-
Accepted 8 January 2013rolled 2205 duplex stainless steel was
performed. The results show that the splits cause
es fo
rowth direction.materials atesting dueifferent cry
graphic orientations or banding of weaker phases. Two examples
of the former case is the splits in a ferritic stainlewhich was
found to be due to texture banding from the rolling procedure [9]
and the splits found in a pipeline steelrolling at low nishing
temperature resulted in a texture that promoted cleavage fracture
parallel to the plane of rolling [10].
0013-7944/$ - see front matter 2013 Elsevier Ltd. All rights
reserved.
Corresponding author. Tel.: +46 8 7906252.E-mail address:
[email protected] (J. Pilhagen).
Engineering Fracture Mechanics 99 (2013) 239250
Contents lists available at SciVerse ScienceDirect
Engineering Fracture Mechanics
journal homepage: www.elsevier .com/locate
/engfracmechhttp://dx.doi.org/10.1016/j.engfracmech.2013.01.002parallel
to the crack growth direction while a TS specimen will have splits
perpendicular to the crack gDelamination of the fracture surface
during fracture toughness testing has been observed for a range
of
mechanism has been explained by the existence of weak planes in
the material which delaminate underthrough-thickness stress [8].
The explanations for these weak planes have either been the
interaction of dnd theto thestallo-ss steelwhere[15]. A common
feature is secondary cracks, so called splits, which are
delamination of the material growing normal tothe plane of the
fatigue crack. They are often seen after fracture toughness and
impact testing when tested through-thicknessalong the rolling
direction (TL) [16]. Their size, numbers and proximity to the
fatigue crack increase with decreasing tem-perature [7]. It has
been proposed that splits affect the scatter in fracture toughness
measurements [3].
The splits are assumed to occur in the ferritic phase or at
austenite/ferrite phase boundaries and the orientation of thesplits
is governed by the orientation of the fracture plane in relation to
the microstructure. A TL specimen will have splitsKeywords:Duplex
stainless steelFracture
toughnessDelaminationSplitsPop-inNormalizationMaster curve
1. Introduction
Reported fracture toughness valuloss of constraint along the
crack front. This can be observed as local difference in
crackgrowth in the specimen. The initiation fracture toughness is
not inuenced by the speci-men thickness. Furthermore, due to the
delamination the material exhibits a stable fractureprocess despite
the presence of cleavage fracture. This is interfering with the
master curvemethod so for evaluating the fracture toughness at
sub-zero temperatures an assessmentof the fracture resistance curve
is instead suggested. For assessing the brittle crack behav-iour at
sub-zero temperatures it is proposed to use the split initiation as
a failure criteria.The splits are also the cause of the pop-in
behaviour observed for the duplex stainlesssteels. The
susceptibility for pop-in is inuenced by the microstructure.
2013 Elsevier Ltd. All rights reserved.
r duplex stainless steels (DSSs) demonstrate high toughness at
low temperaturesLoss of constraint during fracture toughness
testing of duplexstainless steels
Johan Pilhagen , Rolf SandstrmMaterials Science and Engineering,
KTH, Brinellvgen 23, S-10044 Stockholm, Sweden
a r t i c l e i n f o
Article history:Received 2 December 2011Received in revised form
10 September2012
a b s t r a c t
Delamination of the fracture surfaces, so called splits, is an
important phenomenon thatoccurs at sub-zero temperature for
hot-rolled duplex stainless steels during impact andfracture
toughness testing. To evaluate how the splits inuence the fracture
toughness,sub-zero temperature fracture toughness testing of 50, 30
and 10 mm thick plates of hot
userLinien
userLinien
userLinien
userRechteck
userRechteck
userRechteck
userLinien
userLinien
-
Nomenclature
Ap plastic workb0 initial ligament lengthB0 plate thickness in
Fig. 3 or specimen thickness in Eq. (4)B1T reference thicknessBN
net thickness of the specimenC1 coefcient from the power law
regression of the JDa dataC2 coefcient from the power law
regression of the JDa dataE Youngs modulusJe elastic component of
the J-integralJIc initiation fracture toughnessJpl plastic
component of the J-integralJQ tentative initiation fracture
toughnessK stress intensity factor
240 J. Pilhagen, R. Sandstrm / Engineering Fracture Mechanics 99
(2013) 239250KJc(0) measured fracture toughness at the point of
instabilityKJc(1T) size adjusted fracture toughnessKmin threshold
value for cleavage fracturem constraint parameter, function of
crack depth and strain hardening exponentM non-dimensional
deformation limitT0 reference temperatured crack-tip opening
displacement (CTOD)Da crack extensiong dimensionless parametert
Poissons ratiorys yield strength at the test temperaturerY
effective yield strength at the test temperatureCMOD Crack Mouth
Opening DisplacementIn the case of banding of weaker phases there
is an example of another pipeline steel where the
thermomechanically processcaused thin layers of martensitic or
bainitic grains to occur through the thickness parallel to the
rolling direction [11].
