Creep Properties of Grade 91 Steel Steam Generator

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Procedia Engineering 55 (2013) 70 – 77

6th International Conference on Cree ep, Fatigue and Creep-Fatigue Interaction [CF-6 6]

Creep Properties of G Grade 91 Steel Steam Generator Tuube at 923K D.P. Rao Palaparti, E. Isaac S Samuel, B. K. Choudhary∗ ,M. D. Mathew Mechanical Metallurgy Division, Indira Gandhi C Centre for Atomic Research, Kalpakkam-603 102, Tamil Nadu, India

Abstract Creep-rupture properties of T91 steam generator (SG G) tube steel have been examined at 923 K in normalised and teempered condition in the stress range 55-150 MPa. At all stresss conditions, the creep deformation was characterised by a decrrease in creep rate in the transient creep followed by a minim mum creep rate in secondary creep and a rapid increase in creep p rate in tertiary creep stage. A systematic decrease in creep rate with decreasing stress was observed in transient, secondaary and tertiary creep. Stress dependence of minimum creep rrate and rupture life obeyed power law creep. Both minimum creeep rate and rupture life exhibited deviations in terms of loweer respective stress exponent values at low stresses than those obtained o at high stresses. A decrease in creep ductility was observed with increase in rupture life. The fracture mode reemained transgranular at all test conditions. Creep-rupture sstrength of SG tube steel has been found to be comparable to those reported in literature as well as specified in French Nuuclear Design Code, RCC-MR for the steel. © Authors. Published by Elsevier © 2013 2013The The Authors. Published by Ltd. Elsevier Ltd. Selection and/or peer-review Selection peer-review under responsibility Gandhi and Centre for Atomic Research. of the Indira Gandhi Centre for Atomic Research.

under responsibility of the Indira

Keywords: T91 steel; creep-rupture properties; power law crreep; creep ductility; monkman-grant relation

1. Introduction Grade 91 steel, the modified version of 9Cr--1Mo steel, is obtained by the addition of strong carbide//nitride forming elements such as Nb and V and nitrogeen, and is also known as 9Cr-1Mo-V-Nb steel designated as P91 or T91 by ASTM standards [1]. Grade 91 steeel is an important structural material for steam generato or (SG) application in thermal and nuclear power geenerating industries. The steel has been chosen for all a the components, i.e., tube, shell and thick sectionn tubeplate in the steam generators of Prototype Fast Breeder B Reactor (PFBR) in view of its superior elevatedd temperature mechanical properties including tensile, creeep and low cycle fatigue than the alternate 2.25Cr-1M Mo and plain 9Cr-1Mo steels and higher resistance to stress corrosion cracking in water/steam compared to austenitic stainless steels [2]. The 105 h creep-rupture sttrength of grade 91 steel has been found to be higher than 2.25Cr-1Mo and 9Cr-1Mo steels, and remains hig gher or equal to that of type 304 austenitic stainless steeel up to 898 K [3]. Further, the long-term creep properrties of

∗

Corresponding Author: E-mail address: [email protected]

1877-7058 © 2013 The Authors. Published by Elsevier Ltd. Selection and peer-review under responsibility of the Indira Gandhi Centre for Atomic Research. doi:10.1016/j.proeng.2013.03.221

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D.P. Rao Palaparti et al. / Procedia Engineering 55 (2013) 70 – 77

grade 91 steel has been observed to be superior than the conventional steels with Cr content up to 9%Cr, i.e., plain 9Cr-1Mo steel and 12Cr-1Mo-1W-0.3V steel [4]. The high creep and low cycle fatigue strengths of grade 91 steel allow the use of comparatively thinner tubes thereby improving the heat transfer efficiency. The creep-rupture behaviour of Gr. 91 steel is widely reported [3-14] and the steel has been included in the various codes and standards including French nuclear design code RCC-MR [15] being used for the design and construction of Indian PFBR. In this paper, creep deformation and rupture behaviour of indigenously developed T91 SG tube steel at 923 K are presented. The creep behaviour has been examined in terms of the variations of creep strain and creep rate with time, stress dependence of minimum creep rate and rupture life, creep ductility, fracture behaviour, creep rate-rupture life relations of Monkman-Grant type and creep damage tolerance factor. The creep-rupture strength of SG tube steel has also been compared with the values reported in the literature and specified in the design code RCC-MR [15] for the steel. 2. Experimental Grade 91 steel SG tubes of dimension 17.2 mm outer diameter and 2.3 mm wall thickness supplied by M/s. Nuclear Fuel Complex (NFC) Ltd., Hyderabad was used in this investigation. The chemical composition (wt. %) of the steel is given in Table 1. Table 1. The chemical composition (wt. %) of SG tube steel. Elements Amount (wt. %)

