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NISTIR 88-3099

FRACTURE BEHAVIOR

OF A PRESSURE VESSEL STEEL

IN THE DUCTILE-TO-BRITTLE

TRANSITION REGION

Jurgen Heerens

D.T. Read

National Institute of Standards and Technology

(formerly National Bureau of Standards)

U.S. Department of Commerce

Boulder, Colorado 80303-3328

December 1988

100

- U 56

88-3099

1988

C.2

Stimulating America's Progress

1913-1988
NATIONAL INSTITUTE OF STANDARDS &

TECHNOLOGY

Research Information Center

Gaithersburg, MD 20699
NISTIR 88-3099

AJ/src

Q-Cioo

« /U) S’?—

FRACTURE BEHAVIOR ?f

?

OF A PRESSURE VESSEL STEEL

IN THE DUCTILE-TO-BRITTLE

TRANSITION REGION

Jurgen Heerens*

D.T. Read

Fracture and Deformation Division

Institute for Materials Science and Engineering

National Institute of Standards and Technology

Boulder, Colorado 80303-3328

*GKSS Research Center

Geesthacht, West Germany

December 1988

U.S. DEPARTMENT OF COMMERCE, C. William Verity, Secretary

Ernest Ambler, Acting Under Secretary for Technology

NATIONAL INSTITUTE OF STANDARDS AND TECHNOLOGY. Raymond G Kammer. Acting D t

CONTENTS

Page

1. INTRODUCTION 1

2. MATERIAL 2

3. EXPERIMENTAL DETAILS 4

4. RESULTS 4

4.1 Fracture Behavior of the Various Specimens 4

4.1.1 CT-Specimens 4

4.1.2 Blunt-Notched Bend Bars 5 4.1.3 Tensile Specimens 5 4.2 Frac tographic Examination 6

4.2.1 Blunt-Notched Bend Bars 7 4.2.2 Tensile Specimens 7 4.2.3 Location of Initiation Sites 7 4.3 Metallographic Investigations 9

5. DISCUSSION 9

5.1 Fracture Stresses in the Various Specimens 9

5.1.1 CT-Specimens 9

5.1.2 Blunt-Notched Bend Bars 12

5.1.3 Tensile Specimens 12

5.2 Types of Initiation Sites 13

5.3 Scatter of the Fracture Toughness 16

5.4 Some Aspects on the Lower Bound Fracture Toughness and the

Transition Temperature 17

6. CONCLUSIONS 20

7. ACKNOWLEDGMENTS 20

8. NOMENCLATURE 21

9. REFERENCES 22

iii

Fracture Behavior of a Pressure Vessel Steel in the Ducti le - to - Britt le

Transition Region

Jurgen Heerens* and D.T. Read

Fracture and Deformation Division

National Institute of Standards and Technology

The reasons for the scatter of fracture toughness in the ductileto-brittle transition region, as well as the mechanisms leading to

cleavage fracture, have been investigated for a quenched and

tempered pressure vessel steel, DIN 20 MnMoNi 55. The fracture

surfaces indicate that cleavage fracture starts at one small area

in the ligament, the cleavage initiation site. Cleavage initiation

occurs ahead of the crack tip at the location of the maximum normal

stresses. Fractography and metallography show four different types

of initiation sites. The mechanisms which may trigger cleavage

fracture at these initiation sites are discussed. The results indiicate that the scatter of fracture toughness is due to the scatter

in the distance between the cleavage initiation site and the fatigue

crack tip.

Key words: cleavage; fracture stress; inclusions; initiation site;

mechanism; origin; scatter; weak link.

1 . INTRODUCTION

For characterization of the due tile - to -brittle fracture transition

behavior of a steel, the transition curve is normally determined by measuring

the fracture toughness as a function of temperature. In connection with the

experimental determination of the transition curve, the following facts are generally observed:

- The fracture toughness in the lower transition region exhibits high

scatter, which increases with increasing temperature. This scatter ban!

reaches its maximum in the upper transition region. Because of the la:

scatter, many specimens must be tested in order to characterize tie- fracture toughness of the material in this region.

* Permanent Address: GKSS Research Center, Geesthacht, West Germany.

1
- The transition curve frequently depends on the size and geometry of the

specimen as well as on the loading rate used in the fracture test.

Consequently, the test results are not directly applicable to a crack in

a structure which has a constraint or loading rate different from the

test specimen.

Current research on these problems is discussed in [1—3],

The microstructural mechanisms leading to cleavage fracture for quenched

and tempered steels are not yet well understood. The literature reports

different results regarding the mechanisms of cleavage fracture initiation for

this kind of microstructure. The fracture of carbide particles larger than 1

/im seems to trigger cleavage fracture in many mild steels [4]. The carbide

particles in quenched and tempered steels are much smaller than 1 /im and it is

believed that other microstructural parameters, such as the bainite packet

size [5] or inclusion clusters [6], are relevant for cleavage fracture. A

recent investigation [7] on a tempered MnMoNi steel showed that cleavage

fracture proceeds by a micromechanism which involves dislocation interaction.

The present investigation deals with cleavage fracture in a quenched and

tempered pressure vessel steel, 20 MnNoNi 55. This steel is similar to type

A533B. The objective was to investigate the scatter of fracture toughness in

the transition region as well as to examine the micromechanisms which trigger

cleavage fracture. The fracture surfaces of CT-specimens , blunt notched bend

bars, and tensile specimens were investigated. In each specimen one cleavage

initiation site was found. Four different types of initiation sites, which

have different fractographic features, were classified. The results show that

the scatter of the fracture toughness is caused by the variation of the

location of the cleavage initiation sites in the ligament. The fractographic

and metallographic results are used to discuss some possible cleavage

initiation mechanisms as well as some aspects on both the lower bound

toughness and the transition temperature.

