the brittle fracture o f 475°c embrittled cast dup l ex stainless steel

8/2/2019 THE BRITTLE FRACTURE O F 475C EMBRITTLED CAST DUP L EX STAINLESS STEEL

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Fatigue Fract. Engng Mater. Struct. Vol. 20, No. , pp. 565-571, 1997Printed in Great Britain. All rights reserved 8756-758X/97 $6.00+0.00Copyright Q 1997 Fatigue & Fracture ofEngineering M aterials & Structures Ltd

TH E BRITTLE FRACTURE O F 475C EMBRITTLED CAST DUPLE XSTAINLESS STEELT. J. MARROWnd N. BURY

Manchester Materials Science Centre, University of Manchester and UMIST, Grosvenor Street, Manchester, M1 7HS, UKReceived i n finalform 20 December 1996

Abstract-The fracture behaviour of cast duplex stainless steels, heat treated to different ferrite contentsand hardness was investigated using tensile and notched bend tests. The purpose was to identify themicrostructural features which controlled the ductile-to-brittle fracture transition of 475C ernbrittledduplex stainless steel. The results indicate th at tw in nucleated cleavag e has a tensile stress fracture criteriaand the brittle-to-ductile transition temperature depends on ferrite microhardness, ferrite grain size andconstraint.Keywords-Duplex stainless steel; 475C ernbrittlement; Deformation twinning; Brittle-to-ductile trans-ition; Age hardening.

NOMENCLATUREb = Burgers length.D = slip band lengthG = shear modulusy = fracture energy

( T ~ tensile fracture stresstf= critical resolved shear fracture stressti=dislocation friction shear stresst, = critical resolved shear stress to nucleate a crackv = Poissons ratio

INTRODUCTIONIt has been previously shown that brittle fracture of duplex stainless steel is nucleated bydeformation twinning [l]. Brittle fracture is encouraged by the increase in ferrite yield stress atlow tem peratures and with age-harden ing (475C embrittlement). The previous study used a finiteelement model of the deformation around a notch to demonstrate that brittle fracture initiationin notched samples required a cr itical shear stress acting over a critical distance . This critical shearstress was com parable to the shear stress at fracture in sm ooth tensile specimens that failed in abrittle manner.Although providing a quantitative insight into the problem of 475C embrittlement in duplexstainless steels, the model had some flaws. First; the finite element analysis treated the duplexmicrostructure as a hom ogeneous continuum . Second , and m ore importantly; although the criticalshear fracture stress was consistent with twin nucleated cleavage, it did not account for the observedeffect of tensile stress which caused ferrite cleavage in the cen tre of necked tensile specim ens [11.A nucleation-controlled twin-nucleated cleavage model also did not provide a satisfactory expla-nation for the occurrence of a brittle-to-ductile transition, nor for the observations that yieldingin age-hardened duplex stainless steels occurred w ith profuse twinning, yet fracture occu rred above

the yield stress after some plastic strain.565


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566 T. J. MARROWnd N. BURY

Steel andAnnealing temperatureFerrite Content (%)Ferrite Grain size (pm)Ferrite subgrain size (pm)Vickers Hardness (30 g)(agedone w k t 475'C)

This paper addresses some of these concerns and attempts to identify the important micro-structural parameters which control the ductile-to-brittle transition. A similar finite element analysishas been used to determine the effects of the notch, ferrite content, ferrite microhardness, ferritesubgrain size and ferrite grain size in heat treated cast duplex stainless steels.