C

Cr

Mo

Si

Mn

V

Nb

N

S

P

Fe

0.10

9.00

0.92

0.23

0.46

0.18

0.08

0.084

0.008

0.018

Balance

The tubes were normalised at 1338 K for 0.5 h followed by air cooling and tempered at 1053 K for 1 h followed by air cooling. Flat creep specimens of 25 mm gauge length and 6 mm width were machined from the normalised and tempered (N+T) tubes (Fig. 1).

Fig. 1. Flat creep specimen machined from SG tube (all dimensions are in mm).

The inner and outer surfaces in parallel gauge region were kept unaltered and the gauge region had the curvature of the original SG tube. Creep tests were conducted in air environment in ATS make creep machines at 923 K employing stresses ranging from 55 to 150 MPa. The test temperatures were maintained within ± 2 K in all the tests. The elongation values were measured using high temperature extensometers and Mitutoyo make digimatic indicator (for short-term tests) and LVDT transducer (for long-term tests) attached to the bottom of the extensometer at room temperature. The strain resolutions were 1.2×10−4 in the Mitutoyo make digimatic indicator and 4×10−4 in LVDT transducer. Metallographic and fractographic examinations were performed using Philips XL30 scanning electron microscope (SEM). Metallographic specimen was prepared using standard metallographic techniques followed by immersion etching in 1 g picric acid + 5 ml HCl + 100 ml ethyl alcohol.

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3. Results and discussion 3.1. Microstructure The microstructure of N+T SG tube steel was composed of tempered lath martensite and precipitates at the prior austenite grain and martensite lath boundaries and in the intralath matrix regions (Fig. 2). It has been reported that the prior austenite grain and martensite lath boundaries contain relatively coarse Cr rich M23C6 carbides and the intralath matrix regions contain fine MX precipitates of vanadium and niobium [10,12,14]. It has been also reported that the number fraction of Nb(C,N) remains lower compared to V(C,N) precipitates in the normalised and tempered condition.

Fig. 2. Microstructures in normalised and tempered condition showing tempered lath martensite and precipitates at grain and lath boundaries and in the matrix regions.

3.2. Creep-rupture properties Representative creep curves in terms of the variations of creep strain (ε) and creep rate ( ε ) as function of time (t) at stress levels of 55, 65 and 75 MPa at 923 K are shown in Figs. 3 and 4, respectively.

Fig. 3. Typical creep curves showing creep strain-time plots for stress levels 55, 65 and 75 MPa at 923 K.

Fig. 4. Typical creep curves showing creep rate-time plots for stress levels 55, 65 and 75 MPa at 923 K.

At all stress levels, the creep deformation of T91 steel is characterised by a negligible instantaneous loading strain followed by a small transient creep strain, a secondary creep and prolonged tertiary creep regime (Fig. 3). The accumulation of creep strain in tertiary creep is observed to be large compared to that in transient and secondary creep regimes and a rapid creep strain accumulation is observed just prior to rupture. Log İ vs. logt plots exhibit a systematic decrease in creep rate with time in the transient creep followed by secondary creep characterised by a minimum creep rate ( İ m ) and a rapid increase in the creep rate in the tertiary creep (Fig. 4).

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Figure.5 shows the variations of minimum creep rate with applied stress (σ) as double logarithmic plots of

İ m vs. σ. The minimum creep rate increases with increase in applied stress and the stress dependence of

minimum creep rate obeyed Norton’s power law

İ m = A ı

n

,

Fig. 5. Stress dependence of minimum creep rate at 923 K for

T91 SG tube steel.

