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Some notes on the structural design code for FBR components
Int. J. Pres. lies. & Piping 37 (1989) 83-92
Some Notes on the Structural Design Code for FBR Components
Hiroshi Wada a & Masanori Tashimo b Mitsubishi Heavy Industries, Ltd, Kobe Shipyard & Engine Works, a Advanced Reactor Division, b Advanced Nuclear Power Plant Engineering Dept, 1-1-1 Wadasaki-cho, Hyogo-ku, Kobe 652, Japan
ABSTRACT Several codes and standards have been developed and used in different countries for structural design of elevated temperature components of breeder reactors,for example, the A S M E Boiler and Pressure Vessel Code Case N-47 in the United States, R C C - M R in France, the Monju Design Code in Japan, and so on. These codes and standards have been issued in order to be applied to fast breeder reactor components; however, their use has been extended on the basis of a concept adopted in the A S M E Boiler and Pressure Vessel Code, Section III, which has the special aspect that its main subject is the prevention of bursting of vessels and pipes from internal pressure. The authors consider that a design code for breeder reactor components should take into account the special features of the components in geometric and/or loading conditions, i.e. thin wall structure, low internal pressure and high thermal loading. In this paper the authors make some proposals regarding future directions for the codes in their advanced form.
1 INTRODUCTION
Several codes and standards have been developed and used in different countries for the design of elevated temperature components o f breeder reactors. The A S M E Boiler and Pressure Vessel Code Case N-47 in the United States is the first code which fully takes into consideration the creep effects for various failure modes. And in France, R C C - M R for the SPX-II plant has recently been published. This code has features for strain limits and 83 Int. J. Pres. lies. & Piping 0308-0161/89/$03"50 © 1989 Elsevier Science Publishers Ltd, England. Printed in Great Britain
84
Hiroshi Wada, Masanori Tashimo
buckling evaluation comparable to ASME N-47. In Japan a prototype breeder reactor 'Monju' is now under construction. For this plant the Monju Design Code was developed by Power Reactor and Nuclear Fuel Development Corporation (PNC) with the co-operation of the fabricators, and the main components.of the Monju reactor were designed according to this code. The characteristics of the present codes are according to the design of plants to which codes are expected to be applied. For future plants design codes need to be modified to match the plant design. Here we make some proposals regarding future directions for the codes for fast breeder reactor (FBR) plants, taking into consideration the general features of the FBR components, i.e. thin wall structure, low internal pressure and high thermal loading based on the Monju reactor vessel (RV), steam generator (SG), core internal (CI) design and fabrication experience. Proposals for individual limits of design codes, i.e. the primary stress limits, creep fatigue evaluation and the strain limits are given in the following sections.
2 P R I M A R Y STRESS LIMITS Primary stress limits provide the means of preventing plastic instabilities and plastic collapses. For this purpose present codes limit the primary stress by using S m values, which are defined from the yield stress and ultimate tensile stress of the materials. In these limits, the codes allow the primary stress to exceed the yield stress considerably, especially for level D condition and austenitic stainless steels which have large hardening features. This limit seems to be based on the bursting of cylindrical shells by the internal pressure, which is a typical example of plastic instability. This instability is due to the decrease of thickness and the increase of radius of the cylinder, and occurs at very large strains near the uniform elongation of the material. So, for this instability the yield stress of material does not have a large effect, and the primary stress can be allowed essentially until near the ultimate tensile strength. However, another type of instability, usually called plastic collapse, has somewhat different aspects. One of the most typical examples of this type of instability is the collapse of cylindrical shells from axial compression forces. This collapse is due to the radial deformation accompanied by the increase of the tensile hoop strain, and occurs when the strain is very small, near the yield strain, compared with the uniform elongation. Figure 1 shows the ratios of elastically calculated primary stress (from the collapse loads) against the yield stresses of the materials for several structures and loading conditions. Collapse loads are determined by
Structural design code for FBR components
85
2.0 ~el
~y
1.5
1.0
o
0.5 O
0.0
I
I
I
I
I
A
B
C
D
E
Fig. 1. Primary stresses from collapse load. A, Cylinder (R/t = 80) axial compression cal. B, cylinder (R/t = 100) bending cal. C, cylinder (R/t = 200) transverse shear exp. D, complex structure (cylinder-cone conjunction) axial compression cal. E, complex structure (cylinder-cone conjunction) axial compression exp. O, General Membrane. O, Local Membrane.