SteelA SteelA Steel3 SteelB1l00"C 1300'C 1looOC 1300C47*2 7 2 3 51*2 83*5

90M150 1000+200 40W60 75WlOO12.i4 1%4 16i4 16a4

32&8 408i22 327k7 346f6

EXPERIMENTAL DETAILS AND RESULTSMaterials

Two cast duplex stainless steels were used. Energy dispersive X-ray analysis gave the followingcompositions (in wt.%): Steel A (25.5 Cr, 6.3 Ni, 3.8 Mo, 0.8 Mn), and steel B (25.4 Cr, 5.6 Ni,2.0 Mo, 0.7 M n, 0.5 Si). The light element content w as not determined. Both steels were annealedas blanks a t 1100C and 1300C for 1 h and water quenched. Test specimens were then machinedand aged at 475C for up to 640 h. The ferrite contents were determined using image analysis ofmetallographic samples electrolytically etched in 30% KOH solution. The ferrite subgrains wererevealed by an electrolytic oxalic acid etch, an d the ferrite grain size was inferred from thedistribution of the austenite (Fig 1). The results are given in Table 1, with the Vickers hardness(3 0 kg diamond pyramid indenter) after ageing for approximately one week at 475C.Tensile tests (steel A )

Tensile tests (diameter 5 mm) were performed at a strain rate of 0.01 s - l at room temperature(20C) using steel A. The proof stress at 1% plastic strain was taken as the yield stress. Audibleclicking (acoustic emission) was heard well below the yield stress in all samples, increasing in

Fig. 1. Duplex stainless steel microstructures: (a) steel A, annealed at 1100C for 1h and water quenched,(b ) steel A, annealed a t 1 3 0 C for 1 h and water quenched.


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Brittle fracture of du plex stainless steels 567

5Wr-

AGEING TIME (hours)

0 00

s'iracture stress4W 450 50 0MCKERSHARDNESS

Fig. 2. Effect of annealing temperature ( 1100C and 1300C) on the relationship between Vickers hardness(30 kg): (a) ageing time, (b ) yield stress ( 1 % proof stress). The tensile fracture stress for tests which failedwithout yielding, and the critical hardness for the brittle-to-ductile transitions (BDT) are also shown.

frequency at yield with very little emission after yield. The effect of ageing time on yield stress,hardness and fracture behaviour was determined. The results are shown in Fig. 2. Ageing had agreater effect on the hardness of the 1300C annealed steel. Both steels showed the same lineardependence of yield stress and hardness, except for the 1300C annealed steel aged for 640 h(hardness 447 & 20). This material failed below the expected yield stress.A transition from ductile to brittle fracture was observed with increasing hardness. The fracturesurface was either completely ductile (Fig. 3(a)) or completely brittle with ferrite cleavage andductile shearing of the austenite (Fig. 3(b)). No samples failed with both cleavage and ductilefailure of the ferrite. Brittle fracture generally occurred after yielding with plastic strains of theorder of 3-10%. Only the material annealed at 1300C and aged for 640 h failed without yielding.

Fig. 3. The effect of ag eing time on the fracture of steel A, annealed at 1100C. (a) Ductile fracture (agedfor 48 h). (b ) Brittle fracture (aged for 640 h) .

Table 2. The V ickers hardness a nd ferrite microhardness on either side ofthe brittle-to-ductile transition in steel A, annealed at 1100C and 1300Cand aged at 475C

Ageing time (hours)Fracture mode

Vickers hardness 30 kg)


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568 T . J . MARROWnd N. BURYThe Vickers hardness (30 kg diamond pyramid indenter) with ferrite Vickers microhardness(50 g diam ond pyramid indenter) were measured for tests on either side of the brittle-to-ductiletransition. These results are in Table 2.

Tensile and bend tests (steel B)Test specimens from steel B, annealed at 1100C and 1300"C, were aged at 475C fo r 168 h. Thehardness values are given in Table 1. The ferrite (Fe) microhardn ess (50 g) after ageing for bothannealing temperatures was 380&- 30. Tensile tests (dia. 7 mm ) were performed at a s train rate of

