Evolution of the Charpy-V test from a quality control test to a materials evaluation tool for structural integrity assessment

From Charpy to Present Impact Testing D. Franqois and A. Pineau (Eds.) 9 2002 Elsevier Science Ltd. and ESIS. All rights reserved

57

E V O L U T I O N OF T H E CHARPY-V T E S T F R O M A Q U A L I T Y C O N T R O L T E S T TO A MATERIALS EVALUATION TOOL FOR STRUCTURAL INTEGRITY ASSESSMENT KIM WALLIN, PEKKA NEVASMAA, TAPIO PLANMAN, MATTI VALO VTT Manufacturing Technology P.O. Box 1704, FIN-02044 VTr, Finland

ABSTRACT Originally, the Charpy-V test was used mainly as a quality control test. However, after World War II, with the development of the transition temperature philosophy, the Charpy-V test evolved into a tool for material selection and toughness evaluation. With the development of fracture mechanics, further evolution of the interpretation of the Charpy-V test has made it a quantitative materials evaluation tool for fracture mechanics based structural integrity assessment. This presentation will give an outline of the evolution of the Charpy-V test, focussing on the latest developments regarding its use in structural integrity assessment.

KEYWORDS Charpy-V, fracture toughness, brittle fracture, ductile fracture, correlations, material selection.

HISTORICAL PERSPECTIVE Originally, the Charpy-V test was used mainly as a quality control test. However, after World War II, the analysis of the failures in welded merchant ships changed the nature of the Charpy-V test more to a design tool. Out of 4694 ships made during the war, 1289 experienced serious or potentially serious fractures. The ship plates where the crack had initiated showed generally a lower impact energy (KV) at the failure temperature, than the plates where the fracture had arrested. This led to the introduction of the transition temperature concept. It was found that generally, the plates with initiation had an impact energy below 10 ft-lb (13.6 J) and the ones where fracture ended had more than 20 ft-lb (27.1 J) [1,2]. A statistical analysis of the "initiation" plates gave as a result a 95 % upper bound value of 15 ft-lb (20.3 J). These investigations led to the even presently used transition criteria of 2()/21 J and 27/28 J. In further studies, it was found that the 15 ft-lb transition temperature for the ship steels in question also correlated with the ESSO wide plate transition temperature [3]. The 10 ft-lb transition temperature, on the other hand was correlated to the explosion bulge test transition temperature which corresponds to NDT [4]. However, when more modem fully killed fine grained steels were examined, it was found that the correlations were not the same [3,4]. The Charpy-V energy absorption at the ESSO transition temperature or

K. WALLIN, ET AL.

58

NDT was higher for fully killed steels. In terms of the ESSO transition temperature, specifically the 30 ft-lb (40.6 J) energy level was found appropriate [3] (Fig. 1).

[~o.,~" . : ~ , , ~ 41 ,J 9

....: ~

,~u~sc:=,,, c

'

I

L,,~ -___~~ _ _ ~ _ ~ ~

! l

2O

,~. ~o

9 : ' ~ 1 7 6 :t

.1

~..~o -20

. .~

O

*40

.

-30

0

-20

-10

.

.

.

0

10

20

30

T,.~, or T,,., [*C]

Fig. 1 Comparison of ESSO test transition temperature and Charpy-V transition temperatures for conventional and modem steels [3]. With continuing research, it was found that also the materials strength level required different transition criterion to be used. The idea was that irrespective of strength level, each material should have the same deformability. The parameter correlating with deformation, in the Charpy-V test is the lateral expansion (LE). Thus, a constant LE (15 mil = 0.381 mm) was considered as a good transition criteria [5]. In terms of energy, this lead to an increased impact energy demand for higher strength steels. Since LE is a result of plastic deformation, it depends on the strength properties of the material. Specifically, the relation between impact energy and LE is controlled by the materials ultimate strength, (~t', (Fig. 2 [6,7]). For some reason, this relation has not become widely recognised. The energy requirement has been connected to the materials yield strength instead. This pcnalises unduly higher strength steels where the yield/ultimate ratio is larger.

a)

b) o

~

=389MPI o,., ,, ~I0 MPan ou

.o j.. ~

0

- O.OJ

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0.10 ~

0.0o

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c [',:;

0.04

0.02

9 'n~~176176

0.~

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0.06

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o u 9 1034 MPI . o,., 9 1103 MPI

9

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1

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.

.

.

o.o

.

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.