14.0
12.0
o
,
*
10.0
o
o
o
Cylinder
Axial
Compress
i on
(Case
O
o
A)
.¢
,~
8.0
O
L
"~
6.0
c v
O Z
V
4.0
v
v
v
v
v
9
v
v
v
*
•
~
v
v
Axial
Conjunction Compression
(Case
O)
Cylinder-Cone v
2.0
0"O.0
0.1
I
J
I
t
0.2
0.3
0.4
0.5
Nominal
Fig. 2.
Strain
(/q~)
Load-deformation curve.
0.6
Hiroshi Wada, Masanori Tashimo
86
calculation (elastic-plastic large deformation analyses) or by experiments. Figures 2 and 3 show details for cases A and D. Figure 2 shows load against deformation relations; and Fig. 3 shows maximum strain in the structure against deformation relations. For the cases shown in these figures, the ratios are nearly all between 0.5 and 1"0 for general membrane stresses and between 1.5 and 2-0 for local membrane stresses. For cases D and E, we classified the bending stresses at the connective part as secondary stresses, and so the primary stresses were calculated without expecting bending stiffness at the connective part. Figure 3 shows that plastic collapse can occur at relatively small strains, near 1%. As shown in these figures, it seems to be necessary to limit the primary general membrane stresses by a value near the yield stress, and primary local membrane stresses by 1"5 or 2"0 times yield stress, if we intend to prevent plastic collapse by using primary stress limits. Primary stresses which considerably exceed the yield stress may cause plastic collapse. This problem has importance for design codes for FBR components because they are relatively thin structures, and the inner pressure is low enough that bending moment or shear force from earthquake forces may be
*
*
1.0
Cylinder
Axial C o m p r e s s i o n ( C a s e A )
Cyl i n d e r - C o n e Conjunction Axial Compression (Case D ) (
•
max. load
•
)
~ ~
0.5
?_
0.0 #a°
a
o O
r- Z O 0
E = -0.5 E
O O
=E -1.0
O
e
- 1.5
- 2.0 0.0
o11
o'.z
o'.3
Nominal Fig. 3.
Maximum
Strain
o14
0.6
(~)
strain-deformation
curve.
Structural design code for FBR components
87
critical loads for them. And the yield stress seems to have a more important role in primary stress limits of FBR design codes than in the codes for more usual pressure vessels.
3 CREEP FATIGUE EVALUATION Creep fatigue limits are usually the most critical criteria for the structural design of FBR components, especially for design by elastic analysis. For creep fatigue evaluation, almost all design codes adopt the linear damage rule (LDR). This rule has the great advantage of easy application due to its simplicity, and seems to be able to explain the results of creep fatigue tests to some extent. However, it seems to be necessary to improve or to confirm the rule for some items. The first item is extrapolation to long hold time and small strain range for design use. In addition to the linear damage rule, the strain range partitioning (SRP) method is widely used for creep fatigue evaluation, especially in research. It has some problems for application to design, i.e. for complex load histograms and multiaxial stress states. But is appears that SRP can explain the results of creep fatigue tests as well as, or better than, LDR. Figures 4 and 5 show the calculated results of hold time effects on fatigue life (life reduction factor). These calculations are based on a fixed creep equation, creep rupture time, fatigue curve (PP and CP curves for SRP) and cyclic stress-strain curve. Figure 4 shows the results when no elastic 10 °
~.o
\Z Z'*" 10 -l o v-
tl L
10 -z
~t=O.3~ .J
LDR
SRP
Et=l.0~
10-~
I
I
I
I
10'
10 ~
10 ~
10 3
Hold
Fig. 4.
time
(hP)
Life reduction factor calculated by LDR and SRP (q = l).
