness parameters from this method, therefore, has been desired eagerly. The author developed successfully a computer aided dynamic fracture toughness evaluation system, called CAI system already in 1980's. This paper reviews development and progress in the Charpy test, especially focusing on problems of evaluation of dynamic fracture toughness by the Instrumented impact test and on accuracy of measurement in the test.

Key words: Charpy impact test, Instrumented impact test, Dynamic fracture toughness, Toughness, CAI system, Fracture energy

1. INTRODUCTION

In 1901, the Charpy impact test was presented at the

Budapest Congress of International Association for Test-

ing of Materials by G. Charpy of France [1]. Its use has

since spread, and 100 years have now passed. Although

the strength of materials had been evaluated under static

loading conditions, the understanding of the behavior

under impact loading has become a major problem. Frac-

ture energy of a specimen can be measured easily using a

pendulum-type machine. In fact, Russel of the U.S.A.

had already submitted his paper [2] on this idea to the

American Society of Civil Engineers (ASCE) in 1897.

Fremont of France had also presented a method for

measuring energy after the specimen breaks, using

compression of a spring [1]. Because of high activity in

this field in France in that day and Charpy's high position

(he later became the chairman of standardization), the

pendulum-type impact method spread widely and be-came known as the Charpy test. The machine which

Charpy used first had a mass of 50kgf and arm length of

4m (capacity: 250kgfm). The work in the rupture of the

specimen is called •gresilience•h, and it was used as a

measure of toughness.

Although the absorbed energy of fracture can be sim-

ply obtained using the pendulum-type machine, informa-

tion on the fracture process cannot be obtained. Attempts

at recording the load-time or load-deflection curve dur-

ing fracture have been carried out since the 1920s. Kor-

ber and Storp [3] used an optical method, and Yamada

[4] improved and further developed that method. Subse-

quently, the piezoelectricity of rock crystal was applied

to this method by Watanabe [5], and by Tanaka and

Umekawa utilizing Rochelle Salt [6]. Sakui et al. [7]

used a resistance wire strain gage. In recent years, the

method of using a semiconductor strain gage has become

the mainstream. For the recording of displacement, vari-

ous methods of using photoelectric tubes, phototransis-

tors, film potentiometers attached to the hammer rotating

axis, magnetic sensors, and laser beams have been pre-

sented [8]. However, the International Institute of Weld-

ing (IIW) had previously concluded that, even by In-

strumented methods, no quantitative information on frac-

tures could be obtained [9].

However, the Instrumented impact test is presently

aided by the computer and can be used to analyze the

load-deflection curve of a precracked Charpy specimen

and can yield a valid dynamic fracture toughness. These

developments and progress in the 20th century will be

reviewed based mainly on the author's results.

2. MECHANICS OF THE CHARPY IMPACT TEST

Many empirical formulas on the correlation between

the Charpy value and fracture toughness KId or KIC have

been reported [10]. For example, it is basically estimated

from the following relation in the transition region.

KIC,KId=A(CVN)n. (1)

Considering a safety factor, there is another proposal as

follows [10] (CVN is absorbed energy of the standard

V-notch specimen).

KIC=8.47(CVN)0.63(MPa•Em1/2,J), (2)

KId=22.5(CVN)0.17(MPa•Em1/2,J). (3)

There are also many examples of a correlation between

the plastic zone parameter and absorbed energy normal-

ized with yield strength ƒÐY. The following equation

given by Rolfe and Novak for the ductile range is well

known [11,12].

(KIC/ƒÐY)2=0.645{(CVN/ƒÐY)-0.0098}(m,J/MPa). (4)

Figure 1 shows the correlation between the Rolfe-Novak

empirical formula and the measured values [11]. A good

correlation is not observed. The fact that the obtained KIC

Received November 7, 2001

Accepted April 19, 2002

Original paper in Japanese was published in Journal of the Society of Materials Science, Japan, Vol.51, No.7 (2002) pp. 771-779.

141

Toshiro KOBAYASHI

Fig.1. Comparison between the Rolfe-Novak's correlation and measured values of KId, KIC and Cv.

in this figure is not necessarily valid has also been

pointed out. This equation yields ƒÐY in the range of 760-1700MPa and CVN is within 31•`121J. The range is

exceeded slightly by the measured CVN values. Ulti-

mately, there is a limit to the empirical correlation.