In the work by Nilsson [6] the impact toughness of 50 mm and 12
mm plates of 2205 duplex stainless steel were testedfrom room
temperature down to liquid nitrogen temperature for all the major
orientations in Fig. 1. The conclusions werethat the specimens are
toughest when the notch is oriented in the normal plane (LS, TS and
45-S) followed by the casewhen the notch is oriented perpendicular
to the normal plane (LT, TL and 4545). The lowest toughness is
reached whenthe notch is oriented parallel to the normal plane (SL,
ST and S-45) which is the same plane as for the splits. It is
interestingto note that the 12 mm plate has higher toughness than
the 50 mm plate in all orientations except for the
last-mentionedwhere the 12 mm plate transition region has moved to
higher temperature and lower upper-shelf energy compared tothe 50
mm plate. A texture characterisation showed that the ferrite in the
12 mm plate had strong texture while the ferritein the 50 mm plate
was more prone to exhibit random texture. However, no unambiguous
effect of the crystallographic ori-entation could be related to
anisotropy of the impact toughness [6].
DSS duplex stainless steelsLOM light optical microscopeSEM
Scanning Electron Microscope
Fig. 1. Schematic drawing of specimen orientations. L is the
rolling direction, T is the transversal direction and S is the
through-thickness direction.
-
The objective of this paper is to analyse how the splits inuence
the fracture toughness and how the microstructure af-fects the
splitting phenomenon in hot-rolled duplex stainless steels. The
approach was to test single-edge notched bend bars,SE(B), with
different dimensions from a 50 mm plate and compare these with 30
mm plate and 10 mm plate.
2. Material and testing procedure
2.1. Material
The material used in this work was a commercially produced
duplex stainless steel designated 2205 (EN 1.4462, UNSS32205),
produced by Outokumpu Stainless. The material was hot-rolled to the
desired plate thickness (50 mm, 30 mmand 10 mm) followed by
solution treating at 1100 C and water quenching. The chemical
composition can be found inTable 1.
The 50 and 30 mm plate have almost the same tensile properties
at room temperature with yield strength of 496 MPa and492 MPa and
tensile strength of 739 MPa and 737 MPa respectively. For the 10 mm
plate the yield strength is 604 MPa andthe tensile strength is 804
MPa at room temperature. The values for yield and tensile strength
at the testing temperatures forthe 30 and 50 mm plate were obtained
from a cubic polynomial t of the yield and tensile strength data
found in the Euro-pean research project EcoPress [12]. For the 10
mm plate the difference in yield and tensile strength at room
temperaturebetween the 10 mm plate and the polynomial t at room
temperature was assumed to be constant.
Table 1Chemical composition (wt.%) of the plates.
Plate thickness (mm) C Si Mn P S Cr Ni Mo N
50 0.023 0.40 1.40 0.020 0.001 22.4 5.70 3.20 0.1730 0.022 0.44
1.47 0.026 0.001 22.4 5.61 3.12 0.1710 0.016 0.50 1.40 0.023 0.001
22.4 5.20 2.80 0.18
J. Pilhagen, R. Sandstrm / Engineering Fracture Mechanics 99
(2013) 239250 241Fig. 2. Microstructure of the 10 and 50 mm plate:
(a) 10 mm plate and (b) 50 mm plate. Ferrite is the dark phase and
the rolling direction is horizontal withthe gure.
-
Polished and etched light optical microscope (LOM) photos of the
microstructure (transversal plane) for the 10 and50 mm plate are
shown in Fig. 2. It is evident that the 10 mm plate has a more
elongated and aligned austenite lamella struc-ture than the 50 mm
plate due to the more extensive hot-rolling. The austenite spacing
and thickness were measured per-pendicular to the rolling direction
at 20magnication in LOM, see Table 2. The measurements were in the
transversal planeaccording to Fig. 1. In this work, the rolling
direction (L) is normal to the L-plane, the transverse direction
(T) is normal to theT-plane and the through-thickness direction (S)
is normal to the S-plane. The 10 mm plate has the smallest distance
betweenthe austenite lamellae (austenite spacing) followed by the
30 mm plate. The 50 mm plate has a homogenous microstructurewith
respect to thickness of the austenite lamellae and their
spacing.
2.2. Specimen selection
The specimen orientation for all tested specimens was TL
according to Fig. 1 and the specimen dimensions were50 50 400 mm3,
30 60 400 mm3 and 10 20 200 mm3. The rst number corresponds to the
plate thickness. Froma 50 mm plate, 10 mm specimens of dimension 10
20 200 mm3 and 30 mm specimens of dimension 30 60 400 mm3were cut
out. Three types of 10 mm specimens were cut out according to their
location in the thickness direction in the50 mm plate, surface
specimens, intermediate specimens and mid-thickness specimens. The
methods used were water cut-ting or sawing. From the 30 mm plate,
specimens of dimension 30 60 400 mm3 were produced. All surfaces
were ma-chined to comply with the fracture toughness standard ASTM
E 1290-02 [13]. Electrical discharge machining was usedfor the
notch and clip-gauge slots.