I
10'
Hiroshi Wada, Masanori Tashimo
88 10 0 o Z \ z~
10 ~
--------_5
o 1= o L
2
10 -=
.J t =0.3% t=
1.0%
1 0 -3 10 0
101
10 ~ Hold
Fig.
5.
time
10 ~
10'
(hr)
Life reduction factor calculated by L D R and SRP (q = 3).
follow up occurs in the stress relaxation, as in the usual creep tests; and Fig. 5 shows the results when some elastic follow up, described by q = 3, is considered. Here q means the ratio of the total creep strain to the creep strain effective to the stress relaxation. F r o m these figures it can be seen that L D R and S R P give similar reduction factors for the region of the usual creep fatigue tests, i.e. Ae - 1%, hold time -~ 1 h, and no elastic follow up. The difference between L D R and S R P is about 1-5 in life. However, for design applications, smaller strain ranges and longer hold times must be evaluated. F o r example, the difference reaches a factor of 10 for Ae = 0 . 3 % , hold time = 1000h and q = 3. So L D R is possibly very conservative for the region of design applications. The criteria for creep fatigue evaluation need to be investigated from the point of applicability to the design region, and for that purpose, it is necessary to make clear the physical mechanism of the damage of materials, in addition to long duration creep fatigue tests. The second point concerns material properties which could be a measure of the creep fatigue strength. If we use L D R for creep fatigue evaluation, it seems that the creep rupture strength is the measure. O f course creep damage in L D R depends not only on creep rupture time but also on stress relaxation rate and the stress-strain curve; but the rupture strength seems to be the most important property of materials for structural design of F B R components, especially when design by elastic analysis is used. However, it may be more suitable, for example, to use the creep rupture elongation than to use creep rupture strength as a measure of creep fatigue strength, since the rupture elongation can be a measure of low cycle fatigue strength.
Structural design code for FBR components
89
3.0
2.0
1.0 Q; m L ¢:
0.5 316MN, SUS316, 550°c
~
0.3
ContinuousH o l d
Time 10min.
Cycle
g I,-
0.2
0.1
a
i
i
SUS316
©
•
316MN
A
•
k
i
;
101
IILII
i
i
r
,
J
i
ilaL]
10 2
J
J
i
i
10~
i
L
l l l l J
i
i
i
i
i
i
10"
i
i
10 5
Number of cycles t o f a i l u r e , Nf (cycles) Fig. 6. Hold time effects for SUS 316 and 316 MN.
As one such example, Fig. 6 shows a comparison of hold time effects for two materials. One is the usual 316 SS (SUS 316) and the other is a 316 type, but low carbon, medium nitrogen stainless steel (316 MN). As seen in this figure, SUS 316 shows a life reduction by the hold time of 10 minutes, but 316 MN shows no reduction. Comparisons of rupture time and rupture 50
40 o
E
o
O
30
O.
o z,,
o 316MN
• SUS316
20
I
I
I
102
10 ~
10'
Rupture
Fig. 7.
time
(hr)
Creep rupture time.
90
Hiroshi Wada, Masanori Tashimo 100 o 316MN • SUS310 80-
60
¢= o =
40 O
B bJ
O
~0
d
I
102
I
10 ~ Rupture
time
10' (hr)
Fig. 8. Creep rupture elongation.
elongation between these materials are shown in Figs 7 and 8. The 316 M N shows a somewhat longer rupture time, especially for low stresses, but the difference in creep rupture elongation between the two materials seems to be more definite than that in creep time. This may represent the difference in the life reduction by hold time. Of course it is not possible to reach a definite conclusion from this example. But it is preferable that the rupture criteria of a design code reflect directly the material properties which represent the strength for that rupture mode, and this could be useful for material selection in design. From such a point of view it seems to be desirable to improve the criteria for creep fatigue evaluation.