0.8

'

1.o

'

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o.s

~.o

~.s LE

2.0

2.s

[n',m]

Fig. 2 Relation between Charpy-V impact energy and lateral expansion for a) ASTM tup [6] and b) ISO/DIN tup [7]. Another parameter that became popular for determining a transition temperature is the fracture appearance describing the percentage of ductile fracture (SA). A typical criterion is 50 % ductile fracture (FATTs0,-;). The philosophy with this criterion is that at 50 % ductile fracture appearance, the crack arrest properties in the material are assumed to be sufficient to safeguard against fracture propagation in the structure. Also this parameter was correlated

59

Evolution of the Charpy- V Test

against the NDT [5]. Instead of fracture appearance, proportional impact energy could be used as well. Figure 3a shows the proportional correlation between impact energy and SA. The parameters are linearly related, but with an offset of approximately 5 %, since energy is absorbed, even when there is no ductile fracture. Because of the parameters close relation, also the correlation between the two transition temperatures is well defined (Fig. 3b). The small offset, causes the TK50%US.(50 % of upper shelf energy) to be about 8~ lower than

FATr50 %. a)

-

,

-

,

.

,

.

,

-

b)

,

IO0

,

-

,

-

,

.

,

-

,

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,

-

,

-

,

-

,

.

,

.

IO0

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Q

9

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~~.

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.

.

~ 0

: 2O

. 4O

.

.

0 o .

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60

jo

]

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~ 0

9 .

o

4O

%f%"

o

O

I

I

IN)

100

FAI-I'w~ rCl

sA 1%1

Fig. 3 Relation between a) Charpy-V impact energy and fracture appearance [7] and b) TK50%us.and FATI'50 % [8]. The classical transition temperature criteria, were mostly concerned with determining a material condition which would make brittle fracture propagation difficult, regardless of structural details. Thus the transition criteria had to be fulfilled at the lowest service temperature of the structure, i.e. if the service temperature is -20~ the selected transition criteria had to be fulfilled at -20~ New problems arouse when the effect of plate thickness was added to the transition criteria. It was found, e.g. with drop weight tear tests (DW'Iq'), that the transition temperature was also a function of plate thickness [9] (Fig. 4). This thickness effect led to an additional temperature correction to the transition criteria, i.e. the original idea of having a specific transition criterion for the lowest service temperature became muddled. First, the materials yield strength affects the transition criterion to be used and second, the plate thickness leads to an additional temperature adjustment to where the transition criteria must be fulfilled. Presently most workmanship standards, pressure vessel and ship codes are based on such a simple, yet obscure, methodology.

lO

/ o o,,.,.*.,m

~'

70 ~

I

K/

A w r s g e ot t~ Tests

"

-10

~-~ 4O 4

8

Th~.kne.

16

[mr.]

Fig. 4 Effect of specimen (plate) thickness on DWTT FATT50 % [9].