2 . MATERIAL

The material used in this study was DIN 20 MnMoNi 55 steel from a forged

boiler bottom segment. The chemical composition of this steel is similar to

2
ASTM A533B; see table 1. The heat treatment of the material is shown in

figure 1. The size of the heat-treated block was about 600 x 300 x 300 mm. All specimens were machined from three adjacent 25 mm thick slices taken from

one heat treated block. The orientation of the fracture surfaces was the same

for all types of specimens. The microstructure of the heat treated material

is shown in figure 2a. The average ferrite grain size of the alloy is about

30 nm. Rod- like carbide particles which are arranged in parallel rows can be

observed in the individual ferrite grains. The size of the carbide particles

is about 0.2 to 0.5 /im. Large isolated inclusions as well as clusters of

these inclusions can be observed on polished surfaces. A typical example of

these inclusion clusters is shown in figure 2b.

Table 1. Composition of 20 MnNoNi55 (— A533B) in mass percent

C Si Mn P S Cr V

0.19 0.2 1.29 0.007 0.008 0.12 0.02

Cu

0.11

Al

0.015

Ni

0.80

Mo

0.53

Sn

0.012

Co

0.014

As

0.030

Sb

0.030

Ti

0.050

Fe

bal

3
3. EXPERIMENTAL DETAILS

The specimens used in this study are shown in figure 3. For the

determination of the fracture toughness, precracked CT-specimens were used.

All specimens were fatigue precracked at room temperature with a load ratio of

^min/^max = 0-1- The Kmax -value for the final 1 to 2 mm of precracking was about 500 to 600 N/mm3 / 2

. All specimens were side grooved 20 percent after

precracking. The J- integral was calculated according to ASTM E 813—81 [8] by

using the analysis for the multiple specimen technique. The blunt-notched

bend bars were machined and tested under 4 point bending as shown in [9]

.

The

crosshead displacement rate was 0.5 mm/min for the CT- specimen and bend bars.

The tensile specimens were tested with a displacement rate of 1 mm/min. For

cooling the specimens, a bath of ethanol cooled with liquid nitrogen or a bath

of pure liquid nitrogen was used, depending on the test temperature. In

either case the temperature of the specimen was monitored by a thermocouple

attached to the surface of the specimen or located within a small hole in the

specimen. For the metallographic examinations, the surfaces were electropolished using a mixture of 25g Cr0 3 in 133 ml acetic acid. Etching

was done with 2 to 3 percent nital. For qualitative analysis of the

inclusions, an energy dispersive analyzer combined with a scanning electron

microscope (SEM) was used.

4 . RESULTS

4.1 Fracture Behavior of the Various Specimens

4.1.1 CT-Specimens

The critical J- values, J c , for cleavage fracture of 30 CT specimens, were measured in a temperature range of —90 to —45 °C. In order to determine the

scatterband of J c in this temperature range, a number of specimens were tested

at —90 and —60 °C, respectively. The resulting data, figure 4, show that the

scatter in J c increases significantly with increasing temperature. For

temperatures higher than -60 °C the scatterband seems to increase further.

Within the range of —60 to —90 °C, the J-level for initiation of ductile

tearing, J^, was found to be in the range of 100 to 150 N/mm. At this J^-

level, local dimpled rupture can be observed at the crack front. This J-j_-

4
level was determined by fractography using an SEM. The Jj c -value of the

material at room temperature is about 200 to 300 N/mm as shown in a previous

study '10]. The toughness values in figure 4 represent the fracture toughness

of the steel in the upper transition region, that is, the region in which

cleavage fracture with or without ductile tearing can occur.

4.1.2 Blunt-Notched Bend Bars

In order to apply the finite element solutions of [9] to the present

work, cleavage fracture of the specimens must occur under small scale yielding

conditions. For this reason all bend bar tests were conducted at the

relatively low temperature of —196 = C, where all specimens failed by cleavage

fracture under small scale yielding conditions.

4.1.3 Tensile Specimens

The lower yield strength, a 0 2 , and the ultimate tensile strength. cr u , were measured as functions of temperature (fig. 5). The flow properties at

various temperatures are listed in table 2.

Table 2. Flow property of the specimen material at the various test

temperatures

T

°C

°0 .

2

N/mm2 E

N/mm2 n

-90 550 210000 0.09

-85 545 210000 0.09

-70 520 210000 0.095

-60 500 210000 0.1

-55 495 210000 0.1

5
At temperatures below —178 °C all tensile specimens failed by cleavage

fracture; at higher temperatures they failed by ductile tearing.

At —196 °C cleavage fracture of the specimens occurred before the

specimen started necking, whereas at —178 °C cleavage fracture occurred in the

necked area of the specimen, but not at the minimum cross section.

4.2 Fractographic Examination

The fracture surfaces of the specimens were carefully investigated. The

objective was to find evidence for the microstructural mechanisms leading to

cleavage fracture.

At low magnification, the fracture surfaces of each type of specimen

exhibited macroscopic river patterns which originated at one small area on the

fracture surface. As shown in figure 6, these river patterns seem to be

produced by the intersection of macroscopic fracture planes with different

orientations. Fractographic examinations of the center of the macroscopic

river patterns at high magnifications using an SEM revealed fine river

patterns. In all specimens these river patterns could be traced to a single area including one cleavage facet, or, sometimes, a few. These areas are referred to as initiation sites. Four types of initiation sites were found in

the SEM examination of 26 CT specimens.

The cleavage facets at the initiation site had sizes ranging from 55 to

87 /im. In some cases an inclusion was found in the cleavage facet of the

initiation site. An example is shown in figure 7. In cases in which a facet

with an inclusion was found at the initiation site, both fracture surfaces of

the specimen were investigated in order to find out whether this inclusion was fractured. In all cases debonding between the inclusion and the matrix was observed, and the inclusions were apparently not fractured, see figure 8. The

sizes of the inclusions ranged from 2 to 8 nm. In two of the 26 CT specimens, a cluster of large inclusions was found at

the initiation site. These inclusion clusters were similar to those found in

the tensile specimen discussed in Section 4.2.2.

The energy dispersive analysis of the various inclusions show mainly the

elements Mn and S, and traces of Mn, S, and traces of Si, Ti, and Cu.