4 STRAIN LIMITS As for strain limits, both methods of evaluating the strain and the limit value seem to be the subject of reconsideration for future design codes. Present design codes adopt considerably different methods to evaluate the ratcheting strain. ASME N-47 uses the O'Donnell and Porowski model which is an extension of the Bree model to creep region temperature. This model is a theoretical model for elastic-full plastic material loaded with constant membrane and cyclic bending forces. N-47 limits the combination of primary and secondary stresses to within the region in which this model does not show ratcheting behavior. R C C - M R adopts a quite different method. It is based on experimental
Structural design code for FBR components
91
results for cylinders loaded with constant axial and cyclic torsion forces. The Monju design code also adopts the O'Donnell and Porowski model, but it additionally limits cyclic secondary stresses to a specified level even if the primary stress is low enough. These differences seem to be a reflection of the present situation that no model can properly explain the ratcheting phenomena which occur in FBR components. Figure 9 shows comparisons of the ratchet regions which some methods give for the combination of primary and secondary stresses. As shown in this figure, the differences between methods are very large, especially for low primary stresses. For the reasonable design of FBR components, a method should be developed to evaluate ratcheting strain behaviors for low primary stresses. The ratcheting mechanism is very complex, and a simple model may not
-----
40
--
\ E
Bree model (tension+bending) -- Tens ion+ToPsion No ratchet ing
......
Tension+Torsion A ¢ = 1 % / 1 O0 c y c l e
-
RCC - M R
-
exper
i menta
A~ =0.45%
\
30
Q~ w z.
20
a ¢ o o
10 Sus318
( 400°c )
S y = 12.5 k g / m m 2 Sin=
11.2 k g / m m 2
I
I
5
10
Primary
Fig. 9.
Stress
(kc/mm
2)
Comparisons of ratchet regions.
I
/
92
Hiroshi Wada, Masanori Tashimo
be able to explain all kinds of ratcheting. Therefore, it may be necessary to limit the method to some critical parts of FBR components, and to investigate the ratcheting mechanism which occurs in these parts both theoretically and experimentally. Contrary to the methods above, present codes adopt approximately the same limit value for strain limits. However, it seems that the meaning of these values, particularly in relation to the failure due to ratcheting is not clear. Failure may be due to excessive strain itself or to acceleration of another mode of failure (e.g. fatigue or buckling failure) by ratcheting. A new limit should be proposed based on this feature.
Some Notes on the Structural Design Code for FBR Components
Hiroshi Wada a & Masanori Tashimo b Mitsubishi Heavy Industries, Ltd, Kobe Shipyard & Engine Works, a Advanced Reactor Division, b Advanced Nuclear Power Plant Engineering Dept, 1-1-1 Wadasaki-cho, Hyogo-ku, Kobe 652, Japan
ABSTRACT Several codes and standards have been developed and used in different countries for structural design of elevated temperature components of breeder reactors,for example, the A S M E Boiler and Pressure Vessel Code Case N-47 in the United States, R C C - M R in France, the Monju Design Code in Japan, and so on. These codes and standards have been issued in order to be applied to fast breeder reactor components; however, their use has been extended on the basis of a concept adopted in the A S M E Boiler and Pressure Vessel Code, Section III, which has the special aspect that its main subject is the prevention of bursting of vessels and pipes from internal pressure. The authors consider that a design code for breeder reactor components should take into account the special features of the components in geometric and/or loading conditions, i.e. thin wall structure, low internal pressure and high thermal loading. In this paper the authors make some proposals regarding future directions for the codes in their advanced form.
1 INTRODUCTION
Several codes and standards have been developed and used in different countries for the design of elevated temperature components o f breeder reactors. The A S M E Boiler and Pressure Vessel Code Case N-47 in the United States is the first code which fully takes into consideration the creep effects for various failure modes. And in France, R C C - M R for the SPX-II plant has recently been published. This code has features for strain limits and 83 Int. J. Pres. lies. & Piping 0308-0161/89/$03"50 © 1989 Elsevier Science Publishers Ltd, England. Printed in Great Britain
84
Hiroshi Wada, Masanori Tashimo
buckling evaluation comparable to ASME N-47. In Japan a prototype breeder reactor 'Monju' is now under construction. For this plant the Monju Design Code was developed by Power Reactor and Nuclear Fuel Development Corporation (PNC) with the co-operation of the fabricators, and the main components.of the Monju reactor were designed according to this code. The characteristics of the present codes are according to the design of plants to which codes are expected to be applied. For future plants design codes need to be modified to match the plant design. Here we make some proposals regarding future directions for the codes for fast breeder reactor (FBR) plants, taking into consideration the general features of the FBR components, i.e. thin wall structure, low internal pressure and high thermal loading based on the Monju reactor vessel (RV), steam generator (SG), core internal (CI) design and fabrication experience. Proposals for individual limits of design codes, i.e. the primary stress limits, creep fatigue evaluation and the strain limits are given in the following sections.
