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Creep deformation behavior at long-term in P23/T23 steels

K. Sawada, M. Tabuchi and K. Kimura

National Institute for Materials Science, Japan

34th

MPA-Seminar

Materials and Components Behaviour in Energy & Plant Technology

October 9 and 10, 2008 in Stuttgart

Abstract

Creep behavior of ASME P23/T23 steels was investigated and analyzed,

focusing on creep strength degradation at long-term. Creep rupture strength at 625oC

and 650oC dropped at long-term in both P23 and T23 steels. The stress exponent of

minimum creep rate at 625oC and 650oC was 7.8-13 for higher stresses and 3.9-5.3

for lower stresses in the P23/T23 steels. The change of stress exponent with stress

levels was consistent with the drop in creep rupture strength at long-term. The

Monkman-Grant rule was confirmed in the range examined in P23 steel, while the

data points deviated from the rule at long-term in the case of T23 steel. The creep

ductility of P23 steel was high over a wide stress and temperature range. On the other

hand, in T23 steel, creep ductility at 625oC and 650oC decreased as time to rupture

increased. The change in ductility may cause the deviation from the Monkman-Grantrule. Fracture mode changed from transgranular to intergranular fracture in the

long-term at 625oC and 650oC.

1. Introduct ion

Modern ultra supercritical ( USC ) power plants with high thermal efficiency have

been realized due to development of creep strength enhanced ferritic steels. The

development of creep strength enhanced ferritic steels such as ASME Gr.91, Gr.92

and Gr.122 strongly contributed to improvement of thermal efficiency in powerplants.[1] Furthermore, ASME P23/T23 (2.25Cr-1.6W) with lower Cr content has been

developed for reducing the cost of power plants construction.[2] The creep strength of

P23/T23 is superior to 2.25Cr-1Mo steel. Recently, however, allowable tensile stress

of Gr.91, Gr.92 and Gr.122 has been reduced in Japan [3-6] since it became clear that

the creep rupture strength of these steels abruptly drops at long-term. The allowable

tensile stress of P23/T23 has also been reviewed in Japan.[4] Remarkable drop in

creep rupture strength was observed on P23/T23 at higher temperatures in the

long-term. However, it was assumed that the drop in creep rupture strength was due

to mainly an oxidation effect since thick oxide scale was observed after creep.[4]

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Yoshizawa et al. reported that the drop of creep rupture strength for P23/T23 at higher

temperatures can be explained by estimating the effect of oxidation.[7] It has been

confirmed, by measuring the thickness of oxide scale and calculating true stress, that

oxidation causes the drop in creep strength in 2.25Cr-1Mo steel.[8,9]

Kushima et al. reported that preferential recovery of martensite around prior

austenite grain boundary causes the remarkable drop in creep rupture strength for

Gr.91 with tempered martensite.[10] In the same way as Gr.91, the preferential

recovery takes place after long-term creep in ASTM A542(2.25Cr-1Mo) with

martensitic structure.[11] In fact, the drop in creep rupture strength occurs at the

long-term in ASTM A542.[12] That means the preferential recovery contributes to the

creep strength degradation in ASTM A542. It is necessary to consider that in P23/T23

with a bainite structure, creep strength degradation may occur due to microstructural

changes.[13] The bainite structure has a relatively high dislocation density and lathstructure like a tempered martensite, indicating occurrence of the preferential

recovery around prior austenite grain boundary. Komai et al. reported that the creep

strength of P23/T23 abruptly decreases in the long term at 600oC.[13] However, they

also reported that the drop in creep rupture strength does not occur in P23/T23, if it

was tempered for a long time before creep.[13] They pointed out that

non-homogeneous recovery such as the preferential recovery around prior austenite

grain boundaries can not take place in the P23/T23 since long-time tempering can

decrease dislocation density in bainite structure. Bainite structure with a lowerdislocation density may have no inhomogeneity of internal stress due to accumulation

of dislocation, indicating that the structure can homogenously recover during creep.