Therefore, the development of a method of directly

measuring dynamic fracture toughness parameters has

been expected.

At present, as a specimen geometry, thickness and

width of 10mm, length of 55mm, and V-notch depth of

2mm (tip radius ƒÏ=0.25mm) are recommended. How-

ever the physical reasons behind such geometry are am-

biguous. For example, ƒÏ=0.25mm has merely been

recommended for the empirical reason of the least scatter

of experimental impact values. By applying slip-line-

field analysis to this method, Green and Hundy [13]

analyzed the deformation behavior. The plane strain con-

dition of rigid plasticity was assumed; an example, of the

slip line field for general yield is as shown in Fig. 2.

Within ABCD is the rigid region, and the specimen ro-

tates along circular arcs AB and AD by shear. When a

local yield occurs at the notch tip before the general yield

(point A), longitudinal stress distribution ƒÐy in the plastic region (r•…RƒÀ) is as shown in Fig. 3.

ƒÐy=2k[1+ln(1+r/ƒÏ)]. (5)

Here, yield stress ƒÐY=2k [14].

The plastic stress concentration factor KƒÐ(p) is given

as ƒÐymax/ƒÐY=1+ln(1+R/ƒÏ). The maximum value of

KƒÐ(p), i.e., KƒÐ(P)max, becomes as follows (ƒÖ is the opening

angle of the notch):

(ƒÐymax/ƒÐY)max=1+ƒÎ/2-ƒÖ/2=1/ƒÀ=KƒÐ(p)max. (6)

KƒÐ(p)max becomes 2.18 in the 45•‹ V-notch and 2.57 in the

precracked specimen, assuming Tresca's criterion [15].

The bending yield stress=3PY•El/2bh2, b: specimen

Fig. 2. Slip line field under general yielding of Charpy V notch specimen (Green and Hundy) [13].

Fig. 3. Changes of ƒÐy and ƒÐx with notch front distance r

at certain plastic zone size R [15]. (a) R<RƒÀ, (b) R>RƒÀ.

ƒÐymax attains its maximum value (1/ƒÀ) ƒÐY firstly at RƒÀ. The

scale is not same in (a) and (b).

thickness, h: ligament depth (=W-a; W is specimen

width; a is notch length), and l: span distance (40mm)

obtained from PY at the temperature where yield-point

brittle fracture occurs. Then, bending yield stress is con-

verted (tensile stress is generally said to become about

1/2 of bending stress)* to tensile stress, and the critical

cleavage stress of this material is estimated as 2.18•~ the

tensile stress (precisely, the yield stress should be esti-

mated as 0.8 times the value of PY at the point of brittle

fracture). This is effective at least for obtaining one ma-

terial constant. Wells [17] analyzed the COD in the

relation of displacement x to unstable fracture under a

*Otherwise, dynamic tensile yield stress ƒÐyd can be estimated from bending yield load PY in the instrumented test,

based on the following equations presented by Server [16]: ƒÐyd=2.99PYW/B(W-a)2(V-notch); ƒÐyd=2.85PYW/B(W-a)2

(precracked).

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Fig. 4. Representation of the notch profile and obtained

relationship between COD and load point deflection x.

general yielding condition and presented the following equation:

COD=0.8hx/l. (7)

Kobayashi showed the relationship between load point

displacement x and bending angle ƒÆ in the Charpy test

under large deformation where elastic deformation is

disregarded, as in Fig. 4 [11]. The following relation is

established:

x=l/2sinƒÆ/2, (8)

COD=2r(W-a)sinƒÆ/2, (9)

r is the rotation factor and r=0.42 [11]. From these, the

following equation is derived:

COD=2r(W-a)2x/l. (10)

Figure 4 shows an example of the experimental results,

where both COD and x are off-load values. The deviation

from the straight line is probably due to crack extension.

It has been proven that COD may be estimated from dis-

placement x [11]. On the other hand, Rintama and Zim-mermann developed the method of using a laser beam to

measure the COD [18].

It has been considered important that the characteris-

tic distance l0* which Ritchie and Thompson proposed

correlates to the macroscopic fracture behavior and mi-

crostructure [19]. Consequently, it is generally said that

the apparent critical strain energy release rate of the

specimen with a notch tip radius ƒÏ shows the following

relation:

GƒÏ/GC=1/2+ƒÏ/8l0. (11)

Fig. 5. Relationship between critical strain energy release rate and notch root radius (mild steel, 77K).