For J c -values higher than J^, the specimens had ductile tearing along the

6
crack front followed by cleavage. In figure 9, the average extension, Aa, of

ductile tearing at the crack front is shown to increase with increasing J c

. Furthermore, within the temperature range investigated, all specimens fall on

one J-R curve.

For higher J c -values (J c>Jj c ) , the initiation site was frequently found

close to a local zone of ductile tearing, Aa;g. This zone was always

significantly larger than the average of ductile tearing, Aa, along the crack

front; see figure 10. The fractographic results for the CT-specimens are summarized in table 3.

A. 2.1 Blunt-Notched Bend Bars

The fracture surfaces of all four bend bars showed cleavage fracture.

Also in these specimens one initiation site could be found in each specimen.

In all specimens a single cleavage facet was found at the initiation site. No

inclusions were found on the facets. The initiation site was always located

at a certain distance ahead of the notch tip.

4.2.2 Tensile Specimens

All tensile specimens that failed by cleavage fracture showed macroscopic

river patterns like those in the other two types of specimens; see figure 11a.

In all three specimens a cluster of large inclusions was found at the

initiation site. These clusters are similar to the clusters found in two of

the 26 CT specimens. An example of the morphology of these clusters is shown

in figure lib. Debonding between the inclusions and the matrix was observed

in all cases. At the border of the inclusion cluster a number of separate

cleavage facets, which seemed to be the initiation site, were found.

4.2.3 Location of Initiation Sites

The center of the river pattern was assumed to be the cleavage initiation

site of a specimen. For all CT-specimens and blunt notched bend bars the

location of the initiation site was determined. The distance between the

initiation site and the fatigue crack tip was measured; this distance was called r c

. In addition the distance between the centerline of the specimen

and the initiation site was measured; this distance was called z. The :• of these measurements are shown in figure 12.

7
Table 3. Fractographic results for the CT-specimens

Specimen

# [ N/mm

]

T

[°C]

Initiation

Site

r c [mm]

z [mm]

AaL

[mm]

EP 7 184 -90 Facet 0.415 1.0 .

EP 10 180 -90 Facet 0.271 2.5 -

EP 1 170 -90 Facet 0.310 2.2 -

EP 6 157 -90 Facet 0.290 5.0 - EL 11 124 -90 Facet 0.900 4.5 - EL 12 180 -85 Inclusion 0.780 0.3 - EL 4 100 -70 Facet 0.170 2.2 - EL 6 150 -70 Inclusion 0.350 3.0 - EL 3 500 -70 Inclusion 1.700 2.3 -

EP 12 60 -60 Inclusion 0.250 0.0 -

EP 8 120 -60 Cluster 1.500 6.5 -

EQ 3 124 -60 Facet 0.223 3.0 -

EP4 168 -60 Facet 0.490 0.5 -

EL 7 195 -60 Inclusion 0.750 2.1 -

EP 9 215 -60 Inclusion 1.035 2.0 -

EQ 4 286 -60 Facet 0.750 2.5 -

EP 2 304 -60 Facet 0.980 1.5 -

EQ 2 340 -60 Local Tearing 0.800 3.0 0.70

EP 14 434 -60 Local Tearing 1.030 1.2 1.00

EP 15 448 -60 Local Tearing 1.780 1.0 2.01

EP 3 513 -60 Local Tearing 1.250 0.5 1.25

EQ 1 560 -60 Local Tearing 1.800 1.5 1.40

EP 16 630 -60 Facet 1.500 1.1 -

EL 11 190 -55 Facet 0.328 5.0 -

EL 10 250 -55 Cluster 2.320 1.7 -

EL 5 445 -55 Facet 1.320 3.4 -

8
4.3 Metallographic Investigations

To find additional evidence for the cleavage crack nucleation mechanisms,

some CT and tensile specimens were sectioned perpendicular to the fracture

surface. The cross sections were then electropolished and etched. By this

technique, a number of isolated cleavage microcracks were found in the area

close (-200 fim) to the fracture surface. Most of these cracks were deflected

or arrested at grain boundaries, see figure 13. Figure 14 shows the behavior

of the carbide particles at high strains (- 10 percent) in a tensile specimen

tested at —196 °C. Some of the carbide particles have fractured but most seem

only to generate voids. Cleavage microcracks associated with broken carbide

particles were not found.

5. DISCUSSION

5.1 Fracture Stresses in the Various Specimens

Many investigations on cleavage fracture [1-7] have shown that the normal

stresses in a material seem to have a close correlation to the cleavage

fracture behavior of a material. The fact that the cleavage initiation sites can be localized in all types of specimens used in the present study provides

the opportunity to investigate this correlation in detail. For this purpose,

the normal stress at the initiation site at the moment of cleavage fracture

was calculated for all specimens by using the approaches described below. The

normal stress at the initiation site at the moment of cleavage fracture was considered to be the fracture stress of of the specimen.

5.1.1 CT-Specimens

A number of solutions are available for the calculation of the normal

stresses ahead of a stationary crack [11—15]. The solutions are based on different assumptions for shape of the crack tip or the flow property of the

material

.

It has been shown [12, 13] that for a blunted crack the normal stresshave a maximum at a certain distance ahead of the crack tip, see figure 1 In figure 15, the actual distance from the crack tip was normalized.

Therefore, the value A indicates the normalized distance. For a sharp - : , , the normal stresses at the tip become infinite when the specimen is load- :

9
For the calculation of the fracture stresses of the CT specimens,

Schwalbe's [14] solution for normal stresses was used. This solution

describes the stresses at a sharp crack, and, therefore, it does not show a stress maximum ahead of the tip. To account for the fact that cleavage

fracture of a specimen always occurred after crack blunting, the maximum of

the normal stresses was considered to be at the normalized distance Am = 2.7*10—

3

in front of the tip as shown in the finite element calculations of

[12]. The solution given in [14] was used only for distances A > Am . For

distances A < Am , that is, the region between the crack tip and the normal

stress maximum, the normal stresses were assumed to have the same value as at

the normalized distance Am . Figure 15 shows a comparison between this approach

and the solutions taken from [11—13], This approach leads to an overestimation of the normal stresses in the range A < Am , but it gives a good

estimation of the normal stresses at the distances A > Am and for the maximum

itself. It will be seen in the next section that this overestimation in the

stress calculation for the range A < Am is not essential for the results of

this paper.