60

K. WALLIN. ET AL.

FRACTURE MECHANICAL PERSPECTIVE The design criteria took a large step forward with the introduction of fracture mechanics. Useful parameters for the assessment of critical steel (metal) structures based on fracture mechanics are parameters, which are capable of describing a component's resistance against flaws. Such parameters are e.g. the fracture toughness Ktc (Kjc) , the ductile initiation toughness Jtc, the ductile tearing resistance, J-R, and the crack arrest toughness, K~a. In case of steels K~c usually describes the materials resistance against brittle cleavage type fracture, but for e.g. aluminium alloys and some extra high strength steels it may also correspond to ductile initiation toughness. Since the Charpy-V test does not provide a direct measurement of fracture toughness some kind of correlation must be used. Numerous different empirical correlations relating impact energy to fracture toughness have been determined for a variety of materials, over the past years [10-15]. Finding an empirical correlation that would be universally applicable has proven to be quite difficult. Even though both tests describe the materials fracture behaviour, they have many important differences. Due to the differences in the tests, empirical correlations are usually very case dependent. It is also difficult to decide which correlation to use in a given case. The British pressure vessel standard, BS5500 App. D, introduced in 1976, was still quite crude. It is not based on a real fracture mechanical model. Instead, it is based on empirical correlations between Charpy-V impact tests and wide plate tests [16]. A major development was made when Sanz [17] developed the new fracture mechanics based brittle fracture Charpy-V transition criterion for the AFNOR standard [18]. Sanz found that the fracture toughness temperature dependence was quite material insensitive and thus enabled a general correlation between the KIc 100 MPa'~lm and TK28J transition temperatures. The method applies linear elastic fracture mechanics and uses an empirical plate thickness correction which implicitly contains also the assumed flaw size information. Additionally, the model also includes a loading rate correction in the form of a strain rate related temperature shift. The Sanz model was subsequently modified by Sandstr6m [19]. The two models use different CVN-K~,: correlations and treat the empirical section thickness correction differently, but use the same temperature and strain rate dependence for the fracture toughness and treat all stresses as tensile stresses. Sandstr6m furthermore, limits himself to a quarter thickness deep surface crack with an undefined width. The biggest difference in the models, lies in the treatment of the empirical thickness correction. Sandstr6m connects the degree of thickness correction to the stress level, so that the lower the stress level, the smaller the thickness correction. SandstrOm's definition of crack size leads to that the thickness correction should be applied only at stress levels d a y > 0.71. The Sandstr6m model forms the basis of the present Swedish Pressure Vessel Code. Subsequently, Wallin [8,20] re-assessed the Sanz Charpy-V - fracture toughness correlation using the new Master Curve method [21], which accounts for the intrinsic statistical size effects in fracture toughness (description given in state of the art section). This new correlation, and size adjustment, forms the basis of the new EUROCODE 3, annex C brittle fracture assessment. It is also included in the new BS 7910 fracture assessment standard and forms the default level of fracture toughness estimation in the European SINTAP fracture assessment procedure [22]. All of these new more sophisticated methods apply elastic plastic fracture mechanics and enable an assessment with respect to fracture initiation, contrary to the simple methods which only provide assessment against fracture propagation.

6!

Evolution of the Charpy- V Test

STATE OF THE ART The SINTAP method represents the present state of the art in structural integrity assessment. An imperative part of the method is the determination of the materials fracture toughness. It is clear that one cannot simply correlate the impact energy directly with the fracture toughness. One must first clarify which parameters are realistic to correlate. In order to do this the basic features of each test must be examined separately to see which features correspond to the same physical event. This way it is possible to derive comparatively simple, adequately accurate, general correlations between parameters determined from the Charpy-V impact test and brittle fracture initiation, ductile fracture initiation, tearing resistance and even crack arrest toughness (from instrumented Charpy-V test). Examples of this are presented next.

Brittle Fracture Initiation

The Master Curve method enables a complete characterisation of a materials brittle fracture toughness based on only a few small size specimens. The method is based on a more than 15year research at VT'F Manufacturing Technology and has led, e.g., to the ASTM standard E192 l, the first standard that accounts for the statistical specimen size effect and variability in brittle fracture toughness. The Master Curve method has been shown to be applicable for practically all steels with a body-centred cubic lattice structure, generally identified as ferritic steels. The method combines a theoretical description of the scatter, a statistical size effect and an empirically found temperature dependence of fracture toughness (Fig. 5). The fracture toughness in the brittle fracture regime is thus described with only one parameter, the transition temperature To.

I "THE MASTER CURVE APPROACH" 1 MASTER CURVE ANALYSIS

TYPICAL RAW DATA _

-

AD,JUSTMENT /

.,/

SMALL $.oEOMEi~,," "" ~,'''"'5 %

T CcI

~ ~ _

y

'

T COl

THEORETICAL SCATTER DESCRIPTION STATISTICAL SIZE ADJUSTMENT UNIFIED TEMPERATURE DEPENDENCE

Fig. 5 Basic principle of the Master Curve fracture toughness description. The To transition temperature has successfully been correlated to the Charpy-V 28 J transition temperature [8, 20]. The correlation is shown in Fig. 6 a-d. Besides showing the correlation, Fig. 6 also indirectly verifies some of the assumptions of the Master Curve method. It verifies the size adjustment since the correlation is unaffected by specimen size (Fig. 6 a and b) and it

62

K. WALIJN. ET AL.

verifies the validity of the elastic plastic fracture toughness since the correlation is unaffected by parameter type (Fig. 6 a and c). As mentioned earlier, the correlation has been implemented into all major fracture mechanics based assessment methods. a)

EPFMKjc B = 2 5 m m T : TK:-I&;C"~

.

.

.