2.3. Testing procedure
Standard SE(B) specimens were used in the fracture toughness
measurements. Side-grooves were used on all fracturetoughness
specimens and the crack length over the specimen width ratio was
around 0.55 including fatigue precrack. Thefracture mechanic
testing was done with a 100 kN hydraulic testing machine and a
clip-gauge to measure the crack mouthopening displacement (CMOD) of
the specimen. The specimens were submerged during the testing in
ethanol. Liquid nitro-gen was used to cool the ethanol down to the
testing temperature.
242 J. Pilhagen, R. Sandstrm / Engineering Fracture Mechanics 99
(2013) 239250The fracture toughness testing was based on ASTM E
1290-02 for evaluating the crack-tip opening displacement
(CTOD)[13]. The CTOD can be evaluated from the sum of the elastic
(Jel) and plastic (Jpl) component of the J-integral at the point
ofinstability. The point of instability can either be the unset of
unstable crack extension prior to or following stable crackgrowth
or the attainment of a maximum force plateau [13]. The elastic
component can be found from the relation betweenthe stress
intensity factor K and the Jel:
Table 2Mean austenite thickness and spacing.
Plate thickness (mm) Plane Austenite thickness (lm) Austenite
spacing (lm)
10 T 5.8 5.530 T 9.6 6.950 Ca T 11.5 9.250 Sb T 11.6 9.8
a C = mid-thickness specimen.b S = surface specimen.
100 90 80 70 60 500
50
100
150
200
250
300
350
400
Temperature [C]
m [
m]
50 mm30 mm, B0 = 50 mm10 mm, B0 = 50 mm
Fig. 3. Result from the fracture toughness testing of the 50 mm
plate. B0 is the plate thickness.
-
Jel K21 t2
E1
where t = 0.3 is the Poissons ratio for duplex stainless steel,
211 GPa is used for the elastic modulus, E. K is the elastic
stressintensity factor at the point of instability. The plastic
component, Jpl, can be found from the point of instability as
follows:
Jpl gplApBNb0
2
where Ap is plastic area under the force versus crack mouth
opening displacement records, gpl = 3.7853.101(a/W) + 2.018(a/W)2,
BN is the net thickness and b0 is the remaining ligament. The
crack-tip opening displacement can be found from Eq. (3):
d Jel JplmrY
3
where m is a constraint parameter which is a function of the
crack depth and the strain-hardening exponent. rY is the effec-tive
yield strength at test temperature, dened as the mean of yield and
tensile strength.
It has been reported in the literature that pop-ins are quite
common during fracture toughness measurements in rolledduplex
stainless steels [14]. Pop-in is a sudden crack extension with
accompanied crack arrest. For a pop-in to be regardedas signicant
and taken as a point of instability it must be severe enough. Two
common criteria are the increased crack
10 mm plate in the tested temperature interval.
200
Popin
Fig. 4.the sam
J. Pilhagen, R. Sandstrm / Engineering Fracture Mechanics 99
(2013) 239250 243Result from the fracture toughness testing of the
30 mm plate. For the lled data points the evaluated maximum force
plateau and pop-in are frome specimen.100 90 80 70 60 500
Temperature [C]50
100
Maximum force plateau150
[250
m]The difference in pop-in behaviour between the 30 and 10 mm
plate is that for the 10 mm plate the pop-in occurs shortlyafter
the start of the nonlinear behaviour while for the 30 mm plate the
pop-in occurs more closely to the maximum force
300
350
400length [13,14] or the force drop in combination with
increased clip-gauge displacement [13]. In this work a pop-in is
eval-uated according to Section 9.1.2 in ASTM E 1290-02 [13].
3. Results
No unstable failure occurred during the testing and thus the
failure criteria for the tested specimens were chosen to bemaximum
force plateau or pop-in. For the rst case when the maximum force
plateau was reached the crack extension wasmeasured by subsequent
unloading. Post-fatigue cracking at room temperature was also
conducted for measuring the crackextension with LOM after breaking
the specimen. Pop-ins were only observed for the 10 and 30 mm
plate. The testing tem-peratures were between 94 C and 18 C.
The result for the 50 mm plate can be seen in Fig. 3. The 10 mm
specimens have in principle the same measured CTOD asthe 50 mm
specimens and shows that the 50 mm plate has homogenous fracture
toughness as the specimens taken fromdifferent depths of the plate
exhibit similar toughness and scatter. The 30 mm specimens seem to
have a bit higher CTOD.The result shows that the CTOD dened from
the maximum force plateau gives temperature independent fracture
toughnessin the tested temperature interval.