Consequently, there are two possibilities as the reason for the creep strength

degradation in P23/T23. One is that the oxidation causes the reduction of creep life in

the long-term, meaning that the degradation is not due to an intrinsic material property.

Another possible reason is that the drop in creep strength occurs due to

microstructural changes. If the former is the reason for reduction of creep life, the

current allowable stress of P23/T23 does not need to be reviewed.[4] However, if the

latter is the reason, it will be necessary to review the allowable stress again,

considering the drop in creep rupture strength in the long-term.

In this paper, we characterize creep behavior for three heats of P23/T23, and

then discuss whether microstructural change can cause the drop in creep strength or

not. The analysis for creep behavior can provide us the information on the

deformation mechanism and microstructural change.

2. Experimental procedure

The materials examined are ASME T23 and P23.[14] Two heats for T23 tube and

one heat for P23 pipe were used for creep tests. The chemical compositions and heat

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treatment conditions are summarized in Table 1. The Vickers hardness and 0.2%

proof stress at room temperature are listed in Table 2. All steels retained a bainitic lath

structure as shown in Fig.1. Creep tests were performed under constant load in air,

using specimens of 6mm (heat A, B) or 10mm (heat C) in gauge diameter and 30mm

(heat A, B) or 50mm (heat C) in gauge length. Displacement was continuously

measured by extensometer for high temperature use during creep.

3. Results and Discussion

3.1 Creep rupture strength

Creep rupture strength for the three heats is shown in Fig.2. For all of the heats,

there is no large drop in creep strength in the long-term at 500oC to 600oC. However,

creep rupture strength of all heats abruptly decreases in the long-term at 625oC and

650o

C. This tendency is consistent with the results reported in the literature.[13] Thedrop in creep rupture strength in the long-term for heat C is not so remarkable

compared with heat A and B. The gauge diameter of heat C is larger than those of

heat A and B. If oxidation affects creep strength, the influence of oxidation may be

small in heat C since the fraction occupied by oxide scale per cross-section area is

smaller in heat C in contrast with other heats. (gauge diameter : 6mm for heat A and B,

10mm for heat C) This may cause the slighter drop in creep strength at 625oC and

650oC in heat C as shown in Fig.2. On the other hand, differences in initial

microstructure can also affect creep strength degradation. It is expected that initialdislocation density in heat C will be lower than those of heat A and B since the

hardness of heat C is lesser than those of other heats. It is reported in high Cr ferritic

steels that creep rupture strength of steel with a high dislocation density abruptly

decreases in the long-term.[15] In short, the creep strength degradation of heat C may

be slighter than those of other heats due to not only the low fraction occupied by oxide

scale but also due to low dislocation density.

3.2 Creep deformation behavior

Fig.3 shows relationship between minimum creep rate and stress for all steels. A

linear relation is observed for all steels over a wide range of applied stresses at 500oC

to 600oC. There is, however, inflection at a low stress at 625oC and 650oC for all

steels. The magnitude of slope in Fig.3, that is a stress exponent, are summarized in

Table 3. For all steels, the stress exponent is about 6.7 18 at 500oC to 600oC and

about 7.8 - 13 at a higher stress region at 625oC and 650oC. On the other hand, the

stress exponent is about 3.9 5.3 at a lower stress region at 625oC and 650oC. The

change of stress exponent at 625oC and 650oC may indicate transition of the creep

deformation property. For discussion of the creep deformation mechanism, it is also

necessary to determine the apparent activation energy of creep. The relationship

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between minimum creep rate and temperature are shown as an Arrhenius plot in Fig.4.