(1)

Fig. 6. Stress dependence of rupture life at 923 K for T91 SG tube steel.

where A and n are the stress coefficient and the stress exponent, respectively. A deviation in minimum creep rate is observed at low stress of 55 MPa with lower stress exponent than the value of n = 9.95 obtained for high stresses. The variations of rupture life (tr) with applied stress also obeyed power law as

t r = A' ı -n' ,

(2)

where A′ and n′ are the stress coefficient and the stress exponent, respectively (Fig. 6). Like log İ m vs. logσ, stress dependence of rupture life also exhibited a deviation at 55 MPa with lower stress exponent than n′ = 7.95 observed at high stresses. The variation of creep ductility in terms of % elongation and % reduction in area as a function of rupture life displayed decrease in ductility values with increase in rupture life. The variation in % reduction in area with rupture life is shown in Fig. 7 as an example. SEM examination of fracture surfaces of creep tested specimens revealed transgranular fracture characterised by dimples resulting from coalescence of microvoids at all stress conditions studied in this investigation (Fig. 8).

Fig. 7. Variation of % reduction in area with rupture life at 923 K.

Fig. 8.SEM fractograph showing transgranular fracture in specimen creep tested for at 923 K at 112 MPa for 332 h.

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T91 steel followed generalised form of Monkman-Grant relation (MGR) [16] interrelating minimum creep rate and rupture life as α ε m ⋅ t r = C ,

(3)

where C is the Monkman-Grant constant (Fig. 9). The value of slope of logtr vs. log İ m , i.e. α = 0.815 has been observed at 923 K. Generally, the validity of MGR is demonstrated with α equal to unity giving rise to a real constant value for the product of minimum creep rate and rupture life for the range of stresses and temperatures [13]. The observed lower value of α = 0.815 than unity can be ascribed to the loss of creep ductility and its associated effects on creep-rupture properties at 923 K. Unlike MGR, T91 steel obeyed modified Monkman-Grant relation (MMGR) proposed by Dobes and Milicka [17] as shown in the double logarithmic plots of tr/εf vs. İ m (where εf is the strain to failure) at 923 K in Fig. 10. The generalised form of MMGR is expressed as

ε sα′ ⋅

tr = C′ εf

Fig. 9. Rupture time vs. minimum creep rate plot showing applicability of generalised form of Monkman-Grant relation in T91 SG tube steel.

(4)

Fig. 10. Rupture time/strain to failure vs. minimum creep rate plot showing the validity of modified Monkman-Grant relation in T91 SG tube steel.

where C' is the modified Monkman-Grant constant. The validity of MMGR is observed as slope α′ equal to unity in the plot log(tr/εf) vs. log İ m (Fig. 10). When α' = 1, the modified Monkman-Grant constant C' becomes CMMG and MMGR is expressed as

ε m ⋅

tr = constant = CMMG . εf

(5)

The values of CMMG = 0.16 and 0.24 are observed at low and high stresses, respectively for the steel at 923 K. The observed creep deformation characterised by transient, secondary and tertiary creep stages in T91 steel (Figs. 3 and 4) is in agreement with those reported for the steel [5-7,9,13]. The occurrence of minimum creep rate results from a balance between hardening and recovery processes dominant in the transient and tertiary creep regimes, respectively. The stress exponent n = 9.95 obtained at 923 K is rather high compared to that predicted by dislocation climb controlled models of power law creep. However, this is comparable to those reported for the steel [5,6,11,13] in the high stress regime. It has been shown that the high values of stress exponent and activation energy observed in the high stress regime in 9%Cr steels can be rationalised by