Using these approaches, the fracture stresses of the CT- specimens , ag, were estimated as following:

For A > Am , that is the region ahead of the normal stress maximum:

0.3 0.04 ']n/n+l

_ A+0.1 l A J J where n is the strain hardening exponent.

For A < Am , that is the region between crack tip and normal stress maximum:

max

°f = °y =

' 0.3 f0 . 041 n/n+1

_Am+0 .

1

l Am J J

• a o . 2

The normalized distance A is defined as

& =

r c &C./2

(Kc /a o .

2

) 2

( 2 )

( 3 )

10
J c was converted to Kc by

Kc J r »E ^ 0.5

( 4 )

The crack tip opening displacement 8 C was used as defined in reference

[16] by

8 c dn vJq/(7q. (5)

For an estimation of both the strain hardening exponent n and the flow

strength oQ , the procedure proposed in [17] was used. In this procedure the

true s tress - strain curve of a uniaxial tensile test is approximated by a power

law. All data regarding the material flow property used for the calculation

of the fracture stresses are shown in table 2.

The calculations of the fracture stress of are based on the assumptions

of plane strain, small scale yielding and no crack growth. These requirements

are not satisfied for most of the tested specimens. Finite element

calculations [18] show that for side grooved specimens the center of the

specimen deforms under conditions close to plane strain up to load values

above the limit load. Reference [19] showed that the normal stress

distribution in front of the crack tip of a bend specimen can be well

estimated by the small scale yielding solution up to J- levels of about

(W—

a

Q ) o 0 _ 2 /30 . This corresponds to a J- level of about 400 N/mm for the CT

specimen tested here at —60 °C. For this J- level the average length of

ductile tearing at the crack front Aa remains smaller than half the distance

between the fatigue crack tip and the location of the maximum of the normal

stresses in front of the crack tip. These considerations show that for thos*

tests with J c of less than 400 N/mm, the assumptions used to calculate o f a in- valid and the values of af are accurate.

For J c values higher than 400 N/mm the calculated fracture stresses

more uncertain. For specimens which showed cleavage initiation at a Loc zone of ductile tearing, the assumption of an approximately stationary

is not satisfied and the fracture stress was not calculated.

11
5.1.2 Blunt-Notched Bend Bars

For the notched bend bars the fracture stresses were calculated using the

small scale yielding finite element solution given in [9], A comparison

between the f ractographic results and the normal stress distribution at the

notch shows that the initiation site is not always located at the normal

stress maximum. Thus, a slightly different procedure than that proposed in

[9] was used to calculate the fracture stresses. At first the stress

distribution at cleavage initiation was determined by an interpolation between

the loading steps given by the finite element solutions. The fracture stress

was then estimated by determining the normal stress at the location of the

initiation site in front of the tip. This procedure assures that the fracture

stress represents the actual normal stress at the initiation site, whereas the

procedure proposed in [9] would lead to a slight overestimation of the

fracture stress.

5.1.3 Tensile Specimens

For the tensile specimens, of was calculated by

of = Fc /S (6)

where Fc is the failure load of the specimen and S indicates the area of the

cleavage fracture surface.

All fracture stresses for the various specimens are shown in figure 16;

different types of specimens seem to give essentially the same fracture stress

for a particular kind of initiation site. For a single cleavage facet at the

initiation site, the fracture stresses seem to be the same for both CT

specimens and bend bars.

These two specimen types were tested at different temperatures. The

strains at the initiation site at the moment of cleavage fracture are expected

to be quite different. Therefore, the fracture stress seems to be independent

of the temperature as well as independent of the strain at the cleavage

initiation site.

Figure 16 also indicates that the fracture stresses for an inclusion

cluster are nearly the same for both the tensile and the CT specimens. The

results show that clusters of inclusions can cause a significantly lower

12
fracture stress than a single cleavage facet at the initiation site. More

experimental results are needed before final conclusions can be drawn about

the influence of inclusion clusters on cleavage fracture.

5.2 Types of Initiation Sites

The calculations of of are based on isotropic homogeneous deformation

behavior. This assumption is not satisfied for the small volume in which

initiation of cleavage seems to occur. Therefore, the fracture stresses have

to be seen as an average of the normal stresses in the volume surrounding the

initiation site. Despite this restriction, the fracture stresses may be used

for a discussion of some possible micromechanisms which may occur in the few

grains at the initiation site and which may lead to cleavage fracture.

Considering the f ractographic results, four different types of initiation

sites having different fractographic features can be classified.

i; Type A: Cleavage Facet at the Initiation Site.

This type was found mainly for J c -values smaller than Jj c as well as in

all bend bars. This type appears mainlyv at low temperatures because in all

the CT specimens as well as in all bend bars tested at —90 and at —196 °C this

type was found. The fracture stress was ''always higher than 1700 N/mm2 for

“/ A this type of initiation site. A possible . micromechanism which could explain

the fractographic observations, as well as the level of fracture stress, will

now be discussed. It may be assumed that cleavage fracture at an initiation

site takes place in a three-step process as follows:

1. Nucleation of a Cleavage Microcrack. Since the river pattern can be

traced back to an area smaller than the cleavage facet size and

smaller than the grain size, cleavage crack nucleation is a very

localized event which occurs within a grain or at a grain boundary.