The magnitude of slope means apparent activation energy of creep deformation. The

activation energy for each applied stress is also shown in Fig.4. In heat A and B, the

apparent activation energy is respectively 350 420 kJ / mol (QH) and 544 765 kJ /

mol (QL) for a higher stress region and a lower one, respectively. The apparent

activation energy is about 342 431 kJ / mol (Q) for all stress regions in heat C. The

activation energy for lattice self diffusion of ferritic steel (QC) is about 350 kJ / mol.[16]

In heat A and B, QH is consistent with QC. The Q value is almost the same as QCin

heat C. The reason for the higher apparent activation energy at a lower stress region

in heat A and B is not clear. The deformation mechanism may change with stress in

heat A and B due to the apparent activation energy change shown in Fig.4. On the

other hand, it is expected that the deformation mechanism does not change in heat C

in terms of the apparent activation energy although the stress exponent changed at alower stress.

Consequently, it is confirmed that the creep deformation behavior changes at a

low stress at 625oC and 650oC in all steels, considering the changes of apparent

activation energy and stress exponent. The deformation behavior change, however,

does not directly mean deformation mechanism change. For deformation mechanism

change, a microstructure at minimum creep rate should be almost the same at each

stress.[17] However, microstructures such as precipitates, dislocation density and

bainitic lath structure remarkably change during creep in actual heat resistant steels.Time for minimum creep rate under a lower stress is longer than that under a higher

stress, indicating that under a lower stress the microstructure can change remarkably

until creep deformation reaches minimum creep rate, compared to a higher stress

condition. This tendency is more pronounced at a higher temperature. The minimum

creep rate under low stresses will be larger than that expected from the stress

exponent in higher stresses since deformation resistance at minimum creep rate may

be reduced under low stresses due to the remarkable change of microstructure. This

can be one of the reasons for the change of stress exponent and apparent activation

energy under a lower stress shown in Fig.3 and Fig.4.

3.3 Effect of creep deformation behavior change on creep rupture behavior

The time to rupture is plotted against minimum creep rate for all steels in Fig.5. A

linear relation, that is the Monkman-Grant rule[18], was observed. The relation

depends on temperature. The magnitude of slope in the relation(Fig.5-a, b, c) at

500oC and 550oC is different from those at other temperatures. In heat A and B, all of

the experimental data at 575oC and 600oC, and the short-term data at 625oC and

650oC, were expressed as a linear relationship (Fig.5-d, e) . However, experimental

data for the long-term at 625oC and 650oC deviates from the linear relationship. On

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the other hand, there was no deviation from the linear relationship in the long-term at

575oC to 650oC in heat C as shown in Fig.5-f. In the case of heats A and B, actual time

to rupture in the long-term at 625oC and 650oC is shorter than that extrapolated by

short term data (higher minimum creep rate). This suggests that increase in creep rate

at tertiary stage is larger and/or low creep ductility in the long-term (under low

stresses), comparing with the short-term (under high stresses). In short, it is possible

to predict directly time to rupture by minimum creep rate in heat C, while we can not

predict time to rupture at long-term by only minimum creep rate in heat A and B.

3.4 Creep ductility

The elongation for all steels is shown in Fig.6. In heat A and B, the elongation at

500oC and 550oC is kept high even in the long-term, while ductility decreases with

increasing time to rupture at 575o

C to 650o

C. At 625o

C and 650o

C, the ductilityrecovers in the long-term again. On the other hand, no drop in ductility was observed

in heat C for the whole temperature range examined. The drop in ductility of heat A

and B may be caused by brittle fractures such as intergranular fractures. The

Monkman-Grant rule is expressed as follow.

Ctrm & (1)where m& , tr, and Care minimum creep rate, time to rupture and constant. It can be

assumed that the constant corresponds to creep ductility. At 625oC and 650oC, the

constant C should be higher in the short-term since the creep ductility is high in theshort-term as shown in Fig.6. On the other hand, the constant C would be lower in the

long-term due to the low creep ductility. In short, the deviation of data points from the

Monkman-Grant rule in the-long term shown in Fig.5 is due to the decrease in creep

ductility.