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EPFM K~ !1=10...200 mm

b)

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150

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--o

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~~ -11111

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-150

-50

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0

50

-150

IOO

-I00

c)

I.B~K,,: B,25...250mm

0

-50

50

100

rK~. Ccl

TK~ [~

d)

o .300..1200MPa

~

tit K~ !11=10...250 mm

o, = 300...1200 MPI

150

,OOll~

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I

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5O

go

-

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-

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_

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e "1~176 1

~

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.... -2oo

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o CC]

~o

~oo

~so

" -200

-lS0

-100

-60

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SO

100

150

"rK,.~ CC]

Fig. 6 Relation between TKz8J and Master Curve T o for a) elastic plastic fracture toughness from 25 mm thick specimens, b) elastic plastic fracture toughness from other than 25 mm thick specimens, c) linear elastic fracture toughness and d) combination of all results. (Irr. stands for irradiated.) The TKz8J - T O correlation does have a few restrictions. Since it is based upon cleavage fracture, the correlation should not be used for steels with upper shelf energies below 70 J, where the 28 J energy level contain more than 20 % ductile fracture appearance, or where the brittle fracture mode is something else than cleavage fracture (eg. grain boundary fracture, low energy tear etc.). The correlation should neither be used for strongly inhomogeneous materials which may show a so called pop-in behavior in the fracture toughness test [8]. In such cases it is impossible to determine the fracture toughness from the Charpy-V test and therefore actual fracture toughness testing is required. Normally these restrictions are not an obstacle for the use of the correlation to determine the lowest allowable service temperature. Figure 7 [23] gives an example of the application of the correlation for extra high strength steels. The test geometries consisted of U- and D-profiles tested in bending. The U-profiles contained two edge cracks, one on each flange, and the D-profiles contained an elliptic surface flaw on the upper web, or a through thickness crack. In Figs. 7 b-d the estimated temperatures Tc,~ (based upon the fracture load), for all tests showing brittle fracture, are compared with the

63

Evolution of the Charpy- V Test

actual test temperatures Tnx..,.,~. A total of 68 tests showed brittle fracture. Fig. 7 b is a "best" estimate corresponding to 50 % failure probability and the mean Charpy-V - K~-correlation (50 % confidence level). Figs. 7 c and 7 d correspond to a 85 % confidence level with respect to the correlation and failure probabilities of 20 % and 5 % respectively. For the best estimate, 4 1 % of the results are non-conservative. The corresponding percentages are, for the other cases, 9 % and 3 %. Based upon a Monte Carlo simulation the theoretical expectation limits for non-conservative estimates were calculated, assuming all parts of the methodology to be valid. For the best estimate the percentage of non-conservative results should be, with a 50 % probability between 47 % to 56 %. The corresponding percentages are for the other cases 9 % to 15 % and 1% to 4 %. The percentages of non-conservative estimates comply very well with the theoretical expectation values. Overall, the predictions are accurate or conservative

a) TEST SPECIMEN GEOMETRIES (3-POINT LOADING)

u> 0

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O

.50

nn

i _~

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-150

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-120

-100

-80

-60

-40

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20

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T~'cl 10,0

c)~

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-120

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.

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.

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i

.

-60

i

.

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.

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!

20

.

40

T[~

Fig. 7 Prediction of failure behaviour of different test profiles a), for best estimate b), estimate for non-critical case c) and estimate for critical case, compared to test results for extra high strength steels [23].

Crack Arrest The mechanism of brittle crack arrest differs from that of initiation. Thus the scatter and size effects are not the same as those found for brittle fracture initiation. Mechanistically, arrest occurs when the local crack driving force at the crack tip decreases below the local arrest toughness over a sufficiently large portion of the crack front. A single local arrest is not sufficient to arrest the whole crack front, i.e. the scatter is more a function of the mean

64

K. WALLIN, ET AL.

properties of the matrix (and not the local). Therefore the scatter is less than for initiation and there are no statistical size effects in the case of crack arrest. It has been shown that the temperature dependence of crack arrest toughness is the same as for brittle initiation toughness [24]. The main difference is a material dependent temperature shift between the two properties [25]. This similarity to the initiation behaviour, in principle enables a correlation between crack arrest toughness and the Charpy-V test. However, the parameters normally determined from the test produce rather poor correlations with K~.~. Depending on which parameter is chosen, different trends are seen. E.g. in the case of TK41j, yield strength does not affect the correlation, but a low upper shelf energy does (Fig. 8). For LE, also the yield strength will affect the correlation. In the case of FATI', the upper shelf energy has less effect, but the yield strength has an effect.

l

"

|

"

I

"