0.01 s - l at temperatures between - 10C an d -60C. All the samples failed in a brittle mannerafter yielding with plastic strains of up to 20%. A tensile specimen of the steel which had beenannealed at 1100C and tested at - 0C was sectioned along its axis for m etallographic examin-ation. Linear features were observed which spanned the Fe grains. They did not appear to beinfluenced by either the Fe subgrain boundaries or the austenite within the Fe grains (Fig. 4).Bend specimens with a 60" notch (notch root radius 250 pm) were tested in 3 point bending atthe same temperatures, at a displacement rate of 1mm/minute with a loading span of 80m.Aclip strain gauge at the no tch was used to monitor the specimen deflection. All tests showed non-reversible plastic strain, with occasional load drop s before brittle fracture. A model of the elastic-plastic deform ation at the notch root was constructed using LU SAS finite element software. Thedetails of the analysis were similar to those prev iously published [1,2]. The yield stress and failureload were used to determine the maximum shear and tensile stresses at the notch at failure. Astrain hardening exponent of 0.1 was used, which had previously been found to be insensitive toage-hardening in a wrought duplex stainless steel [11.The finite element model calculated that the shear stress ahead of the notch root decreasedrapidly with increasing distance up to 140pm and then decreased by less than 20% between140pm and 1.1.mm. The tensile stress increased by approxim ately 10% over the sam e distances.The stresses increased with strain hardening. The behaviour close to the no tch surface may be anartefact of the m esh used in the model. The shear stress and tensile stress at the d istance of 200 pmwere taken to represent the average stress state operating over a distance of the o rder of the ferritegrain size. The stresses at failure were calculated using the failure load. The results are shown inFig. 5 , compared with the shear and tensile stresses at failure in the tensile tests. The shear stresswas taken as half the m aximum true tensile stress, which was calculated using the tensile specimenfracture surface area. There is good ag reement between the shear stress at fracture in the smoothand notched tests, and no agreement of the tensile stress. The re is no significan t difference betweenthe results for the two microstructures.

Fig. 4. Linear features in the gauge length of a tensile specimen of steel B,annealed at 11Oo"C, aged for168 h at 475C and tested at - 0C.


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Brittle fracture of duplex stainless steels 569

2000

rma m

500

0

tAA

i i

x 1100C Sbear Fracture Stress (smooth)o 1100C Shear Fracture Stress (notched)+ 1300C Shear Fracture Stress (smooth)13OO0C Shear Fracture Stress notched)

A 1100C Tensile Fracturo Stress (smooth)o 1100C Tensile Fracture Stress(notched)A 13OOOC Tensile Fracture Stress smooth)rn 13OO0C Tmsile Fracture Stress (notched)

-60 -40 -20 0TEMPERATURE(C)Fig. 5. A comparison of the shear and tensile stresses in notched and smooth specimens at the failureload in steel B, annealed at 1100C and 1300C,aged for 168 h at 475C and tested between - 10Cand- 0C. The shear fracture stress (smooth) and tensile fracture stress (smooth) were measu red in tensilespecimens. The shear fracture stress (notched) and tensile fracture stress (notched) were calculated for thenotched end specimens.

DISCUSSIONThe brittle-to-duc tile transition

The coarse ferrite grain size ensured that the brittle fracture propagating in one grain wassufficient to cause failure. This simplified the identification of the brittle-to-ductile transition.Annealing steel A at 1100C and 1300C produced microstructures with different ferrite contents,but with a comparable ferrite subgrain size and a small difference in ferrite grain size (Table 1).The change in ferrite chemical composition with increasing ferrite content increased the rate ofage-hardening, producing brittle fracture within a shorter time in the 1300C annealed steel. Brittlefracture in a wrought duplex stainless steel was caused by deformation twinning [ 11. The acousticemission during loading and the linear features in the ferrite (Fig. 5) are both consistent withdeformation twinning. It is proposed that ferrite cleavage in cast duplex stainless steels isdeformation twin nucleated.It is important to determine whether twin nucleated cleavage is nucleation controlled (difficultcrack nucleation, easy crack propagation) or propagation controlled (easy crack nucleation,difficult crack propagation). Twin nucleated cleavage is generally considered to be nucleationcontrolled [3,4]. If cleavage is nucleated by the nucleation controlled Stroh mechanism [S], thenthe critical shear fracture stress, zf, is