The result from the 30 mm plate is shown in Fig. 4. All the
specimens from the 30 mm plate exhibited multiple pop-insbut only
two specimens had signicant pop-ins as shown in Fig. 4. The 30 mm
plate has higher CTOD than the 50 mm plate.
For the 10 mm plate multiple pop-ins occurred during the testing
and if the maximum force plateau were used instead ofpop-in the
CTOD would increase substantially, see Fig. 5. The fracture
toughness seems to be temperature independent for
-
plateau, compare Fig. 4 with Fig. 5. According to the method
used for pop-in evaluation, the pop-in for the 10 mm plate isalso
more severe compared to the 30 mm plate.
The fracture surfaces of all specimens are dominated by the
splits and extensive shear-lip deformation between the splits.This
is independent of plate thickness and specimen dimension. Their
sizes vary from small ones only observable in SEM tolarger ones
that are visible to the naked eye. These splits grow both in and
perpendicular to the plane of the fatigue crack as
100 90 80 70 60 50 40 30 20 100
50
100
150
200
250
300
350
400
Temperature [C]
[
m]
Maximum force plateauPopin
Fig. 5. Result from the fracture toughness testing of the 10 mm
plate. For the lled data points the evaluated maximum force plateau
and pop-in are fromthe same specimen.
244 J. Pilhagen, R. Sandstrm / Engineering Fracture Mechanics 99
(2013) 239250seen in Fig. 6, and can be several mm deep. The true
length of these splits during testing is unknown because they grow
dur-ing subsequent post-fatigue (30 < KI < 70 MPa
pm).
Examination of the fracture surfaces shows a distinct difference
between the 10 and 30 mm plate and the correspondingspecimens from
the 50 mm plate. The 10 and 30 mm plate have numerous and clearly
dened splits, see Fig. 6a, compared tothe 10 and 30 mm specimens
from the 50 mm plate which have fewer, shallower and less dened
splits, see Fig. 6b. The50 mm specimens have similar splits as the
30 mm plate, but not as deep and well dened.
The fracture surface between these splits is a combination of
brittle fracture and microvoid coalescence (MVC). Close tothe
splits the MVC fracture process is dominating. The brittle
fractures consist of cleavage fractures in the plane of the
fatiguepre-crack, see Fig. 7, which are arrested locally. There are
also areas mixed with the MVC that has curved facets and no
direc-tion of crack advancement which indicates a quasi-cleavage
type of fracture [15]. The amount of growth between the
splitsincreases with increasing distance between the splits.Fig. 6.
SEM photos of the fracture surface for 30 mm specimens
(25magnication): (a) 30 mm plate and (b) 30 mm specimen from the 50
mm plate. Blackdots indicate the boundary between crack growth and
post fatigue cracking.
-
45 specimens) the cleavage crack is free to propagate in the
plane of the notch with little resistance from the austenite.
Thiscauses the appearance of a weaker plane. The lower impact
toughness of the S-plane for the 12 mm plate compared to the
J. Pilhagen, R. Sandstrm / Engineering Fracture Mechanics 99
(2013) 239250 24550 mm plate in the work of Nilsson [6] might be
explained by the difference in texture. The increased splitting in
the 10 and30 mm plate compared to the cut out specimens from the 50
mm plate is then explained by the more elongated and
nermicrostructure found in the 10 and 30 mm plate, recall Fig. 2
and Table 2.
4.2. Loss of constraint
In an article by Embury et al. [16] the splits inuence on impact
toughness was evaluated by the introduction of weakerindividual
laminate subunits in the specimen (mild steel plates soldered
together). It was concluded that the transition re-gion was
markedly shifted to a lower temperature and this effect increased
with increasing laminate subunits. The increasedtoughness was
explained with the impact specimens behaving as the sum of the
individual subunits which caused a relax-ation of triaxial
stress.
In the work by Chan [17] the same behaviour was observed for
AlFeX alloys during fracture mechanics testing. It wasfound that
the specimens exhibited similar splits as the tested duplex
stainless steels and this caused measured KIc values to4.
Discussion
4.1. Cause of splits
For the reported cases of splits in various materials in the
literature, the splits were single or few and found in the
mid-thickness region where the triaxial stress is the highest. This
is not the case for the tested DSS where the splits were numer-ous
and seemed to be evenly distributed over the thickness (at least
for the splits clearly seen by the naked eye).
The likely explanation for the splits in 2205 is that when
cleavage crack initiation occurs in the ferrite the cleavage
prop-agation is locked along the austenit lamellae which cause
delamination. For crack propagation in the S-plane (SL, ST and
S-
Fig. 7. SEM photo showing cleavage fracture. 30 mm plate
specimen tested at 55 C.deviate from theoretical plane-strain
stress state toward a plane-stress stress state. This mechanism was
called thin sheettoughening by Chan and it was later conrmed that
for delaminated AlFeX alloys the KIc fracture toughness is
controlledby the thin sheet ligaments [18], thus rendering the
measured toughness independent of specimen thickness.