Fig. 6. Typical examples of the load-time curve and the

fracture aspect obtained in Charpy V notch impact test.

Testing temperature: -30•Ž, energy 68.8ft-lb,

shear fracture: 46%

(Fig. 5, ƒÏ•âl0). If it is assumed that GƒÏ=l*0=GC, then l*0=4l0. The physical meanings of the above values (for ex-

ample, l0) are still unclear [20]. At ƒÏ=0, however, the G

value is slightly larger than GƒÏ at some small ƒÏ. This may

be due to the process zone size effect at the notch tip. It

seems that the probability of the largest particle control-

ling the fracture involves changes according to the proc-

ess zone size. For example, in the KIC test, where the

stressed volume at the crack tip is nearly 1, the stressed

volume is about 20 in the Charpy test and about 400 in

the tensile test. It is therefore reported that the evaluated

critical fracture stress of obtained by each test increases

in the following order: tensile<Charpy<KIC [21]. This

seems to be important as a probabilistic aspect of fracture.

More details will be studied in the future.

143

Toshiro KOBAYASHI

3. DEVELOPMENT OF THE INSTRUMENTED CHARPY IMPACT TEST

The analysis of deformation and fracture behaviors, particularly the ductile-brittle transition behavior of steel, by recording the load-time or load-deflection curve dur-ing impact has been attempted since the 1920s [22]. Fig-ure 6 shows a representative load-time curve and fracture surface in the transition region [23]. Properties such as the general yield load (Py), maximum load (Pm), brittle fracture load (Pf), and brittle fracture arrest load (Pa) are clearly recorded. The absorbed energy is generally dis-cussed by dividing it into the premaximum load energy to obtain the nominal crack initiation energy (Ei) and into the postmaximum load energy to obtain the nominal crack propagation energy (Ep). It is possible to obtain considerable information from such analysis.

Kobayashi et al. classified the transition behavior of steel into 6 temperature regions based on the results of the analysis of the load-deflection curve and fracture surface morphology (Fig. 7) [23]. The outline of the definition of transition temperatures and the correlation with the results of a large-scale test are also shown in this

figure. In the transition region of the V-notch test, a cleavage crack appears after a fibrous thumbnail, i.e., ductile crack, formation beneath the notch root. There-fore, it is not possible to obtain KId directly from the fracture load (Pf). The situation is the same for obtaining the arrest fracture toughness KIa from the brittle crack arrest load (Pa).

Because of such limits, the use of a press-notched specimen inserted with a knife-edge was proposed in order to accelerate cleavage crack formation at the notch root [24]. However, there is some fear in this case that the material may be changed from a virgin state upon inserting the knife-edge. Kobayashi et al. presented a method of forming a carburized or nitrified layer on the surface and succeeded in obtaining a KId value corre-sponding to the one obtained in a large-scale test [23]. However, even in their methods, there are some difficul-ties in that the change of the material quality upon heat treatment must be avoided. Subsequently, although the precracked Charpy test had become the mainstream, it has been difficult in this method to obtain the dynamic fracture toughness based on linear fracture mechanics, except in extremely low-temperature regions. Moreover,

Fig. 7. Summaries of the test results on the transition behavoir in Charpy impact test.

144


a considerable inertial effect appears at such tempera-tures.

Under such circumstances, elastic-plastic fracture mechanics was proposed by Rice [25], and the validity of the J-integral value was recognized. The J-integral is generally valid for the initiation of a ductile crack. Then, the K (JIC) value converted from JIC can be used to esti-mate the lower boundary: of K measured in a large-scale test. Figure 8 shows such situations schematically [26]. The reference fracture toughness value KIR gives an es-timate of the lower bound of dynamic fracture toughness. There are two elastic-plastic fracture toughness parame-ters: JIC the crack initiation resistance and the R curve or material tearing modulus (Tmat); the gradient of the R curve represents crack propagation resistance. To deter-mine toughness, both parameters must be considered [20]. Although they are difficult to separate, the Charpy value includes both parameters and represents the toughness as a whole.