zf= Ti+ 7,z = 2741G E)

where T~ s the friction stress, G is the shear modulus, v is Poissons ratio, b is the Burgers lengthand D is the length of a dislocation pile-up equivalent to the arrested twin. If the twin length islarge then the friction stress due to age-hardening is the most significant component of the fractureshear stress. For example, with a pile-up length, D , of 200 pm, Burgers length of 0.25 nm, Poissonsratio of 0.28 and shear modulus of 81 GPa, then z is 40 MPa. The maximum shear stress atfracture is of the order of 400 MPa (Fig. 5) . The friction stress would therefore account for


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570 T. J. MARROWnd N. BURYapproxim ately 90" of the fracture stress. The critical shear stress to nucleate cleavage is therefo reexpected to be very insensitive to the ferrite conten t and ferrite grain size. The re is no conclusiveevidence that the brittle-to-ductile transition in ferritic and duplex stainless steels is mechanismcontrolled [ ,2], and the brittle-to-ductile transition may therefore depend on plastic relaxationof the elastic strain at the tip of an arrested twin. This would be contro lled by the effects of age-hardening and temperature on the mobility and ease to cross-slip of dislocations [S]. The lengthof the arre sted twin would therefore have no effect on the brittle-to-duc tile transition, which wouldbe determined by the fe rrite yield streng th, measurab le as the ferrite micro hardn ess.Alternatively, if the re is a mechanism of easy crack nuc leation by deform ation twinn ing [7],the brittle-to-ductile ransition will depend on both the crack nucleus size and the m aximum tensilestress in the ferrite. Duplex stainless steels are clean steels and no evidence has been foundsuggesting that cleavage is nucleated a t large b rittle inclusions. According to the Cottrell mechanismof crack nucleation at intersecting twins and dislocations [ 1, the brittle-to-ductile transitionoccurs at a c ritical tensile stress, of,

a f n b= 2y (3)where n and b are the numb er and the Burgers length of the dislocations in a dislocation pile-upequivalent to the arrested twin. The fracture energy is y . The nucleated crack length thereforeincreases with the length of the twin, decreasing the critical tensile fracture stress. The maxim umtensile stress in the ferrite wou ld then depend on the ferrite yield streng th and the level of constraint.An effect of co nstraint, microstructure and ferrite microhardn ess on the brittle-to-ductile transitionis expected.Hence, both mechanisms predict that the brittle-to-ductile fracture transition depends on theferrite microhardness. The critical microhardness may or may not be affected by microstructure,depending on the crack nucleation mechanism . The results for steel A indicate that brittle fractureoccurs above a critical ferrite microhardness of between approxim ately 550 and 590 in the steelannealed at 1100C and between 490 and 540 in the 1300C annealed steel. Annealing at thehigher temperature increased the average ferrite grain size, with no change in the ferrite subgrainsize (Table 2). The decrease in the critical ferrite microh ardness of the brittle-to-duc tile transitionwith increasing twin length is not cons istent with a nuc leation contro lled cleavage mechanism andsupports p ropagation controlled cleavage.The notch effect.

The preceding discussion concluded tha t b rittle fracture requ ired a critical tensile stress. This issupported by the observation of cleavage in the centre of the fracture surface of necked tensilespecimens [l]. The finite element model, however, shows that brittle fracture required a criticalshear stress, acting over a distance of the o rder of the g rain size, which was similar in both notchedand smooth specimens. Fracture occurred after yielding, and the critical shear stress was similarto, but h igher than the yield stress. It is suggested tha t the tensile stress near the no tch tip generallyexceeds the critical tensile stress for cleavage propagation, and brittle fracture requires cracknucleation within a critical volum e ahead of the no tch tip. This volume may reaso nably be assum edto depend on the microstructure and to be related to the probability of forming a favourablyoriented crack nucleus. This fracture criterion is equivalent to a critical shear stress acting over acritical distance. In smooth tensile specimens, strain hardening is necessary to achieve a sufficienttensile stress for the propagation of cracks nucleated at yield. The strain hardening rate in duplexstainless steels is low [ 1 and consequently the calculated critical shear stress for brittle fractureis close to the shear stress a t yield. The low failure stress observed in steel A, annealed at 1300C


