PVP2009-77279

Proceedings of the ASME 2009 Pressure Vessels and Piping Division Conference
PVP2009
July 26-30, 2009, Prague, Czech Republic
Proceedings of PVP2009
2009 ASME Pressure Vessels and Piping Division Conference
July 26-30, 2009, Prague, Czech Republic
PVP2009-77279
PVP2009-77279
IASCC EVALUATION METHOD OF IRRADIATED COLD WORKED 316SS BAFFLE
FORMER BOLT IN PWR PRIMARY WATER
Kiyotomo NAKATA
Japan Nuclear Energy Safety
Organization (JNES), Japan
[email protected]
Kenichi TAKAKURA
Japan Nuclear Energy Safety
Organization (JNES), Japan
[email protected]
Noboru KUBO
Mitsubishi Heavy Industries
(MHI), Ltd., Japan
[email protected]
Koji FUJIMOTO
Mitsubishi Heavy Industries
(MHI), Ltd., Japan
[email protected]
ABSTRACT
Irradiation Assisted Stress Corrosion Cracking (IASCC) is
a matter of great concern as degradation of core internal
components in light water nuclear reactor. To clarify the
IASCC initiation conditions of baffle former bolt (BFB),
constant load stress corrosion cracking (SCC) tests were carried
out in simulated PWR primary water (290, 320, 340°C) using
C-ring type specimens.
Based on the SCC test results, IASCC initiation time
becomes shorter with increasing fluence and increasing applied
stress, IASCC initiation threshold stress becomes lower with
increasing fluence. A test temperature effect was observed in
SCC initiation time, but it was not clear the effect of test
temperature for SCC initiation threshold stress. These results
suggest that IASCC initiation threshold criteria can be
described with stress in specimen and fluence.
This paper describes the whole evaluation procedure to
secure structural integrity of irradiated baffle structure in PWR
primary environments, including the threshold stress diagram
of IASCC initiation and the irradiation creep formula.
.
INTRODUCTION
Austenitic stainless steels are widely used as structural
alloy in reactor pressure vessel internal components because of
their high strength, ductility, and fracture toughness. However,
exposure to neutron irradiation results in changes in
microstructure, mechanical properties and microchemistry of
the material.
Kimihisa SAKIMA
Mitsubishi Heavy Industries
(MHI), Ltd., Japan
[email protected]
IASCC caused by the effect of neutron irradiation
during long term plant operation in high temperature water
environments is considered to be one of the major concerns of
in-core structural materials for Light Water Reactors (LWR).
IASCC is recognized as the most important degradation
phenomena affecting the integrity of LWRs It takes the form of
intergranular stress corrosion cracking (IGSCC). Actually, the
baffle former bolts cracking have been caused occasionally in
the overseas plants, e.g. BUGEY-2 in France, since 1988.
Although such events are not found in the Japanese plants so
far, it cannot be denied that the same events might occur in
future. A number of studies on IASCC have been performed
over the last two decades however the mechanism of IASCC is
not fully understood. Furthermore, practical engineering
databases about IASCC to evaluate the core shroud’s internal
integrity are of an insufficient amount.
Japan Nuclear Energy Safety organization (JNES) has
been conducting a project related to IASCC as part of safety
research & development study for the aging management &
maintenance of the nuclear power plant. The objective of
IASCC project is to prepare an IASCC evaluation guide for the
regulatory side, which uses it to evaluate utility’s ageing
management technical evaluation report. The project was
begun in JFY 2000, completed in JFY 2008. The target
component is the BFB, as can be seen from Figure 1, the BFB
is the parts forming the baffle structure, witch composes the
core region and primary coolant passage. The SCC initiation
test in simulated PWR primary water had been conducted, on
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the bases of SCC test results, the structural integrity evaluation
guide for BFB is proposed as JNES document.
TEST PROCEDURE
Materials
IASCC lifetime evaluation should preferably be carried out
for BFBs from actual plants, however, these BFBs have today a
fluence of only 20dpa at highest (corresponding to operating
period of 20 years). BFB material irradiated up to fluence of
60dpa (corresponding to operating period of 60 years) are not
available. Therefore, flux thimble tubes (FTTs), which are a
tube in which a neutron flux measurer, and located in the center
of a reactor, is used as sample instead of BFBs. The FTTs are
made of type 316CW (cold worked) stainless steel similar to
BFBs, but it has a higher fluence than BFBs due to different
location in a reactor. For this reason, much research work on
IASCC, using FTTs have been performed and reported [1-10].