2 P R I M A R Y STRESS LIMITS Primary stress limits provide the means of preventing plastic instabilities and plastic collapses. For this purpose present codes limit the primary stress by using S m values, which are defined from the yield stress and ultimate tensile stress of the materials. In these limits, the codes allow the primary stress to exceed the yield stress considerably, especially for level D condition and austenitic stainless steels which have large hardening features. This limit seems to be based on the bursting of cylindrical shells by the internal pressure, which is a typical example of plastic instability. This instability is due to the decrease of thickness and the increase of radius of the cylinder, and occurs at very large strains near the uniform elongation of the material. So, for this instability the yield stress of material does not have a large effect, and the primary stress can be allowed essentially until near the ultimate tensile strength. However, another type of instability, usually called plastic collapse, has somewhat different aspects. One of the most typical examples of this type of instability is the collapse of cylindrical shells from axial compression forces. This collapse is due to the radial deformation accompanied by the increase of the tensile hoop strain, and occurs when the strain is very small, near the yield strain, compared with the uniform elongation. Figure 1 shows the ratios of elastically calculated primary stress (from the collapse loads) against the yield stresses of the materials for several structures and loading conditions. Collapse loads are determined by
Structural design code for FBR components
85
2.0 ~el
~y
1.5
1.0
o
0.5 O
0.0
I
I
I
I
I
A
B
C
D
E
Fig. 1. Primary stresses from collapse load. A, Cylinder (R/t = 80) axial compression cal. B, cylinder (R/t = 100) bending cal. C, cylinder (R/t = 200) transverse shear exp. D, complex structure (cylinder-cone conjunction) axial compression cal. E, complex structure (cylinder-cone conjunction) axial compression exp. O, General Membrane. O, Local Membrane.
14.0
12.0
o
,
*
10.0
o
o
o
Cylinder
Axial
Compress
i on
(Case
O
o
A)
.¢
,~
8.0
O
L
"~
6.0
c v
O Z
V
4.0
v
v
v
v
v
9
v
v
v
*
•
~
v
v
Axial
Conjunction Compression
(Case
O)
Cylinder-Cone v
2.0
0"O.0
0.1
I
J
I
t
0.2
0.3
0.4
0.5
Nominal
Fig. 2.
Strain
(/q~)
Load-deformation curve.
0.6
Hiroshi Wada, Masanori Tashimo
86
calculation (elastic-plastic large deformation analyses) or by experiments. Figures 2 and 3 show details for cases A and D. Figure 2 shows load against deformation relations; and Fig. 3 shows maximum strain in the structure against deformation relations. For the cases shown in these figures, the ratios are nearly all between 0.5 and 1"0 for general membrane stresses and between 1.5 and 2-0 for local membrane stresses. For cases D and E, we classified the bending stresses at the connective part as secondary stresses, and so the primary stresses were calculated without expecting bending stiffness at the connective part. Figure 3 shows that plastic collapse can occur at relatively small strains, near 1%. As shown in these figures, it seems to be necessary to limit the primary general membrane stresses by a value near the yield stress, and primary local membrane stresses by 1"5 or 2"0 times yield stress, if we intend to prevent plastic collapse by using primary stress limits. Primary stresses which considerably exceed the yield stress may cause plastic collapse. This problem has importance for design codes for FBR components because they are relatively thin structures, and the inner pressure is low enough that bending moment or shear force from earthquake forces may be
*
*
1.0
Cylinder
Axial C o m p r e s s i o n ( C a s e A )
Cyl i n d e r - C o n e Conjunction Axial Compression (Case D ) (
•
max. load
•
)
~ ~
0.5
?_
0.0 #a°
a
o O
r- Z O 0
E = -0.5 E
O O
=E -1.0
O
e
- 1.5
- 2.0 0.0
o11
o'.z
o'.3
Nominal Fig. 3.