The metallographic results show that small carbides can fracture and

therefore can produce a microcrack when the material is deformed,

see figure 14. The critical stress necessary to drive a carbidesized cleavage microcrack into the adjacent ferrite matrix may be

estimated by

13
74 eV d~° • 5 (l-«/ 2

)

a (7)

as proposed by Orowan [20], Using the effective fracture surface

energy 7 = 14 J/m2

, which is reasonable for a cleavage microcrack

which encounters the ferrite - carbide interface [ 21 ], and using a microcrack size of d = 0.5 /jm, which is approximately the size of

the biggest carbides, the critical stress, ac , is about 5000 N/mm2

. This stress value is much higher than the calculated fracture

stresses for any initiation site in this material. Therefore it seems unlikely that fracture of the single carbide determines the

fracture of the specimen. This conclusion is also in agreement with

the f ractographic results, which show no evidence that a fractured

particle is i Ived in the cleavage microcrack nucleation. These

findings suppc the assumption that mechanisms other than particle

fracture cause cleavage microcrack nucleation and cleavage fracture

of a specimen.

2. Cleavage Microcrack Extension. Once the cleavage microcrack has

nucleated by a mechanism other than particle fracture, the

microcrack may grow within one grain or across a few adjacent grains

which have a low mismatch in crystallographic orientation. The

stress necessary to drive a cleavage microcrack is very low for this

situation, because the effective fracture surface energy 7 for a cleavage crack in a ferrite single crystal is as low as 2 J/m2 [3,

22]. At this stage, the cleavage microcrack may form a cleavage

facet which can be observed at the initiation site by fractographic

examination

.

3. Cleavage Microcrack at a Grain Boundary. The fact that grain-sized

cleavage microcracks can be found in the material (fig. 13)

indicates that a grain boundary can be an obstacle for a cleavage

microcrack. It can be expected that the resistance of a grain

boundary increases with increasing crystallographic mismatch. This

mismatch may promote an arrest or deflection of the cleavage crack,

as shown in figure 13. Fractography shows that the borders of the

14
cleavage facets are formed by a deflection of the spreading cleavage

crack. Therefore, the borders of the cleavage facets might be grain

boundaries between grains which have a relatively high mismatch in

crystallographic orientation. For an estimation of the critical

stress necessary to drive a cleavage crack across such a grain

boundary, eq (7) may be used. The critical surface energy 7 for a cleavage microcrack encountering this kind of grain boundary may be

assumed to be within the range of 100 to 190 J/m2 as proposed in [3,

23]. Thus, the critical stress for a cleavage microcrack which has

the size of the cleavage facet (60 ^m) is in the range of 1250 to

1700 N/mm2

. This is in agreement with the fracture stresses

calculated from the present experiments; see figure 16. Therefore

it is possible that the critical step for cleavage fracture is the

ability of a cleavage microcrack to cross the first high-mismatch

grain boundaries which surround the nucleation site.

As a consequence of these cleavage initiation mechanisms, it can be assumed

that cleavage crack nucleation is necessary but not sufficient for cleavage

fracture. Only for situations in which nucleation occurs at a stress level

high enough to drive the crack across the grain boundaries does fracture of

the specimen occur. Below a certain stress level, all nucleated cleavage

cracks become arrested at the grain boundaries. This may be the reason why

all fracture stresses associated with type A initiation site are higher than

1700 N/mm2

. The size of the cleavage facets is approximately two times larger

than the average ferrite grain size. The reason for this has not been closely

investigated yet. Possibly the prior austenite grain size is related to the

facet size.

Type B: Inclusion at the Cleavage Facet is the Initiation Site.

Figure 16 indicates that the appearance of an inclusion on the cleavage - initiation site occurs mainly at higher temperatures because no bend bars or CT specimens tested at —196 or —90 °C showed this type of initiation. In

addition it seems that the appearance of an inclusion is associated with a slightly lower fracture stress.

15
Type C: Cleavage Initiation at a Cluster of Inclusions.

This type was found in all tensile specimens but only in 2 of 26 CT

specimens. The fracture stresses for this type of initiation are significantly lower than the fracture stresses for Types A and B; possibly the

cluster of inclusions causes a stress elevation in the surrounding volume and

therefore promotes cleavage fracture. The cleavage initiation mechanism

itself seems to be similar to Type A, because also in this case single

cleavage facets at the initiation sites are located close to the cluster, see figure 9b. The fact that this type of initiation site was found mainly in the

tensile specimen can be explained by the large active volume of this specimen.

This large active volume leads to a high probability of occurrence of such a cluster, whereas for a CT specimen, which has a small active volume, this

probability is very low.

Type D: Cleavage Initiation at Local Zone of Ductile Tearing.

This type becomes dominant at J c -values higher than Jxc- As shown in

Section 5.4 the length of the local zone of ductile tearing, Aap, is always

equal to the distance between the original fatigue crack tip and the normal

stress peak in front of the crack tip at the moment of cleavage fracture.

Therefore, it seems that the tearing crack grows locally ahead of the main

crack front and triggers cleavage fracture at the location of maximum normal

stress corresponding to the original stationary fatigue pre-crack.

5.3 Scatter of the Fracture Toughness

Figures 17a and b plot the J c -values versus the r c -values for the CT

specimen tested at —90 and —60 °C respectively. Two lines indicate the range

in which the normal stresses in front of the crack tip are higher than 95% of

(jy ax . This range is approximately determined by 2*10—

3

< Am < 3.5*10 3

, see figure 15.

Most of the r c -J c data points fall into this range; only a few are located slightly in front of the normal stress maximum. This result shows

that the J c -values are essentially controlled by r c , the distance between the

fatigue crack tip and the cleavage initiation site. This distance, r c , characterizes how far the normal stress peak must move into the ligament in

order to find the cleavage initiation site. Comparing the r c -values for

16
-60 °C with those -90 °C, it can be seen that the scatter of rc decreases

significantly with decreasing temperature, and so does the scatter of J c

. This shows that the scatter of the toughness values, J c , is caused mainly by

the scatter of the location of the cleavage initiation sites in the ligament.

The reduction of the scatter of the r c -values with lower temperature

indicates that the density of the cleavage initiation sites at the crack tip

increases with lower temperature, possibly because of the increase of the

normal stresses due to the temperature dependence of the yield strength.