4. EVALUATION OF DYNAMIC FRACTURE TOUGHNESS

The author has developed a method of obtaining the propagation resistance Tmat by the analysis of the load-deflection curve of a single precracked specimen, as well as the dynamic elastic-plastic initiation fracture toughness value Jd (instead of JId when validity has not been verified). This method enabled the dynamic fracture toughness parameters to be obtained for the first time in the 1980s [11, 22, 27-29]. It is also applicable to the Instrumented impact test of brittle materials [30-32]. This method has been put into practical use in Japan and is called the Computer-Aided Instrumented Charpy Impact Testing System (CAI System) [33-35]. The ISO or JIS standards on the instrumentation of the Charpy impact machine (metallic materials-JIS B 7755, plastic materi-als-JIS B 7756) were established in 1993. The stan-dardzation of the method of determining the dynamic fracture toughness will be carried out hereafter.

I: Safe region on unstable fracture.II: Prevention region of unstable fracture initiation.III: Control region of unstable fracture initiation.

IV: Unstable fracture region.

Fig. 8. Schematic illustration on the relation between various K values and temperature.

The author had already completed such a system be-

fore 1986 [29]. First, the essential principle of the CAI

System for a ductile material will be introduced [36,37].

If the true crack initiation point (which usually exists

between Py and Pm) on the load-deflection curve of a

precracked specimen is known, it will be possible to ob-

tain JIC based on Rice's simple equation [38]. That is, JIC

is simply obtained using 2Ei/B (W-a), where Ei is the

potential energy until crack initiation. This point has

generally been detected in the static loading test using the potentiometric method, AE method, and unloading

compliance method. However, these methods are not

effective under dynamic loading. Therefore, the

load-deflection curve was first smoothed [27] using the

moving average method [29], and the point of sudden

change of the assumed elastic compliance at any point on

the load-deflection curve vs. initial elastic compliance

was detected (Fig. 9; compliance changing rate method).

This point corresponds to the true point of crack initia-

tion. This has been verified on many materials by the

experiment using a stop block [37,39].

In addition, in order to estimate the amount of crack

growth (ƒ¢a) from deflection, the key curve method was developed, where curve fitting of the load-deflection

curve is performed using the n-th power hardening law

[29,35]. Figure 10 shows the J-ƒ¢a curve for each case

[29]. For J in the controlled crack growth range (ƒ¢a•…

1mm), a good coincidence is ascertained. In the case of

accompanying crack growth, Garwood et al.'s correction

[40] and JM (modified J-integral) values are also shown.

The coincidence between predicted J-ƒ¢a curves and

those obtained in experiments using the stop block is

ascertained.

On the other hand, Ct (=Cs+Cm) is elastic compliance

which is the reciprocal of the initial slope of the

load-deflection curve (Cs is specimen compliance and Cm

is machine compliance). It is considered that for stored

energy before the maximum load, energy stored only in

the specimen should be extracted. In short, it is consid-

ered that Ei'=Ei•~(Cs/Ct) should be used instead of Ei in

Rice's equation. Each compliance shows impact velocity

dependence, as shown in Fig. 11 [41], Cs(mm/N) depends

on impact initial velocity V0 and is given as [41]

Deflection, ƒ¢m

Fig. 9. Schematic explanation of the compliance

changing rate method.

145

Toshiro KOBAYASHI

Fig. 10. J-ƒ¢a curves obtained from various methods.

Fig. 11. Relationship between impact velocity and compliances, Ct, Cm and Cs (Notched specimen: a/W=0.6).

Cs=(l2/BEW2)[Y+0.29{W2/(W-a)2}log(1/Vo)-0.399, (12)

where l: span distance (mm), B: specimen thickness (mm), Vo: impact initial velocity (m/s), E: Young's modulus (N/mm2), W: specimen width (mm), a: crack length (mm), and Y=27.11(ao/W)3-8.56(ao/W)2+1.77(ao/W)+0.829.

In the CAI System, the above energy correction is made. Although some doubts may be included in the application of this method up to the elastic/plastic range, such doubts will be neglected [41]. Such a correction becomes more important in the case of brittle materials.

In the impact test of brittle materials such as ceramics, oscillations due to the inertia effect increase considerably and it becomes difficult to measure load accurately [32]. One way to achieve such measurement is to apply the low blow impact test, where the inertial effect is greatly suppressed [30]; the significance of the impact test is retained in this case, because the impact velocity is re-duced to 1/2 and the strain rate is also reduced to only 1/2. Figure 12 shows examples of load-deflection curves in the low blow impact test using a JIS transverse speci-men [30]. Applied energy Eo must be over 3 times the fracture energy to maintain the original load-deflection

PSZ

SiC

Fig.12. Typical load-deflection curves of each specimen.

relationship [42]. This is a necessary condition for con-ducting the low blow impact test [42,43].