I

I

~

|

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,~

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Oo

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1

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-60

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-

-

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/

-40

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0

o

0

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45O

~w IJl

o, (MPa]

Fig. 8 Effect of yield strength a) and upper shelf energy b) on the relation between TK[~100 and T K 4 1 J [24]. The best information regarding crack arrest can be obtained from the instrumented Charpy-V impact test in terms of the 4 kN crack arrest load temperature (Fig. 9 [25]). This is a parameter that physically corresponds to an actual crack arrest event and provides therefore a true measure of the materials crack arrest properties. The conventional parameters are reacting to several other factors than crack arrest and therefore, they can never be accurately correlated to crack arrest on a general basis.

L O A D C H A R A C T E R I S T I C S IN INSTRUMENTED IMPACT TEST

a)

.

.

.

.

.

.

.

.

_

b)

.

F,, - ~ n e r ~

F.

F.

F.

F.

12

F.

yk~g

T~.- T

+ 13.5"C

. max. force . c~av~ Inlflatlo~ .

crack anvst

~ s s

0 9

,

1

1

1

2

.

1

,

3

DISPLACEMENT (mm)

|

4

.~::F

Os.

p-

,

,

-50 -50

0

50

100

T~.fC] Fig. 9 Definition of crack arrest load (F~) in instrumented impact test a) and correlation between 4 kN arrest load temperature and TK[~0o b)

[25].

150

Evolution of the Charpy- V Test

65

Ductile Fracture

The Charpy-V notch test provides information about the energy needed to fracture a small specimen in half. On the upper shelf this energy relates to ductile fracture resistance and it is possible to correlate it to the J-R -curve. Recently, 112 multi-specimen J-R -curves from a wide variety of materials were analyzed and a simple power law based description of the J-R curves was correlated to the CVNus energy (Fig. 10 [26]). This new correlation corresponds essentially to a 5 % lower bound and conforms well with the earlier correlations, regardless of the definition of the ductile fracture toughness parameter. The correlation gives a conservative tearing resistance estimate in terms of the J-integral as follows:

mmJ

(1)

p~ T - 2 0 / [kJ/m: j, oC] - 400 J . . . .

(2)

J = Jlmm 9Aa m ... lkJ/m 2 , where Jlmm = 0.53" CVNt.s

"~

.ex

and m=0.133.CVNt, s

0.:,6

pf T - 2 0 / ov .ex.~, - 2 0 0 0 ) - 4664 +0.03 "'" [J, ~

MPa]

(3)

With the above expressions, practically all previous ductile fracture CVN-"K~c" correlations can be explained [26]. The expressions can be used either to estimate Jlc from equation (4) or to estimate the J-integral at any amount of ductile tearing between 0...6 mm for temperatures lower than 300~ Jr(- / l/m J r ( - + 0.2 mm =0 2"Of ~.Jlmm .)

(4)

where the flow stress can be approximated by:

af = a v .

a)

,

,

,

i

lSOO I J,_. o.7,-c~'-1

,

1+

,

, 6 .,.

150MPa

(5)

Oy b)

o.8

~ O.6

, %_=1r-/.o

t . J t e r t u re d a t a

] o ~ o .-6.... E 0.4 (12 r

0

50

100

150

200

250

300

0

o 9

200

400

tk:)O O00

o=0.1 t.JWature data VI"T d m

1(100 1200

1400

1600

a,.,. [~m']

Fig. 10 Relation between CVNtJs and the J-integral value at 1 mm crack growth a) and relation between J-R curve shape (described by power m) and the J-integral value at 1 mm crack growth b) [26].