With the same argument as Embury et al. and Chan, the local
constraint in the tested specimens should decrease whensplits are
initiated and propagating. The local constraint should then depend
on the distance between two splits and on theseverity of respective
split. The crack growth should be higher in a region with higher
constraint (triaxial stress state) [1921]. Fig. 8 shows the
mid-thickness fracture surface of a 50 mm and 10 mm specimen and
the difference in crack growth dueto the presence of splits is
demonstrated. This relation between crack growth and splits has
been observed for all testedspecimens.
Furthermore, if the suggested explanation for the splits
(Section 4.1) is valid then it means that delamination is
favouredover cleavage crack propagation in the plane of the fatigue
pre-crack. This causes the specimen to exhibit stable
fracturebehaviour instead of a catastrophic failure, which
increased the fracture toughness.
To gain further information on how the splits affect the
fracture toughness, JR curves were calculated according to
thenormalization data reduction technique described in chapter A15
in ASTM E 1820-06 [22]. This method is based on the keycurve
approach and the principle of load separation and has been shown to
give equivalent JR results with the elastic com-pliance method
[23]. The normalization data reduction technique is not explained
here and the reader is referred to refer-ence ASTM E 1820-06 [22]
for more information. This method was only successful on the 50 mm
plate test data because themultiple pop-ins that occurred for the
30 mm and the 10 mm plate caused a discontinuous JDa graph with
large crack
-
246 J. Pilhagen, R. Sandstrm / Engineering Fracture Mechanics 99
(2013) 239250Fig. 8. SEM photos over crack growth in 50 mm plate:
(a) 50 mm specimen (14 magnication) and (b) 10 mm specimen (25
magnication). Black dotsindicate the boundary between crack growth
and post fatigue cracking. Areas with lower crack growth are marked
with an asterisk ().extension jumps and an oscillation in J values.
For JRmeasuring of steels which are prone to pop-ins, the elastic
compliancemethod may be less sensitive to pop-ins than the
normalization method [24].
For all JR curves of the 50 mm plate (including 50, 30 and 10 mm
specimens) the shape of the curve was the same. Afterinitial crack
extension the J value increased linearly with increasing crack
extension instead of having a power law relation,see Fig. 9. The
power law regression line is the tted line between the two
exclusion lines according to J = C1DaC2. The ten-tative initiation
toughness JQ is thus the intercept of the 0.2 mm offset line and
the tted line. JQ becomes JIc if the thicknessand initial ligament
is less than 25 JQ/rY and if the slope of the line in Fig. 9 at DaQ
is less than rY [22]. rY is the mean valueof the yield and tensile
strength at testing temperature. All specimens evaluated with the
normalization method except fortwo passed these conditions. These
two 10 mm specimens exceeded the constraint on the initial ligament
which is explainedby the present of larger than normal splits for
the 10 mm specimens.
0 0.5 1 1.50
50
100
150
200
250
Crack extension, [mm]
J, [k
N/m
]
JNorm.Power law regression line0.15 mm exclusion line1.5 mm
exclusion line0.2 mm offset line
Fig. 9. Resulting JR curve from the normalization data reduction
technique for a 50 mm specimen at T = 55 C. The number of data
points in the gure hasbeen reduced for clarity.
-
J. Pilhagen, R. Sandstrm / Engineering Fracture Mechanics 99
(2013) 239250 247The KJIc values for the 50 mm plate with different
specimen dimensions are shown in Fig. 10. With respect to the
numberof specimens and the scatter, the results indicate that the
specimen thickness does not inuence the initiation
fracturetoughness. The splits give almost constant fracture
toughness independent of the specimen thickness. A result
comparableto Guo et al. [25].
The splits also seem to inuence the post-fatigue cracking of the
specimens where some regions had large fatigue crackextension where
others were not affected at all by the fatigue cracking, see Fig.
6. This gives uncertainties on the extent of thecrack growth during
the testing when optical measuring the crack growth with LOM. The
highly irregular crack growth foundin the specimens, see Fig. 8,
makes it difcult to dene the crack extension, especially if the
split should be accounted for.
Using other methods like the potential drop or elastic
compliance technique instead of optical measuring of the
crackextension may not solve the problem with the splits. It is
plausible to assume that the free surfaces of the splits change
boththe electrical resistance in the specimen and the elastic
compliance of the specimen. This will give an erroneous
measure-ment of the crack extension that occurred in the plane of
the fatigue crack.
4.3. Pop-in
From the examination of fracture surfaces (broken up in liquid
nitrogen) of duplex stainless steel specimens that exhib-ited
pop-in, Wiesner [26] found no brittle crack propagation in the
plane of the fatigue crack. The pop-in was instead attrib-uted to
the splits that had occurred during the testing. In Fig. 11 this
lack of brittle crack propagation in the plane of thefatigue crack
is demonstrated. The depth of the split can be seen in Fig. 11b.