On the other hand, although the area under the load-deflection curve corresponds to the absorbed energy, it has been pointed out that energy, other than fracture energy of the specimen itself, is included in excess in the Charpy test of brittle materials [30]. Components of en-ergy Et are

Et=Es+Em=Ef+Ek+Em. (13)

Es and Em are energies stored in the specimen and the machine, respectively. Ef and Ek are deformation and fracture energy of the specimen and toss energy of bro-ken halves after fracture, respectively. An example of the results of the analysis of energy using the equation for compliance and the method in which collision between the hammer and specimen is assumed is shown in Fig. 13 [30]. It is obvious that Ef, assumed to be the true ab-sorbed energy, becomes about 38.1% in PSZ and 59.5% in SiC. The calculation of KId based on the static formula is possible for the precracked specimen in the low blow

146


Fig. 13. Analysis of absorbed energy in each specimen.

impact test. Such a program is included in the CAI soft-

ware.

However, a contactless condition appears in the

high-velocity impact test, between the hammer, specimen

and anvil. This phenomenon is called loss of contact [44].

Therefore, the application also becomes impossible even

in the analysis which takes the dynamic stress intensity

factor into consideration [31]. In such a case, there is no

alternative than to use the impact response curve method

with a strain gage fixed at the precracked tip [32]. Oth-

erwise, one may also consider the use of the one-point

bend test, where fracture is induced only by inertial load

[45]. Regarding the measurement of dynamic fracture toughness, the present situation is summarized in Fig. 14.

In particular, for the Jd value in the transition region, al-

though it is not strictly defined, the author applies the ƒÀ

IC method presented by Irwin for ductile crack initiation

[46]. This yields an estimate of the lower bound value.

The value of KIC(J) estimated using the ƒÀIC method is

given by

KIC(J)=KIC•ã1+1.4(KIC/ƒÐY)4/B2. (14)

Another important problem is the validity of the J

Fig. 14. Temperature dependence of plane strain fracture toughness (KIC, KId) and estimation methods of fracture toughness in each region.

Fig. 15. Change of Jin/ƒÐfs with ligament width.

value obtained. Whether or not a valid JId value can be

obtained from such a small specimen under the bending

condition has been investigated. According to ASTM

E813 or JSME S001, the valid condition for the J value

(Jin) and flow stress (arithmetic mean value of yield

stress and tensile stress) ƒÐfs is shown by

B,bo•†25(Jin/ƒÐfs), (15)

where bo is the ligament width (W-ao), ao is the crack

length and B is the specimen thickness. Figure 15 shows

the experimental results for bo of reactor pressure vessel

steel A508 [47]. In the case of dynamic loading, the coef-

ficient of 25 in equation (15) was reduced to 20 for bo.

However, for B, the coefficient became about 28. It has

been proven that the valid criterion under the static con-

dition can also be applied under the dynamic condition.

It is possible to obtain a valid JId for this steel at room

temperature by using the standard Charpy size specimen

if side grooves are added [47].

5. PROBLEMS OF MEASUREMENT

In ASTM, the method for obtaining dynamic fracture

toughness by the precracked Charpy test has been exam-

ined [26]. The low blow impact test in which fracture

147

Toshiro KOBAYASHI

occurs after 3ƒÑ (ƒÑ is the oscillation period) was proposed

[26,48] preliminarily, because it is difficult to record the true fracture load for fracture within 3ƒÑ.

ƒÑ=1.68(LWEBCs)1/2/Co. (16)

Co: speed of sound in specimen (5000m/s in steel), Cs:

specimen compliance, E: Young's modulus, L: span dis-

tance, B: specimen thickness, and W: specimen width. ƒÑ

becomes 33ƒÊs in steel.

However, many criticisms have been presented con-

cerning this criterion [49], and ASTM has since lost

momentum in its standardization activities. At present,

ISO seems to be more active in establishing the standard

[22]. In general, the response speed of electronic instru-

mentation is a problem. In the instrumentation used by

the author, f0 .9dB=800kHz (frequency at 0.9dB, where

the amplitude attenuates 10%; in the standard, it must be

over 100kHz), and the critical response time of the sys-

tem is estimated from the following equation:

TR=0.35/f0.9dB (17)

Therefore, the instrumentation has been confirmed to

have a sufficient measurement accuracy of TR=0.44ƒÊs.