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invoking the concept of back/resisting stress into power law relationship [18]. Like high stress exponent values in the stress dependence of creep rate, stress dependence of rupture life also exhibits high stress exponent values n′ = 7.95 at 923 K. The observed deviations in the minimum creep rate (Fig. 5) and rupture life (Fig. 6) due to acceleration of creep deformation and consequent reduction in rupture life at low stresses is in agreement with those reported for the steel [11,14]. It has been suggested that the dominance of recovery processes results in degradation in the long-term creep-rupture properties in Gr. 91 steel [12,14]. Sawada et al. [12] observed dominance of subgrain coarsening and associated decrease in dislocation density, preferential recovery at prior austenite grain boundaries resulting in the formation complex nitride phase, the Z-phase and consequent reduction in major strengthening MX particles in the long-term creep regime [12]. The breakdown of creep strength at low stresses has been attributed to static recovery of subgrains controlled not by the aggregation of MX particles but by M23C6 precipitates in Gr. 91 steel, recently [14]. The degradation in creep-rupture properties and the dominance of recovery processes at low stresses are reflected in the significant loss of creep ductility as shown in Fig. 7. A considerable shift in the stress range signifying low stress regime for % reduction in area is seen compared to those displayed by minimum creep rate and rupture life (Figs. 5 and 6). This shift in stress range that can be identified as low stress regime has also been observed for the validity of modified Monkman-Grant relation shown in Fig. 10. The observed low values of CMMG of 0.16 and 0.24 at low and high stress regimes, respectively clearly indicate low contribution of secondary creep strain towards the overall creep ductility of the steel. The large contribution of tertiary creep strain is also reflected in the high creep damage tolerance factor observed for T91 steel. Creep damage tolerance factor λ is a significant parameter that assesses the susceptibility of a material to localised cracking at strain concentrations [19-21]. It is suggested to be a better measure of creep ductility as it is related to the ability of a material to redistribute stresses [19]. Creep damage can occur by various mechanisms such as loss of external cross section, loss of internal cross section (formation, growth and linkage of cavities at grain boundaries), degradation of microstructure (thermal-coarsening of particles and/or substructure-induced acceleration of creep) and oxidation [20]. Each damage micromechanism, when acting alone, results in a characteristic shape of creep curve and a correspondingly characteristic value of λ [20]. For example, damage due to growth of cavities by coupled diffusion and power-law creep results in λ values between 1.5 and 2.5, whereas it can be as high as 5 or more when thermal-coarsening of particles cause damage. λ is defined [19,20] as the ratio of strain to failure εf to secondary creep strain as

λ=

εf . ε m ⋅ t r

(6)

On rearrangement, λ can be expressed as

εf = λ ⋅ ε m . tr

(7)

The ratio of strain to failure and rupture time (εf/tr) is the average creep rate, and the double logarithmic plot of εf /tr vs. ε m gives the value of intercept as λ as shown in Fig. 11. Increase in damage tolerance factor λ = 4.2 at high stresses to λ = 6.4 at low stresses also indicate dominance of recovery in the low stress regime in T91 steel. 3.3. Comparison of creep-rupture strength The creep-rupture strength of SG tube steel at 923 K has been compared with those reported in the literature in Fig. 12. The average and the minimum creep-rupture strength values given in design code RCC-MR [15] at 923 K for the steel are also superimposed in Fig. 12.

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Fig. 11. Average creep rate vs. minimum creep rate showing high value of creep damage tolerance factor (λ) in T91 SG tube steel.

Fig. 12. Comparison of the creep strength of T91 SG tube steel with strength values reported in the literature and specified in RCC-MR at 923 K.

SG tube steel exhibited creep-rupture strength comparable to those reported in literature and the average creep-rupture strength values given in the design code RCC-MR for the steel. The above comparison suggests that the indigenously developed SG tube steel satisfies the long-term creep-rupture strength requirement for steam generator application. 4. Conclusions T91 SG tube steel negligible instantaneous strain, a small transient creep stain and a secondary creep followed by a prolonged tertiary creep regime. The secondary creep was characterised by a minimum creep rate. The stress dependence of minimum creep rate and rupture life obeyed power law and exhibited high values of stress exponents. Both minimum creep rate and rupture life exhibited deviations at low stresses with lower respective stress exponent values compared to those observed at high stresses. The fracture mode remained transgranular. Loss of creep ductility was observed at low stresses. The steel followed generalised form of Monkman-Grant relation. T91 steel obeyed modified Monkman-Grant relation, and exhibited high creep damage tolerance factor. The steel exhibited creep-rupture strength comparable to those reported in literature as well as given in RCC-MR.

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