For this particular specimen the fracture tough-ness test was
stopped with subsequent fatigue cracking immediately after a pop-in
for revealing the amount of crack growththat had occurred during
the pop-in. As seen in Fig. 11a no change in fracture surface has
occurred between the two fatiguesurfaces except for crack-tip
blunting and delamination (split). Pop-in that occur due to the
formation of splits has been de-
100 90 80 70 60 50 40
100
150
200
250
300
350
Temperature [C]
K JIc
[MPa
m]
B = 50 mmB = 30 mmB = 10 mm
Fig. 10. Initiation fracture toughness for 50 mm plate of 2205.
The conversion from J-integral units to stress intensity factor
units is done according toKJIc =
p(JIc E/(1 t2)).ned as type II pop-in [8].The pop-in occurs due
to the sudden change in compliance of the specimen when the splits
are formed. A microstructure
with more elongated austenite lamellae is likely to cause a more
severe pop-in due to the longer uninterrupted crack pathbetween the
cleavage crack initiation and the crack front. This mechanism also
explain why it is common to nd splits thathave grown backwards in
to the fatigue pre-cracking region when a pop-in has occurred [8],
see Fig. 11a. This phenomenon ismore common and more visible for
the 10 mm plate compare to the 30 mm plate. For the 50 mm plate the
pop-in is likely tobe too small to be observed. The difference in
texture for the ferritic phase between the plates could also
contribute to theincreased pop-in behaviour for the thinner plates
[6].
The validity of type II pop-in in fracture toughness testing has
been evaluated in the work by Wiesner and Pisarski [8]. Itwas
concluded that the fracture toughness value at this pop-in is not a
true material property for the crack orientation sub-jected to
testing nor for the through-thickness orientation and can thus be
ignored. However, it is recommended that if thecomponent exhibit
through-thickness stresses then the through-thickness fracture
toughness (S-L/T/45 orientation in Fig. 1)of the material should be
evaluated [8,10].
4.4. The validity of ASTM E 1921 for the tested material
The ASTM E 1921-05 [14] standard is commonly used for evaluating
the fracture toughness in the ductile to brittle tran-sition
temperature region for ferritic materials. The standard decides a
reference temperature which characterizes the frac-ture toughness.
The reference temperature, T0, is dened as the temperature where
the median KJc for a 25 mm thickness
-
248 J. Pilhagen, R. Sandstrm / Engineering Fracture Mechanics 99
(2013) 239250specimfractupartic
Spe
KJc(0) ithickn
Eva50 mm[2,28]ditionmet a
wheredimenplate
Hounstabetwestrainquesti
Fig. 11(fatigueexamplen (B1T) is 100 MPapm. This test method was
originally developed for ferritic steels in the transition region
where the
re toughness is primarily controlled by the statistical event of
crack initiation from the cleavage of brittle second phaseles like
carbides and inclusions [14,27].cimens with different thicknesses
can be size adjusted to the reference thickness (B1T) by Eq.
(4).
KJc1T Kmin KJc0 Kmin B0B1T
1=44
s the measured toughness, Kmin is the threshold toughness
assumed to be 20 MPapm [14] and B0 is the specimen
ess.luating the data for the 50 and 30 mm plate according to
this method result in a T0 temperature of 135 C for theplate and
143 C for the 30 mm plate. This is in agreement with other reported
results for the 2205 base metal
. However, the standard for determining the reference
temperature T0 is based on the criteria of high constraint
con-along the crack front which theoretically requires small scale
yielding condition [14,29]. To assure that this criterion ist
fracture for the master curve method, a maximum KJc capacity is
dened:
KJclimit Eb0rys
M1 t2
s5
b0 is the remaining ligament length of the specimen, rys is the
yield strength at testing temperature and M is a non-sional
deformation limit which is set to 30 in the standard [14].
According to the KJc(limit) criteria the 30 and 50 mmspecimens have
a sufcient constraint condition along the crack front.wever, if one
considers a specimen where splits occur before or at stable crack
growth and the specimen later fails byble cleavage fracture, it is
reasonable to assume that the effective thickness of the specimen
is closer to the distanceen splits than the specimen thickness. The
measured fracture toughness is likely to increase due to the
reduced con-t in the material which results in higher plastic work
done before a critical cleavage initiation occurs. One can
thereforeon if the KJc(limit) is still valid for a specimen that
delaminate during testing.
. SEM photo over pop-in fracture during testing of the 10 mm
plate: (a) overview over the fracture surface showing (from right):
pre-cracking), crack-tip blunting and post fatigue cracking and (b)
depth of split in (a). Note that several splits are initiated
during a pop-in event and in thise seven similar splits along the
crack front were found. L, T and S according to Fig. 1.