These higher normal stresses at the crack tip may give more cleavage cracks

the ability to spread out than at higher temperatures. The change in

temperature from —60 to —90 °C causes an increase of the normal stresses of

about 10 percent. This change is relatively small. It is hard to believe

that this small stress elevation is the only reason for the large change in

fracture behavior. Possibly the temperature influences more than just the

stress elevation. However, the small variation of normal stresses is the only

clear reason for the variation of the density of the initiation sites, and as

a result, the variation in fracture behavior of the CT specimen. If this were

so, then any kind of small variations in yield strength would lead to large

changes in brittle fracture behavior.

5.4 Some Aspects on the Lower Bound Fracture Toughness and the Transition

Temperature

The approach discussed in Section 5.3 does not explain the existence of a lower bound fracture toughness because, in principle, a cleavage initiation

site may be located very close to the crack tip. Considering the correlation

of J c with rc , this situation would lead to a vanishing J c value as r c approaches zero. In all the tests conducted in this study, no r c values smaller than 170

^m have been found. This indicates that the initiation of cleavage close to

the crack tip is very unlikely.

An interpretation of this result can be made by the RKR-Model [24] which

postulates that cleavage can only occur if the fracture stress is reached or exceeded over a critical distance, called X. According to the normal st r< behavior at the crack tip, this requirement can only be satisfied when .! i

17
high enough to provide a normal stress peak sufficiently wide enough to cover

the initiation site.

The lower bound toughness would then be characterized by a situation in

which the critical length, X, which covers an initiation site, equals the

width of the normal stress peak at the stress level necessary to initiate

cleavage fracture.

Results of RKR calculations are shown in figure 4 for critical distances

X = 100 and 200 /im using the fracture stress found for the CT specimens for

the type A initiation site. It is not clear yet whether this model predicts

the lower bound toughness of the specimen because there are not enough

experimental data for small r c values. Recent results [25] indicate that the

minimum of the r c values, that is r cm i n , decreases with decreasing

temperature; see figure 18. This is in agreement with the approach described

above because with decreasing temperature the width of the normal stress peak

at a given stress level increases due to the increase of the yield strength.

Therefore, the requirement of reaching or exceeding the fracture stress over

the critical distance X can be satisfied at lower J- levels for which the

normal stress peak is close to the crack tip.

Cleavage fracture will occur at an initiation site if the normal stress

exceeds a certain stress level, called the activation stress. Scatter results

because cleavage occurs when the normal stress peak exceeds the lowest

activation stress of all initiation sites ahead of the crack tip. If the

activation stress were never exceeded, the specimen would fail by ductile

fracture. The normal stress peak is a function of the yield strength and work

hardening of the steel at the temperature and constraint level existing at the

initiation site.

For a discussion of the activation stress, the fracture stresses of the

CT specimens may be used. Figure 16 shows that the majority of the type-A

initiation sites have a fracture stress higher than = 1720 N/mm2

; only 2 of

26 specimens had a lower fracture stress. A lower fracture stress oq of 1500

N/mm2 was found to be associated with initiation sites of type C. The fact

that only 2 of 26 specimens exhibited type C initiation sites indicates that

the density of this type of initiation site is low compared to the density of

initiation sites which become activated at stress levels higher than a For

a situation in which the normal stress peak has a stress level lower then a

18
but higher than oq, it can be expected that only the type C initiation sites

could be activated. Due to the low density of type C initiation sites, the

r c -values for this situation will show a large scatter, and therefore J c will

show a large scatter. For a situation in which the normal stress peak becomes

lower then oq , cleavage fracture of a specimen becomes impossible.

If the normal stress peak in the CT-specimen can be estimated by eq (2),

then the temperatures at which the stress peak has a level of ag and oq

respectively can be calculated using the temperature versus yield strength

relationship shown in figure 5:

a . = a0.22.921 (14.81) n/(n + 1) (8a)

c = o0.22.921 (14.81) n/(n + 1)

This estimation gives temperatures of = —30 °C and > room temperature.

At the present stage of this study it is not possible to evaluate this

approach because no data are available yet for this temperature range. But

figure 4 seems to indicate that the scatter of J c becomes very large for

T > —30 °C as predicted by this estimate.

At the present state of this study there are no experimental data

available for specimens thicker than B = 20 mm, but the present data suggest .. few points concerning thickness effects and the transition behavior which ir.av be checked in a further study.

With increasing thickness, the number of cleavage initiation sites vi'l.i: a certain distance from the crack tip will increase because of the incr. . , length of the crack front. This can lead to a reduction of the rc -valu< 1 therefore to a reduction in the scatter of the J c -values, as observed • u experimental studies.

19
Figure 19 indicates that cleavage initiation occurs mainly in the center

region of the specimen. The effective thickness area in which cleavage can occur may be called B e ff; B e fg decreases with increasing distance from the

crack tip. The initiation sites located far from the crack tip are related to

higher J c -values than at initiation sites located close to the tip (see r c -J c relationship in figure 17) . Finite element solutions show that with

increasing J the loss of triaxiality of the stresses grows gradually from the

side surface of the specimen into the specimen. This loss of triaxiality

leads to lower normal stresses. Therefore, the location of the cleavage

initiation sites tends to be located nearer to the center of the specimen with

increasing distance from the crack tip. The J-value at which Be ff approaches

zero seems to characterize the situation at which the normal stress peak

becomes lower than adue to the loss of triaxiality of the stresses.

6. CONCLUSIONS

One cleavage initiation site was found on the fracture surface of each

specimen, including fatigue cracked CT- specimens , blunt notched bend bars, and

tensile specimens.

Four different types of cleavage initiation sites were classified:

Type A: Cleavage facet at the initiation site

Type B: Nonfractured inclusion at the initiation site

Type C: Cluster of inclusions at the initiation site

Type D: Initiation at a local zone of ductile tearing

In the CT- specimens , cleavage crack initiation occurs at the location of

the maximum normal stresses ahead of the crack tip. The scatter of fracture

toughness in the upper transition region is due to the scatter of the distance

between the fatigue crack tip and the cleavage initiation site.