Another problem is the accuracy of load measure-

Fig. 16. Effects of materials (a) and thickness (b) on the load calibration factor C. C is normalized with a calibration factor of A508 steel with 10mm thickness, CA508.

ment for fracture toughness determination. In JIS B 7755,

the measurement accuracy has been stipulated as •}2%.

The method of calibrating the output signal of the strain

gage fixed to the hammer (striker) has been provided in

JIS. In this case, the author has found that the calibration

value changes according to the test material and speci-

men thickness (Fig. 16). This becomes important, for

example, in conducting a test using a small specimen.

FEM analysis of this phenomenon shows that the strain

distribution at the striker changes according to the mate-

rial and specimen thickness [50]. Therefore, it should be

calibrated taking the material and the specimen geometry

into account.

On the other hand, the standard does not precisely

describe where the strain gage should be attached. Gen-

erally, in a striker where the strain gage is fixed, a slit is

cut to the striker to release the constraint in the sur-

rounding mass and to accelerate elastic deformation.

Moreover, hollow grooves are made for the protection of

the strain gage (Fig. 17). The result of examining these

effects is shown in Fig. 18 (•gupper•h in Fig. 18 shows the

position of the strain gage on the slit (see Fig. 17)). The

load-deflection curve recorded from the gage position of

15mm is smooth and does not oscillate, but other gages

show different oscillating curves. The end position of the

slit is near 30mm, and it was revealed that the open

hammer ends vibrate periodically in a bending manner

and that the gages near the slit end suffer such vibrating

Hollowed striker

Non-hollowed strikerFig. 17. Schematic illustration of instrumented striker for instrumented Charpy impact test (mm).

148


Fig. 18. Typical load-deflection curves recorded from specified strain gage positions in hollowed and non-hollowed strikers for the V-notched 6061-T6 Al alloy Charpy specimen.

effects [42,51,52]. Therefore, it is better to place the strain gage near the tip of the striker far from the end of the slit. The position at 15mm coincides with this condi-tion in our case. The effect of hollow grooves was negli-gible, as shown in Fig. 18. This was similar to the effect of different C- and U-type hammers [53].

The difference in the tip radius is recently becoming a problem, that is, the 2R or 8R striker. In the lower tem-perature range of the transition curve, there is little dif-ference, however in the higher temperature range of the transition curve, a large difference is observed (Fig. 19) [54]. Since the deformation behavior of an 8R striker at higher temperature becomes similar to that in 4-point bending [55], the adoption of a 2R striker seems reason-able.

Another problem is that, for example, when an alu-minum alloy is tested, it may be better to use an alumi-num striker than a steel striker. According to the test re-sult, a smoother curve was obtained [54]. The reason for this result may be the effect of mechanical impedance. Many improvements to the testing machine will be made hereafter. The dynamic calibration of load is another problem that must be studied in the future [56].

6. CONCLUSION

The Charpy impact test proposed by Charpy in 1901 spread widely during the last century, and is an excellent method of evaluating the notch toughness of

Fig. 19. Ductile-brittle transition curves obtained using R2 and R8 strikers on A508 steel.

materials simply and inexpensively. On the other hand,

the instrumented Charpy impact test, which provides

more information on fracture, has also been developed.

At present, the instrumentation itself has been standard-

ized and a method of measuring the dynamic fracture

toughness is the next issue. The EU is enthusiastic about

such standardization, particularly the European Structural

Integrity Society (ESIS). The author has already devel-

oped the CAI system and has put it to practical use in

Japan. Worldwide cooperation is desirable to establish

the new standard.

The Symposium on Pendulum-Type Impact Test-

ing-A Century of Progress was held by ASTM in 1999

[50]. In Germany, the special issue on 100 Jahre

Charpy-Versuch was published in Mat.-wiss. u. Werk-

stoffe Vol.32 (2001), No.6. The Charpy Centenary Con-

ference (CCC2001) was held in France in October, 2001

[52]. The author has studied this method for a long time, but he expects further development in this century

as well.

Acknowledgment-The author would like to thank

MSRI for the permission to publish the English version

of this article which originally appeared in Zairyo [vol.

41 (2002)] in Japanese.

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Toshiro KOBAYASHI

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