-
J. Pilhagen, R. Sandstrm / Engineering Fracture Mechanics 99
(2013) 239250 249For the tested material the situation is however
different and more complex. The splits seem to be cleavage cracks
andtherefore a change in effective thickness occurs after cleavage
initiation. All the specimens exhibited a stable fracture
mech-anism where increasing displacement was needed for increasing
crack growth despite the presence of cleavage fracture inthe plane
of the fatigue pre-crack. Reported fracture toughness results of a
30 mm plate of 2205 DSS (TL orientation) statethat no brittle
failure (i.e. complete cleavage fracture) is observed down to 180 C
[30]. This is interfering with the weakest-link theory and the
assumption in the master curve method that the specimen failure is
primarily cleavage initiation con-trolled [27,31]. Therefore the
master curve analysis is not valid in this case because there is no
direct link between cleavagefracture initiation and specimen
failure.
A similar situation exists for the use of the maximum force
plateau as a failure criterion. The maximum force plateau oc-curs
when the rate of strain hardening of the material is balanced by
the rate of decrease of the remaining cross section [21].However,
this point is not directly associated with the crack extension in
the specimen and likely to depend on specimengeometry (size and
crack depth) and displacement rate. The maximum force plateau has
therefore been removed fromthe CTOD test standard in the later
revision of the test standard [32]. For specimens that only show
stable crack growththe CTOD at the end-of-test (deot) can be used
for quality control and specications of acceptance [32].
The specimens were stable after reaching maximum force plateau
under displacement control which indicates that thespecimens have a
rising R curve, recall Fig. 9. Therefore, it is suggested that for
assessing the fracture toughness at sub-zerotemperatures for the
2205 base metal the fracture resistance curve should be determined
(JR or dR). For assessing the frac-ture toughness regarding brittle
crack behaviour for the tested material at sub-zero temperatures it
is suggested to evaluatethe split initiation toughness. The
potential drop method might have the potential to detect the split
initiation when it occursin the specimen.
5. Conclusions
Fracture toughness testing of a duplex stainless steel were
performed on SE(B) specimens between 94 C to 18 C.Specimens were
taken from 10, 30 and 50 mm plate. From the 50 mm plate, 10 and 30
mm specimens were also cut out.The purpose was to examine how the
splits inuence the fracture toughness and how the microstructure
affects the splittingphenomenon in hot-rolled duplex stainless
steels.
Thinner plates have more extensive splitting than thicker plates
likely due to the more aligned and elongated austenitelamellae.
The consequence of initiated and propagating splits is the
creation of free surfaces which causes the specimen to behaveas two
or more individual subunits. This causes loss of constraints along
the crack front which increase the fracturetoughness. The loss of
constraints is also visible by the local difference in crack growth
in the specimen where the crackgrowth decreases with decreasing
distance between splits. From the normalization method, fracture
initiation valueswere obtained and the result shows that specimen
thickness does not inuence the initiation fracture toughness.
The pop-in phenomenon that can occur during fracture toughness
testing of rolled duplex stainless steels is coupled tothe
initiation of splits. This phenomenon is frequently observed for
the 10 and 30 mm plates but not for the 50 mm plate.The frequency
and severity of the pop-ins seems to increase with decreasing plate
thickness due to the renement of themicrostructure.
The tested specimens exhibit a stable fracture process despite
the presence of cleavage fracture likely due to the presenceof
splits. For assessing the fracture toughness regarding brittle
failure, the master curve method is suggested to not to bevalid for
the 2205 base metal at the tested conditions. For evaluating the
fracture toughness at sub-zero temperatures thefracture resistance
curve can be used.
A proposed failure criterion for assessing brittle crack
behaviour for these materials is the split initiation which may
beobserved with the potential drop method.
Acknowledgements
The authors would like to express their gratitude to the VINN
Excellence Center Hero-M and Outokumpu Stainless fornancing this
study. Outokumpu stainless is also gratefully acknowledged for
delivering the material and for all help. Valu-able support from
Mikael Johansson and Jan Y. Jonsson at Outokumpu Avesta Research
Centre is acknowledged.
References
[1] Dlouhy I, Chlup Z, Holzmann M. Crack length effect on
fracture behaviour of duplex steels. In: Brown MW, de los Rios ER,
Miller KJ, editors, Fracturefrom defects, ECF 12 1998;2:72732.
[2] Sieurin H, Sandstrm R. Fracture toughness of a welded duplex
stainless steel. Eng Fract Mech 2006;73:37790.[3] Sieurin H,
Sandstrm R, Westin ME. Fracture toughness of the lean duplex
stainless steel LDX 2101. Metall Mater Trans A 2006;37A:297581.[4]
Dhooge A, Deleu E. Fracture toughness of duplex and super duplex
stainless steels at low temperatures. Stainless steel world,
September 1995.[5] Sieurin H, Westin M E, Liljas M, Sandstrm R.
Fracture toughness of welded commercial duplex stainless steel.
Duplex 2007, 1820 June 2007, Grado,
Italy.