7 . ACKNOWLEDGMENTS

The first author would like to thank the GKSS - Forschungszentrum

,

Geesthacht, West Germany and the National Institute of Standards and

Technology, Boulder, for supporting his sabbatical at NIST. The authors

20
gratefully acknowledge many helpful discussions with H.I. McHenry. We would

also like to thank P.T. Purtscher and Y. -W. Cheng for their contributions and

J.D. McColskey for technical support.

8 . NOMENCLATURE

A normalized distance from the crack tip.

a crack length

a 0 crack length after precracking

Aa average of ductile tearing

Aag local amount of ductile tearing

B specimen thickness

Bn net thickness

d size of a cleavage microcrack

dn factor which correlates J c with 8 C [16]

6 crack tip opening displacement

6 C crack tip opening displacement at the moment of cleavage fracture

E Young's modulus

Fc Load at the moment of cleavage fracture

7 effective fracture surface energy for a moving cleavage crack

J J- integral

J c J- integral at the moment of cleavage fracture

Ji J -value at which first dimpled rupture can be observed

Jj c fracture toughness after [8]

Kfflax maximum stress intensity factor used in precracking

n strain hardening exponent

Pmax maximum load during precracking

^min minimum load during precracking

r c distance between fatigue crack tip and cleavage initiation site

r cmin smallest rc -value observed at a given temperature

S area of the cleavage fracture surface

T temperature

T'p transition temperature

W width of the specimen

X critical distance for cleavage initiation

21
o 0 2 lower yield strength

au ultimate tensile strength

lowest fracture stress for initiation Type A of the CT-specimen

oq lowest fracture stress for initiation Type C of the CT-specimen

lowest fracture stress of all initiation sites at the crack tip

ITlclX "\ t r- t i i i a__ stress level of the normal stress peak at the crack tip m a CTCT specimen r max . . .. max , . <7^ an estimation ot a^

by using equation 2.

<7f fracture stress

9 . REFERENCES

[1] J. Merkle; An Examination of Size Effects and Data Scatter Observed in

Small Specimen Cleavage Fracture Toughness Testing, NUREG/CR- 3672

,

ORNL/TM- 9088 , (1984).

[2] R.O. Ritchie; A.W. Thompson; On Macroscopic and Microscopic Analysis for

Crack Initiation and Crack Growth Toughness in Ductile Alloys, Metall

,

Trans

.

. Vol. 15A, (1985), pp . 233-248.

[3] G.T. Hahn; The Influence on Microstructure on Brittle Fracture Toughness,

Metall. Trans. Vol. 15A, (1984), pp . 947—959.

[4] C.J. McMahon, M. Cohen; Acta Metall

.

. Vol. 3, (1965), pp . 591—604.

[5] P. Brozzo, G. Buzzichelli, A. Mascanzoni, M. Mirabile; Microstructure and

Cleavage Resistance of Low Carbon Steel, Metal Science . Vol. 11, (1977),

pp. 123-129.

[6] G. Irwin, Brittle-Ductile Transition Behavior in Reactor Vessel Steels,

WRSI Meeting, NBS -Gaithersburg , Oct. (1986).

[7] D.A. Curry; Influence of Microstructure on Yield Stress and Cleavage

Fracture Stress at -196 °C of SA508 Class 2 Pressure Vessel Steel, Metal

Science . Vol. 18 (1984), pp . 67—76.

[8] ASTM E813-81, Standard Method for Jjc> A Measure of Fracture Toughness,

ASTM- Standards , Part 10, Metals, (1982).

[9] J.R. Griffiths; Dr. J. Owen, Journal of Mech. Phvs . Solids . Vol. 19

(1971), pp. 491-431.

[10] J. Heerens, K.H. Schwalbe, A. Cornec ; Modification of ASTM E 813 Standard

Test Method for an Improved Definition of Jjq Using a New Blunting Line

Equation. Fracture Mechanics—18th Symposium . ASTM STP 945, edited by

R.P. Reed and D.T. Read, American Society for Testing and Materials,

Philadelphia, 1988, pp . 374—389.

22
[11]

D.M. Tracy; Finite Element Solution for Crack Tip Behavior in Small Scale

Yielding, J. Eng. Maths. Tech. . Vol. 98, (1976), pp . 146—151.

[12] J.R. Rice, M.A. Johnson; Inelastic Behavior of Solids , eds . M.F. Kanninen

et al., McGraw-Hill, (1970), pp . 641—672.

[13] R.M. McMeeking; Finite Deformation Analysis of Crack Tip Opening in

Elastic-Plastic Materials and Implication for Fracture. J. Mech. Solids . Vol. 25 (1977), pp. 375-381.

[14] K.H. Schwalbe; Comments to Finite Element Solution of Crack-Tip Behavior

in Small Scale Yielding, Transactions of the ASME . April (1977), pp

.

186-188.

[15] J.R. Rice, C.F. Rosengren; J. Mech. Phys . Solids . Vol. 16 (1968), pp

.

1 -12 .

[16] C.F. Shih; Relationship Between the J-Integral and the Crack Tip Opening

for Stationary and Extending Cracks, J. Mech. Phys. Solids . Vol. 29,

(1981), pp. 305-326.

[17] A. Cornec, J. Heerens , K.H. Schwalbe, Bestimmung der RiBaufweitung CTOD

und der Rissabs tumpfung SZW und dem J- Integral, GKSS 86E15, (1985).

[18] H.G. DeLorenzi, F. Shih; 3D Elastic-Plastic Investigation of Fracture

Parameters in Side Grooved Compact Specimens, Int . Journal of Fracture . Vol. 21, (1983), pp. 195-220.

[19] R.M. McMeeking, D.M. Parks; On the Criteria for J-Dominance of Crack-Tip

Fields in Large Scale Yielding, Elast . -Plast . Fracture. ASTM STP 668 . (1979), pp. 175-194.

[20] E. Orowan, Rep. Progress Physics . Vol. 12, (1948), p. 185.

[21] D.A. Curry, J.F. Knott; Effects of Microstructure on Cleavage Fracture

Stress in Steel, Metal Science . Vol. 12, (1978), pp . 511—514.

[22] F.C. Frank, A.W. Stroh, Proc . Phys . Soc

.