Maximum
Strain
o14
0.6
(~)
strain-deformation
curve.
Structural design code for FBR components
87
critical loads for them. And the yield stress seems to have a more important role in primary stress limits of FBR design codes than in the codes for more usual pressure vessels.
3 CREEP FATIGUE EVALUATION Creep fatigue limits are usually the most critical criteria for the structural design of FBR components, especially for design by elastic analysis. For creep fatigue evaluation, almost all design codes adopt the linear damage rule (LDR). This rule has the great advantage of easy application due to its simplicity, and seems to be able to explain the results of creep fatigue tests to some extent. However, it seems to be necessary to improve or to confirm the rule for some items. The first item is extrapolation to long hold time and small strain range for design use. In addition to the linear damage rule, the strain range partitioning (SRP) method is widely used for creep fatigue evaluation, especially in research. It has some problems for application to design, i.e. for complex load histograms and multiaxial stress states. But is appears that SRP can explain the results of creep fatigue tests as well as, or better than, LDR. Figures 4 and 5 show the calculated results of hold time effects on fatigue life (life reduction factor). These calculations are based on a fixed creep equation, creep rupture time, fatigue curve (PP and CP curves for SRP) and cyclic stress-strain curve. Figure 4 shows the results when no elastic 10 °
~.o
\Z Z'*" 10 -l o v-
tl L
10 -z
~t=O.3~ .J
LDR
SRP
Et=l.0~
10-~
I
I
I
I
10'
10 ~
10 ~
10 3
Hold
Fig. 4.
time
(hP)
Life reduction factor calculated by LDR and SRP (q = l).
I
10'
Hiroshi Wada, Masanori Tashimo
88 10 0 o Z \ z~
10 ~
--------_5
o 1= o L
2
10 -=
.J t =0.3% t=
1.0%
1 0 -3 10 0
101
10 ~ Hold
Fig.
5.
time
10 ~
10'
(hr)
Life reduction factor calculated by L D R and SRP (q = 3).
follow up occurs in the stress relaxation, as in the usual creep tests; and Fig. 5 shows the results when some elastic follow up, described by q = 3, is considered. Here q means the ratio of the total creep strain to the creep strain effective to the stress relaxation. F r o m these figures it can be seen that L D R and S R P give similar reduction factors for the region of the usual creep fatigue tests, i.e. Ae - 1%, hold time -~ 1 h, and no elastic follow up. The difference between L D R and S R P is about 1-5 in life. However, for design applications, smaller strain ranges and longer hold times must be evaluated. F o r example, the difference reaches a factor of 10 for Ae = 0 . 3 % , hold time = 1000h and q = 3. So L D R is possibly very conservative for the region of design applications. The criteria for creep fatigue evaluation need to be investigated from the point of applicability to the design region, and for that purpose, it is necessary to make clear the physical mechanism of the damage of materials, in addition to long duration creep fatigue tests. The second point concerns material properties which could be a measure of the creep fatigue strength. If we use L D R for creep fatigue evaluation, it seems that the creep rupture strength is the measure. O f course creep damage in L D R depends not only on creep rupture time but also on stress relaxation rate and the stress-strain curve; but the rupture strength seems to be the most important property of materials for structural design of F B R components, especially when design by elastic analysis is used. However, it may be more suitable, for example, to use the creep rupture elongation than to use creep rupture strength as a measure of creep fatigue strength, since the rupture elongation can be a measure of low cycle fatigue strength.
Structural design code for FBR components
89
3.0
2.0
1.0 Q; m L ¢:
0.5 316MN, SUS316, 550°c
~
0.3
ContinuousH o l d
Time 10min.