66

K. WALLIN. ET AL.

Treatment of Sub-size Specimens

When the plate thickness is less than 10 mm, testing with standard sized Charpy-V notch specimens is impossible. In such cases the testing must be based on subsized specimens. The difficulty lies in extrapolating the result from the subsized specimen to correspond to the result from a standard sized specimen. Basically two different methodologies can be used. The extrapolation can be based either directly upon the measured parameter, e.g. impact energy KV or on some transition temperature criterion. The presently used simple codes apply the direct extrapolation of the impact energy. Specifically, for a 5x10 mm 2 specimen, normally the energy is multiplied by 1.5 to make it correspond to a full size specimen. The problem with the direct extrapolation lies in the fact that the specimen thickness yields different effects in different regions of the transition (Fig. 11 [8]). On the lower shelf subsized specimens yield proportionally higher impact energies as compared to standard size specimens. They may even produce higher absolute energies than a full size specimen. On the upper shelf the behavior is reversed so that subsized specimens yield either proportionally equal or even lower impact energies than standard sized specimens [27-29]. The reason for this is that the different fracture micromechanisms yield different specimens thickness effects. In the transition region there is a competition between ductile and brittle fracture micromechanisms thus yielding a very complex combined thickness effect. This effectively invalidates the method of direct extrapolation which is used today. A much more reliable extrapolation of the brittle fracture behaviour is obtained by using a transition temperature criterion (Fig. 12 a [301). For the upper shelf energy a separate correction, which accounts for the transition to shear fracture with decreasing specimen thickness, has also recently been developed (Fig. 12 b [30]). Thus sub-sized specimens can be used both to estimate brittle fracture initiation as well as ductile tearing.

100

,-,

6O

o o

o o (9

~ "

o o

o o

2O

I ~ ,,,, (Jl

Fig. 11 Dependence between impact energies measured from full- and sub-size CVN specimens.

BEYOND STATE OF THE ART Even though the correlations presented here take into consideration the physical aspects of the fracture parameters being correlated, they are still empirical in nature. The next generation of CVN correlations will undoubtedly take the step forward to a fully theoretical interpretation of the Charpy-V test. Already, efforts are undergoing to model the Charpy-V impact energy with the help of the French local approach [31] and other similar micromechanism based

67

Evolution of the Charpy- V Test

models. Eventually, also fracture assessment may change so that conventional fracture mechanical parameters like KI(- and J may become obsolete and only material parameters used in the micromechanical models are needed. Then, the Charpy-V test may become a true materials evaluation tool.

a)

40 2O

THICKNESS ~ C T I O N , ,

,

FOR CVN TRANSITION , ,

b)

=

T,.~ 9 T=.,_. - sl.4.~{z.(Bno)"-1)

o

'•

-~-- B = S mm I ~ B-, 3.3 mm [ i ~>- 9 - 2.s mmol~

o

3O0

~

t4o

i

t-

-40

~ I00

-8O -100

0

2

4

6 O

8

[nw)

10

12

0o

100

150

290

250

30O

3~

~,,~(o,8x8) [~cm*]

Fig. 12 The effect of specimen thickness on CVN 35 J/cm 2 transition temperature a) and upper shelf energy b) [30].

SUMMARY AND CONCLUSIONS Originally, the Charpy-V test was used mainly as a quality control test. However, after World War II, the analysis of the failures in welded merchant ships changed the nature of the Charpy-V test more to a design tool. The ship plates where the crack had initiated showed generally a lower impact energy (KV) at the failure temperature, than the plates where the fracture had arrested. This led to the introduction of the transition temperature concept. The classical transition temperature criteria, were mostly concerned with determining a material condition which would make brittle fracture propagation difficult, regardless of structural details. Thus the transition criteria had to be fulfilled at the lowest service temperature of the structure. Problems arouse when the effect of plate thickness was added to the transition criteria. This led to an additional temperature correction to the transition criteria, i.e. the original idea of having a specific transition criterion for the lowest service temperature became muddled. First, the materials yield strength affects the transition criterion to be used and second, the plate thickness leads to an additional temperature adjustment to where the transition criteria must be fulfilled. Presently most workmanship standards, pressure vessel and ship codes are based on such a simple, yet obscure, methodology. The SINTAP method represents the present state of the art in structural integrity assessment. An imperative part of the method is the determination of the materials fracture toughness. By analysing the Charpy-V test on a physical basis, it has been possible to derive comparatively simple, adequately accurate, general correlations between parameters determined from the Charpy-V impact test and brittle fracture initiation, ductile fracture initiation, tearing resistance and even crack arrest toughness (from instrumented Charpy-V test). Examples of this have been presented. The next generation of CVN correlations will undoubtedly take the step forward to a fully theoretical interpretation of the Charpy-V test. Then, the Charpy-V test may become a true materials evaluation tool.

68

K. WALLIN,ET AL.

ACKNOWLEDGEMENTS This work is a part of the Structural Integrity Project (STIN), belonging to the Finnish Research Programme on Nuclear Power Plant Safety (FINNUS), performed at VTI" Manufacturing Technology and financed by the Ministry of Trade and Industry in Finland, the Technical Research Centre of Finland (V'lq') and the Radiation and Nuclear Safety Authority (STUK).

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