Figure 7 shows optical micrographs of gauge portions, which are uniformly

deformed and far from necked regions, after creep at 550oC. No creep cavities were

observed in heat A, B and C after short-term(Fig.7(a), (c), (e) ) and long-term

creep(Fig.7(b), (d), (f)). We confirmed in heat A, B and C that creep cavities were

located in prior austenite grains of necked regions after short-term and long-term

creep at 550oC, meaning that the transgranular fracture occurs at 550oC. Figure 8

demonstrates optical micrographs of gauge portions, which are far from necked

regions, after creep at 625oC. A large number of creep cavities was located on prior

austenite grain boundaries in heats A (Fig.8(b), (c)) and B (Fig.8(e)) after long-term

creep although a small amount of cavities was observed after short-term creep

(Fig.8(a), (d)). This means that at 625oC, the fracture mode changes from

transgranular to intergranular with time to rupture in heat A and B. However, in heat B,

only a small amount of cavities was observed even after very long-term creep,

indicating recovery of ductility. (Fig.8(f)) In heat B, the change of number of cavities

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with time to rupture qualitatively agrees with that of ductility shown in Fig.6. While

there are no cavities after short-term (Fig.8(e)) and long-term creep (Fig.8(f)) in heat C.

In short, in the case of heat C, the transgranular fracture occurs even at 625oC

regardless of test conditions. It was reported that fracture mode change from

transgranular to intergranular fracture causes inflection in stress vs. time to rupture

curve for tempered martensitic steel.[19] However, note that in heat B, the creep

rupture strength at 625oC drops in the long-term although only a small amount of

cavities is observed after long-term creep as shown in Fig.8(f). In short, not only

fracture mode change but also other microstructural factors should be considered to

clarify the degradation mechanism of creep strength. The inflection in stress vs. time

to rupture curve is observed in heat C as shown in Fig.2. This indicates that in the

case of heat C, not the fracture mode change but other microstructural change

contributes to the inflection since the fracture mode change was not observed in heatC shown in Fig.8. The hardness of heat A and B is higher than that of heat C shown in

Table 2. The strength of the grain interior in heat A and B should be higher than that in

heat C since microstructural factors such as dislocation density and lath width, which

retain the strength of the grain interior, contribute to the hardness. In short, it is easy

for creep voids to form at prior austenite grain boundaries as an accommodation

process since the strength of the grain interior may be relatively larger than that of the

grain boundary in heats A and B, while in heat C, creep void formation may not be

required for the accommodation since the grain interior can easily deform due to thelower creep resistance.

We have summarized characteristics of creep deformation behavior and its effect

on creep rupture strength in Fig.9. In heat C, change of the stress exponent of

minimum creep rate directly contributes to the inflection in stress vs. time to rupture

curve. In heat A and B, actual data of creep life in the long-term deviates from the

trend expected from change of the stress exponent of minimum creep rate. The low

creep ductility causes the deviation, considering relationship between creep ductility

and the Monkman-Grant rule as discussed in section 3.4. It is expected that not

deformation mechanism change but microstructural change may cause the change in

the stress exponent of minimum creep rate in all heats as mentioned in section 3.2.

3.5 Comparison of creep rupture strength between all steels

Figure 10 shows creep rupture strength of all steels. The creep rupture strength

of heat A is clearly higher than those of heat B and heat C at 500oC to 600oC. At

625oC and 650oC, the creep rupture strength of heat A abruptly drops in the long-term,

compared with heat B and heat C. Therefore, the creep rupture strength of heat A

becomes similar to those of heat B and heat C in the long-term. The hardness of heat

C is lower than those of heat A and heat B in the as tempered condition as shown in

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Table 2. The difference of hardness can contribute to that of creep rupture strength

shown in Fig.10. However, the creep strength of heat A is higher than that of heat B

although the hardness of heat A is the same as that of heat B. There may be

difference in bainitic structure and/or precipitates between heat A and heat B. The

difference in bainitic structure and precipitates can cause that in creep rupture

strength. We have to clarify the detail of bainitic structure and precipitates for all heats

in future.