Cycle
g I,-
0.2
0.1
a
i
i
SUS316
©
•
316MN
A
•
k
i
;
101
IILII
i
i
r
,
J
i
ilaL]
10 2
J
J
i
i
10~
i
L
l l l l J
i
i
i
i
i
i
10"
i
i
10 5
Number of cycles t o f a i l u r e , Nf (cycles) Fig. 6. Hold time effects for SUS 316 and 316 MN.
As one such example, Fig. 6 shows a comparison of hold time effects for two materials. One is the usual 316 SS (SUS 316) and the other is a 316 type, but low carbon, medium nitrogen stainless steel (316 MN). As seen in this figure, SUS 316 shows a life reduction by the hold time of 10 minutes, but 316 MN shows no reduction. Comparisons of rupture time and rupture 50
40 o
E
o
O
30
O.
o z,,
o 316MN
• SUS316
20
I
I
I
102
10 ~
10'
Rupture
Fig. 7.
time
(hr)
Creep rupture time.
90
Hiroshi Wada, Masanori Tashimo 100 o 316MN • SUS310 80-
60
¢= o =
40 O
B bJ
O
~0
d
I
102
I
10 ~ Rupture
time
10' (hr)
Fig. 8. Creep rupture elongation.
elongation between these materials are shown in Figs 7 and 8. The 316 M N shows a somewhat longer rupture time, especially for low stresses, but the difference in creep rupture elongation between the two materials seems to be more definite than that in creep time. This may represent the difference in the life reduction by hold time. Of course it is not possible to reach a definite conclusion from this example. But it is preferable that the rupture criteria of a design code reflect directly the material properties which represent the strength for that rupture mode, and this could be useful for material selection in design. From such a point of view it seems to be desirable to improve the criteria for creep fatigue evaluation.
4 STRAIN LIMITS As for strain limits, both methods of evaluating the strain and the limit value seem to be the subject of reconsideration for future design codes. Present design codes adopt considerably different methods to evaluate the ratcheting strain. ASME N-47 uses the O'Donnell and Porowski model which is an extension of the Bree model to creep region temperature. This model is a theoretical model for elastic-full plastic material loaded with constant membrane and cyclic bending forces. N-47 limits the combination of primary and secondary stresses to within the region in which this model does not show ratcheting behavior. R C C - M R adopts a quite different method. It is based on experimental
Structural design code for FBR components
91
results for cylinders loaded with constant axial and cyclic torsion forces. The Monju design code also adopts the O'Donnell and Porowski model, but it additionally limits cyclic secondary stresses to a specified level even if the primary stress is low enough. These differences seem to be a reflection of the present situation that no model can properly explain the ratcheting phenomena which occur in FBR components. Figure 9 shows comparisons of the ratchet regions which some methods give for the combination of primary and secondary stresses. As shown in this figure, the differences between methods are very large, especially for low primary stresses. For the reasonable design of FBR components, a method should be developed to evaluate ratcheting strain behaviors for low primary stresses. The ratcheting mechanism is very complex, and a simple model may not
-----
40
--
\ E
Bree model (tension+bending) -- Tens ion+ToPsion No ratchet ing
......
Tension+Torsion A ¢ = 1 % / 1 O0 c y c l e
-
RCC - M R
-
exper
i menta
A~ =0.45%
\
30
Q~ w z.
20
a ¢ o o
10 Sus318
( 400°c )
S y = 12.5 k g / m m 2 Sin=
11.2 k g / m m 2
I
I
5
10
Primary
Fig. 9.
Stress
(kc/mm
2)
Comparisons of ratchet regions.
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Hiroshi Wada, Masanori Tashimo
be able to explain all kinds of ratcheting. Therefore, it may be necessary to limit the method to some critical parts of FBR components, and to investigate the ratcheting mechanism which occurs in these parts both theoretically and experimentally. Contrary to the methods above, present codes adopt approximately the same limit value for strain limits. However, it seems that the meaning of these values, particularly in relation to the failure due to ratcheting is not clear. Failure may be due to excessive strain itself or to acceleration of another mode of failure (e.g. fatigue or buckling failure) by ratcheting. A new limit should be proposed based on this feature.