. Vol. 65B, (1952), p. 811.

[23] G.T. Hahn, B.L. Auerbach, W.S. Owen, M. Cohen; Fracture . B.L. Averbach et

al., eds. John Wiley and Sons, New York, (1959), p. 91.

[24] R.O. Ritchie, J.F. Knott, J.F. Rice; On the Relationship Between Critical

Tensile Stress and Fracture Toughness in Mild Steels. J . Mech , Phys

.

Solids . Vol. 21, (1973), pp . 395-410.

[25] J. Watanabe, T. Iwadate , Y. Tanaka, T. Yokobori, K. Ando; Fracture

Toughness in the Transition Region, to be published in J . of Frac . M< Nov. 1987.

23
Figure 1 Heat treatment of the forged material.
Figure 2. Microstructure of the material.

a. cluster of large inclusions

b. small carbides are arranged in rows within the ferrite grain

25
,

N/mm

Figure 3. CT- specimens , blunt notched bend bars and tensile specimens as used

in this study.

T, °C

Figure 4. Fracture toughness of the material in the upper transition region.

26
2

ujuj/n

‘

bf

Csl

b°'

T, °C

Figure 5. Ultimate tensile strength and yield strength versus temperature

27
20 mm

Figure 6. Fracture surface of a CT-specimen showing macroscopic river pattern

which emanates from one small area on the fracture surface in front

of the fatigue crack tip.

Figure 7. Cleavage fracture surface showing fine river pattern. A single

cleavage facet is located at the center of the river pattern.

28
Figure 8. A nonfractured inclusion at the center of the river pattern. Both

fracture surfaces of a CT-specimen are shown.

29
N/mm

A a, mm

Figure 9. J-Aa curve of the CT-specimens within the temperature range of

—55 to —85 °C. Each data point indicates one specimen. The

blunting line was calculated as proposed in [17].

30
Figure 10. Cleavage fracture initiation at a local zone of ductile tearing.

Two different CT-specimens are shown.

31
1mm 20 Ofxm

Figure 11a. Macroscopic river pattern of a tensile specimen, tested at

-196 °C.

30 m

Figure lib. Cluster of large inclusions in the center of the river pattern; a number of single cleavage initiation sites form the center of the

fine pattern at the border of the cluster.

32
E

E

a* L-°

Z, mm

Figure 12. Location of the cleavage initiation site in the ligament,

data point indicates one test.)

(Each

33
Figure 13. Cleavage microcracks which become arrested or deflected at the

grain boundaries. (The micrograph was taken at a distance of 200

to 300 /xm ahead of the fatigue crack tip.)

lum

Figure 14. Fractured carbide particles in a tensile specimen tested at

—196 °C. The plastic strain is about 10 to 15 percent.

34
Figure 15. Normal stresses in front of a blunted crack tip.

[14] was used in this study.

The solution ir

35
gUJiii/N

2500

2000

1500 -

1000

Bend Bars{

o o

o

o

Tensile

{

_L

CT-Specimens

OOOo

®

® OOOO

o

"O-

_L

o Type A

® Type B • Type C

I

&

oo

®

°A

-206 “196 “90 -80 -70 -60 -50

r, °c

Figure 16. Fracture stresses of the various specimens for the various types of

initiation sites.

36
N/mm

Figure 17. Relationship between the location of the initiation sites and tin- value of fracture toughness. Lines indicate the approximate

location of the normal stress peak in front of the crack tip.

37
,

N/mm

Figure 18. The rm^ n

-value for different temperatures (data were taken from

[25]).

38
m

m

1000

800

E

E

eg

o\

^max _

^CT <crA

o

\

o

Beff/2

o

\

Ox

0

>

o

o

k_

O

0

o

CO

6 0 0 04

400

o

rc min

0

°o °

O O

Xex

o

^

Cr>^A

O O

o

O- o

O O

o

200

II

10

Z, mm

Figure 19. Relationship between effective thickness B e ff and J-integral, as indicated by the locations of the cleavage initiation sites found

in this study.

39

U. S. GOVERNMENT PRINTING OFFICE: 1989 -674-456/ 5029

(rc

-

8/2)(1-
NBS-1 14A i re v . 2-ac > U.S. DEPT. OF COMM. 1. PUBLICATION OR

REPORT NO.

2. Performing Organ. Report No. 3. Publication D ate BIBLIOGRAPHIC DATA

SHEET (See instructions) NISTIR 88-3099 December 1988

4. TITLE AND SUBTITLE

Fracture Behavior of a Pressure Vessel Steel in the Ductile to

Brittle Transition Region

5. AUTHOR(S)

J. Heerens, D.T. Read

6. PERFORMING ORGANIZATION (If joint or other than NBS, see i n stru ction s) National Institute of Standards and Technology

W A T KNHH-- B tt RE AH -04=- $ T A W B Afl B5 DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

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8. Type of Report & Period Covered

9. SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (Street. City, State. ZIP)

10. SUPPLEMENTARY NOTES

Document describes a computer program; SF-185, FIPS Software Summary, is attached.

11. ABSTRACT (A 200-word or less factual summary of most significant information. If document includes a significant

bibliography or literature survey, mention it here)

The reasons for the scatter of fracture toughness in the duct ile-to-br ittle

transition region, as well as the mechanisms leading to cleavage fracture, have

been investigated for a quenched and tempered pressure vessel steel, DIN 20

MnMoNi 55. The fracture surfaces indicate that cleavage fracture starts at one small area in the ligament, the cleavage initiation site. Cleavage

initiation occurs ahead of the crack tip at the location of the maximum normal

stresses. Fractography and metallography show four different types of initiation

sites. The mechanisms which may trigger cleavage fracture at these initiation

sites are discussed. The results indicate that the scatter of fracture

toughness is due to the scatter in the distance between the cleavage initiation

site and the fatigue crack tip.

12. KEY WORDS (Six to twelve entries ; alphabetical order; capitalize only proper names; and separate key words by semicolons)

cleavage; fracture stress; inclusions; initiation site; mechanism; origin;

scatter; weak link

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