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Brittle fracture of duplex stainless steels 571and aged for 640h implies that the ferrite microhardness was sufficiently high to propagatemicrocracks nucleated by the small number of twins formed below th e yield stress.Both microstructures of steel B were tested below their brittle-to-ductile transition temperature.Fracture occurred in the notched sam ples when the grains aro un d the notch roo t yielded and aftersome strain hardening of the sm ooth tensile specimens. The yield stress of an age-hardened duplexstainless steel is dom inated by the ferrite yield stress and dep ends on th e ferrite conten t, as observedin steel A (Fig. 2). However, the yield stress of the 1300C annealed steel B was not significantlyhigher than the 1100C annealed steel B, despite the significant increase in ferrite content. Bothsteels has similar ferrite microhardn ess and yield occurred by twinning. Th e increase in grain sizeand consequent decrease in the critical shear stress for yield by twinning, since the twin length isthe ferrite grain size, is considered to have counteracted the effect of ferrite content. Both heattreatments of steel B therefore had similar fracture behaviour, although a difference in the brittle-to-ductile transition temperature or hardness is expected, similar to that observed in the two heattreatments of steel A.The tensile stress at the tip of a sharp crack is higher than that ahead of a blunt notch. Thefracture toughness of fatigue cracked age-hardened duplex stainless steels should therefore bedetermined by the development of a critical yielded volume, which is analogous to a critical shearstress acting over a critical distance. Work is in progress to develop a model for the effects ofmicrostructure, ferrite hardness and temperature on the fracture toughness of duplex stainlesssteels. Propagation controlled cleavage requires a n easy crack nu cleation mechanism. At present,there is no experimental evidence positively identifying this mechanism.

CONCLUSIONSBrittle fracture in age-hardened (475C embrittled) duplex stainless steels is twin nucleated an dpropagates at a critical tensile stress. The brittle-to-ductile transition temp erature depend s on the

stress state, the ferrite microhardn ess and the ferrite grain size.Acknowledgements-The auth ors would like to than k Professors G. W. Lorimer and R. . Taylor for the provision offacilities at the Man chester Materials S cience Centre.

REFERENCES1. T. J. Marrow and C. Harris (1996) The fracture mechanism of 475C embrittlement in a duplex stainless2. T. J . Marrow (1996) The fracture mechanism in 475C embrittled ferritic stainless steels. Fatigue Fruct.3. D. Hull (1960) Twinning and fracture in single crystals of 3% silicon iron. Acta Metall . 8, 11-18.4. M. Sarfarazi and S. K. Gho sh (1987) O n the microstructural theories of stress-induced cleavage5. A. N. Stroh (1955) The formation of cracks in plastic flow 11. Proc. Roy. SOC.232A, 548-560.6. C. J. McM ahon (1967) The microstructural aspects of tensile fracture, Fundamental Phenomena in the7. R. Lagneborg (196 7) Yielding and fracture of F e-30%Cr alloys subjected to 475C-em brittlement. Acta8. A. H. Cottrell (1958) Theory of brittle fracture in steel and similar metals. Trans Metall. SOC.A IM E .

steel. Fatigue Fract. Engng Muter. Struct. 19, 935-947.Engng Muter. Struct. 19, 919-933.

microcracking in crystalline solids. Engng Fract. Mech. 27, 215-230.Material Sciences, Vol. 4, pp. 247-284. Plenum , New York .Polytech. Scand. Ch. 62.212, 192-203.

the brittle fracture o f 475°c embrittled cast dup l ex stainless steel

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