4. Conclusions

Creep tests were performed under several conditions in three steels for ASME

P23/T23. Creep rupture data and minimum creep rate were analyzed for evaluation of

creep strength degradation at long-term. The results can be summarized as follows.

(1) At 625oC and 650oC, the drop in creep rupture strength at long-term regions was

observed in all steels (heat A, B and C) . The drop in heat A and B with higher

initial hardness was more remarkable than that in heat C with lower hardness.

(2) At 625oC and 650oC, the stress exponent of minimum creep rate changed at a

lower stress in all steels (stress exponent : 7.8-13 for higher stress, 3.9-5.3 for

lower stress) . The stress region at which the stress exponent was changed is

consistent with that for the drop in creep rupture strength. In heat C, the

Monkman-Grant rule was confirmed over a wide temperature and stress rangeexamined. While actual data deviated from the rule at long-term in heat A and B,

the rule was confirmed at short-term.

(3) At 500oC and 550oC, the ductility was high even in the long-term in heats A and B,

while at 575oC to 650oC, the ductility decreased with increasing time to rupture.

The ductility increased at longer term again at 625oC and 650oC in heat A and B.

On the other hand, the ductility was high over a wide stress and temperature range

examined in heat C. The drop in ductility relates with intergranular fracture in heat

A and B. For heat A and B, the low ductility in the long-term caused the deviation of

experimental data from the Monkman-Grant rule.

(4) In heat C, the change of stress exponent of minimum creep rate directly

contributes to the drop in creep rupture strength since the Monkman-Grant rule

was confirmed. The stress exponent change may be attributed to microstructural

change. On the other hand, the drop in creep rupture strength was more

remarkable than that expected from both the stress exponent change and the

Monkman-Grant rule in heats A and B. The decrease in creep ductility in heats A

and B causes the deviation of data point from the Monkman-Grant rule. This

deviation contributes to the large drop in creep rupture strength in heats A and B.

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Acknowledgement

The authors would like to express their sincere gratitude to K. Yokokawa, O.

Kanemaru, K. Kubo, T. Ohba, Dr. H. Kushima, H. Miyazaki and all the members

involved in the Creep Data Sheet project for their collaboration in creep tests.

References

[1] F. Masuyama : ISIJ Int., 2001, vol.41, pp.612-625.

[2] F. Masuyama, T. Yokoyama, Y. Sawaragi and A. Iseda : Proc. on Materials for

Advanced Power Engineering, 1994, part I, pp.173-181.

[3] K. Kimura : Proc. of 2005 ASME Pressure Vessels and Piping Division Conference,

2005, PVP2005-71039.

[4] K. Kimura : Proc. of 2006 ASME Pressure Vessels and Piping Division Conference,2006, PVP2006-ICPVT11-93294.

[5] Thermal Power Standard Code, Ministry of Economy, Trade and Industry (METI),

Japanese Government, Tokyo, Dec. 14 (2005).

[6] Thermal Power Standard Code, Ministry of Economy, Trade and Industry (METI),

Japanese Government, Tokyo, July 10 (2007).

[7] M. Yoshizawa, M. Igarashi and A. Iseda : CAMP-ISIJ, 2005, vol.18, pp1549.

[8] M. Nakashiro, S. Kihara, Y. Tumita and I. Kajigaya : J. Soc. Mat. Sci., Japan, 1994,

vol. 43, pp.203-209.[9] R. Viswanathan and J. Foulds : Trans. ASME, J. Pressure Vessel Tech., 1998, vol.

120, pp.105-115.

[10] H. Kushima, K. Kimura and F. Abe : Tetsu-to-Hagane, 1999, vol. 85, pp.841-847.