-
[6] Nilsson SA. Anisotropy in duplex stainless steel SS 2377.
Swedish Institute for Metals Research, IM-2551, February 1992. Can
be ordered at: .
[7] Erauzkin E, Irisarri MA. Effect of the testing temperature
on the fracture toughness of a duplex stainless steel. Scripta
Metall Mater 1991;25:17316.[8] Wiesner SC, Pisarski GH. The
signicance of pop-ins during initiation fracture toughness tests.
3R International 1996;35:63843.[9] Chao CH. Mechanism of
anisotropic lamellar fractures. Metall Trans A 1978;9A:50914.[10]
Pisarski GH, Hammond R, Watt K. Signicance of splits and pop-ins
observed during fracture toughness testing on line pipe steel. In:
Proceedings of
IPC2008, seventh international pipeline conference, Alberta,
Canada, 2008. p. 47382.[11] Yang Z, Huo C, Guo W. The charpy notch
impact test of X70 pipeline steel with delamination cracks. Key Eng
Mater 2005;297300:23916.[12] Ericsson C, Sandstrm R, Sieurin H,
Lagerqvist O, Eisele U, Schiedermaier J, et al., Material and
welding data. Properties and test results, background
document 3.5: duplex stainless steel, EcoPress, European
research 5th, framework; 2003.[13] ASTM E 1290-02: Standard test
method for crack-tip opening displacement (CTOD) fracture toughness
measurement. vol. 03.01, 2007 ed.[14] ASTM E 1921-05: Standard test
method for determination of reference temperature, T0, for ferritic
steels in the transition range. vol. 03.01.[15] Engel L, Klingele
H. An atlas of metal damage. London: Wolfe Publishing Ltd.;
1981.[16] Embury DJ, Petch JN, Wraith EA, Wrigth SE. The fracture
of mild steel laminates. Trans Metall Soc AIME 1967;239:1148.[17]
Chan SK. Evidence of a thin sheet toughening mechanism in AlFeX
alloys. Metall Trans A 1989;20A:15564.[18] Chan SK. Conrmation of a
thin sheet toughening mechanism and anisotropic fracture in AlFeX
alloys. Metall Trans A 1989;20A:233744.[19] Marrow JT, Humphreys
OA, Strangwood M. The crack initiation toughness for brittle
fracture of super duplex stainless steel. Fatigue Fract Eng
Mater
Struct 1997;20:100514.[20] Kikuchi M, Ishihara T. Study on
thickness effect of three-point-bend specimen. JSME Int J Ser A
2006;49:4117.[21] Anderson LT. Fracture mechanics: fundamentals and
applications. 2nd ed. CRC Press; 1995.[22] ASTM E 1820-06: Standard
test method for measurement of fracture toughness. vol. 03.01.[23]
Landes DJ, Zhou Z, Lee K, Herrera R. Normalization method for
developing JR curves with the LMN function. J Test Eval
1991;19:30511.[24] Zhu KX, Levis NB. Experimental determination of
constraint dependent JR curves for X80 pipeline steel using
normalization method. In: ASME
international mechanics engineering congress and exposition.
2006: p. 695703.[25] Guo W, Dong H, Lu M, Zhao X. The coupled
effects of thickness and delamination on cracking resistance of X70
pipeline steel. Int J Pres Ves Pip
2002;79:40312.[26] Wiesner SC. Toughness requirements for duplex
and super duplex stainless steels. In: Duplex stainless steels
975th world conference. 1997: p. 979
90.[27] Wallin K, Laukkanen A. New developments of the Wallin,
Saario, Trrnen cleavage fracture model. Eng Fract Mech
2008;75:336777.[28] Wallin K. Quantifying Tstress controlled
constraint by the master curve transition temperature T0. Eng Fract
Mech 2001;68:30128.[29] Ruggieri C, Doods Jr HR, Wallin K.
Constraint effects on reference temperature, T0, for ferritic
steels in the transition region. Eng Fract Mech
1998;60:1936.[30] U. Eisele, Schiedermaier J. Limit state models
derived, background document 7.2: ductile fracture, EcoPress,
European research 5th, framework; 2003.[31] Wallin K. The master
curve method: a new concept for a brittle fracture. Int J Mater
Product Technol 1999;14.[32] ASTM E 1290-08: standard test method
for crack-tip opening displacement (CTOD) fracture toughness
measurement. vol. 03.01, 2010 ed.
250 J. Pilhagen, R. Sandstrm / Engineering Fracture Mechanics 99
(2013) 239250
Loss of constraint during fracture toughness testing of duplex
stainless steels1 Introduction2 Material and testing procedure2.1
Material2.2 Specimen selection2.3 Testing procedure
3 Results4 Discussion4.1 Cause of splits4.2 Loss of
constraint4.3 Pop-in4.4 The validity of ASTM E 1921 for the tested
material
5 ConclusionsAcknowledgementsReferences