[11] K. Kimura, K. Sawada, K. Kubo and H. Kushima : Proc. on Life Management and

Maintenance for Power Plants, 2004, vol.2, pp.465-476.

[12] K. Kimura, K. Sawada, K. Kubo and H. Kushima : Proc. of 2004 ASME/JSME

Pressure Vessel and Piping Conference, 2004, PVP2004-2566.[13] N. Komai and T. Imazato : Proc. of 8th Liege Conf. on Materials for Advanced

Power Engineering, 2006, part II, pp.997-1009.

[14] NIMS Creep Data Sheet, No.54, National Institute for Materials Science, Tsukuba,Japan, 2008.

[15] A. Iseda, H. Teranishi and F. Masuyama : Tetsu-to-Hagane, 1990, vol.76,

pp.1076-1083.

[16] K. Maruyama and H. Oikawa : J. Japan Inst. Metals, 1991, vol.55, pp.1189-1193.

[17] J. Cadek : Creep in Metallic Materials, Elsevier, Amsterdam, 1998.

[18] F. C. Monkman and N. J. Grant, Proc. ASTM, 1956, vol.56, pp.595.

[19] J. S. Lee, H. G. Armaki, K. Maruyama, T. Muraki and H. Asahi : Mater. Sci. Eng. A,2006, vol.A428, pp.270-275.

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Table 1 Chemical compositions(mass%) and heat treatment conditions.

C Si Mn P S Cr Mo W V

heat A (T23) 0.05 0.2 0.12 0.014 0.002 2.33 0.25 1.54 0.23

heat B (T23) 0.06 0.22 0.43 0.013 0.003 2.27 0.12 1.67 0.26

heat C (P23) 0.07 0.18 0.24 0.009 0.002 2.58 0.27 1.60 0.22

Nb Al N B Normalizing Tempering

heat A (T23) 0.05 0.002 0.007 0.0033 1050oC / 10min A.C. 770oC / 60min A.C.

heat B (T23) 0.06 0.002 0.004 0.0021 1050oC / 30min A.C. 770oC / 60min A.C.

heat C (P23) 0.04 0.009 0.006 0.0034 1050oC / 30min A.C. 770

oC / 60min A.C.

Table 3 Stress exponent of minimum creep rate.

625oC 650oC500oC 550oC 575oC 600oC

high stress low stress high stress low stress

heat A 7.9 6.7 18 12 13 3.9 9.0 4.5

heat B 7.7 8.0 13 9.0 7.8 4.2 9.6 4.2heat C 12 8.2 9.1 8.5 9.5 5.3 10 4.1

Vickers hardness0.2% proof stress

/ MPa

heat A 207 607

heat B 207 620

heat C 182 547

Table 2 Mechanical property at room temperature.

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heat A

50m

heat B

50m

heat C

50mFig.1 Optical micrographs before creep.

Fig.2 Creep rupture strength for all steels.

20

40

60

80

100

300

500

101

102

103

104

105

500oC

550oC

575oC

600oC

625oC

650oC

Stress/MPa

Time to rupture / h

ASME T23

heat A

20

40

60

80

100

300

500

101

102

103

104

105

500oC

550oC

575oC

600oC

625oC

650oC

Stress/MPa

Time to rupture / h

ASME T23

heat B

20

40

60

80

100

300

500

101 102 103 104 105

500oC

550oC

575oC

600oC

625oC

650oC

Stress/MPa

Time to rupture / h

ASME P23

heat C

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Fig.3 Relationship between minimum creep rate

and stress for all steels.

Fig.4 Relationship between minimum creep

rate and temperature for all steels.

10-7

10-6

10-5

10-4

10-3

10-2

20 40 60 80100 300 500

500oC

550

o

C575oC

600oC

625oC

650oC

Minimumcreeprate/h-1

Stress / MPa

ASME P23

heat C

10-7

10-6

10-5

10-4

10-3

20 40 60 80100 300 500

500oC

550oC

575oC

600o

C625oC

650oC

Minimumcreeprate

/h-1

Stress / MPa

ASME T23

heat A10

-7

10-6

10-5

10-4

10-3

20 40 60 80100 300 500

500oC

550oC

575oC

600oC

625oC

650oC

Minimumcreeprate

/h-1

Stress / MPa

ASME T23

heat B

10-7

10-6

10-5

10-4

10-3

10-2

0.001 0.0011 0.0012 0.0013 0.0014

200MPa (382kJ/mol)

180MPa (385kJ/mol)

160MPa (765kJ/mol)

140MPa (633kJ/mol)

100MPa (563kJ/mol)


T -1/ K -1

ASME T23

heat A

10-7

10-6

10-5

10-4

10-3

10-2

0.001 0.0011 0.0012 0.0013 0.0014

200MPa (374kJ/mol)

180MPa (350kJ/mol)

160MPa (420kJ/mol)

140MPa (552kJ/mol)100MPa (544kJ/mol)


T -1/ K -1

ASME T23

heat B

10-7

10-6

10-5

10-4

10-3

10-2

0.001 0.0011 0.0012 0.0013 0.0014



120MPa (431kJ/mol)

100MPa (371kJ/mol)

T -1/ K -1

ASME P23

heat C


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Fig. 5 Relationship between minimum creep rate and time to rupture.

Fig.6 Creep ductility for all steels.

10-7

10-6

10-5

10-4

10-3

10-2

101

102

103

104

105

575oC

600oC

625oC

650oC

Minimum creep rate / h-1

Timetorupture/h

ASME P23

heat C

-0.98

10-7

10-6

10-5

10-4

10-3

10-2

101

102

103

104

105

575oC

600oC

625oC

650oC


Timetorupture/h

ASME T23

heat A

-1.2

10-7

10-6

10-5

10-4

10-3

10-2

101

102

103

104

105

575oC

600oC

625oC

650oC


Timetorupture/h

ASME T23

heat B

-1.1

10-7

10-6

10-5

10-4

10-3

10-2

101

102

103

104

105

500oC

550oC


Timetorupture/h

ASME P23

heat C

-0.84

-1.0

10-7

10-6

10-5

10-4

10-3

10-2

101

102

103

104

105

500oC

550oC


Timetorupture/h

ASME T23

heat B

-0.68

-0.78

10-7

10-6

10-5

10-4

10-3

10-2

101

102

103

104

105

500oC

550oC


Timetorupture/h

ASME T23

heat A

-0.63

-0.83

(a) (b) (c)

(d) (e) (f)

0

10

20

30

40

50

60

101

102

103

104

105

500oC

550o

C

575oC

600o

C

625oC

650o

C

Elongation/%

Time to rupture / h

ASME P23

heat C

0

10

20

30

40

50

60

101

102

103

104

105

500o

C550oC

575o

C600oC

625o

C650oC

Elongation/%

Time to rupture / h

ASME T23

heat A0

10

20

30

40

50

60

101

102

103

104

105

500o

C550oC

575o

C600oC

625o

C650oC

Elongation/%

Time to rupture / h

ASME T23

heat B

8/13/2019 16 Sawada

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8/13/2019 16 Sawada

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m&log

m&log

m&log

m&log

log rtlog rtlog

log

Ktrm =&

log rtlog rtlog

log

Ktrm =&

Heat C at 625oC and 650oC

Heat A and B at 625oC and 650oC

Fig.9 Schematic illustration for effect of deformation behavior on rupture behavior.

Deviation fromthe Monkman-Grant rule

Fig.10 Comparison of creep rupture strength in all steels.

20

40

60

80

100

300

500

101

102

103

104

105

500oC

550oC

575oC

600oC

625oC

650oC

Stress/MPa

Time to rupture / h

Black : heat A

Red : heat B

Blue : heat C

16 Sawada

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