In this study, constant load SCC tests have been conducted
for both irradiated FTTs (maximum 70dpa) and BFBs
(maximum 20dpa) to evaluate IASCC lifetime of BFBs at high
irradiation level.
Chemical compositions for test materials are shown in
Table 1, and condition of cold working and grain size are
shown in Table 2. The degree of cold work is 20% for BFBs,
and 12% for FTTs. Additionally, there is a remarkable
difference in the grain size of the two materials. The FTTs have
a finer grain size, which can be attributed to recrystallization
caused by repeated cold working and heat treatments.
SCC Test Pieces
Constant load SCC test is required for assessment of
IASCC lifetime. In this study, C-ring type specimens were
adopted for following three reasons:
(1) SCC initiation tests require relatively high applied stress.
It has been reported that stress above 0.2% yield strength of
material is required for SCC initiation.
(2) Specimen and test jig have to be as small as possible
considering handling of the materials in a hot cell laboratory.
(3) Bending stress at neck is dominant in an actual BFBs,
and test under bending stress is possible in the case of C-ring
test specimen in accordance with the actual case.
The shape of the C-ring type specimen is shown in Figure
2. The outer diameter is 7.65mm and the wall thickness is
1.27mm for all specimens. The length of the specimen is 10mm
for the FTTs, but 7 to 10mm for the BFBs due to limitations of
available material.
C-ring type specimens from the BFBs were machined from
head, neck under head, straight portions, and threaded rejoin of
the BFBs, and finished removing surface by mill scale. But Cring type specimens from the FTTs were prepared by simply
cutting the tube into prescribed length (10mm), and not
finished removing surface.
Two pin holes with diameter of 1.2mm are made in the
upper and lower parts of the specimen to fix the specimens
firmly into the test jig.
SCC Test Condition
Constant load SCC tests were conducted in simulated
PWR water environment. Test conditions are shown in Table 3.
The tests were conducted at 290, 320, and 340°C. The
maximum total test period was 2,000 hours.
Figure 3 is a schematic illustration of the test jig structure
with specimens installed. Three C-ring type specimens can be
tested simultaneously. SCC initiation time (rupture of C-ring
type specimen) was detected by a laser displacement sensor and
by a change in the indication of the load cell.
The applied stress in the C-ring type specimen was
evaluated by finite element method (FEM) analysis using
stress-strain curves obtained using tensile tests for each fluence
level in high temperature. Based on the FEM analysis result,
the applied load and stress was determined assuming that
specimen cracks initiation occurs on the outer surface of C-ring
type specimens.
RESULTS AND DISCUSSION
Constant load SCC tests were conducted using the C-ring
type specimens (Total 81 specimens; 27 specimens from the
BFBs and 54 specimens from the FTTs) to investigate the effect
of fluence, temperature, and stress to determine SCC initiation
time and SCC threshold stress.
Observation of fracture surfaces after constant load SCC
test
Figure 4 through 5 show results of stereo microscope
observations of outer surface cracks and SEM observations on
fracture surfaces near the outer surface of the C-ring type
specimens after the constant load SCC test.
The surface cracks in the BFBs specimen were observed a
more zigzag appearance while the surface cracks in the FTT
specimen were straighter. This difference is considered to be
caused by difference in grain size.
The fracture surfaces of all the specimens show typical
“intergranular” features.
Dependence of SCC initiation time on applied stress
Figure 6 shows the relationship between the SCC initiation
time and applied stress. The top, middle, and bottom sides in
Figure 6 corresponds to constant load SCC test results at 290,
320, and 340°C, respectively.
As the results, the SCC initiation time decreases with
increasing applied stress (σ) and fluence.
Dependence of SCC initiation stress on fluence and applied
stress
Figure 7 shows the results from SCC initiation tests and
describes the relationship between presence of SCC crack and
test conditions, i.e. applied stress and fluence. The top, middle,
and bottom sides in Figure 7 correspond to results at 290, 320,
and 340°C, respectively.
The result clearly indicates that the applied stress required
for SCC initiation decreases with increasing fluence.
Additionally, it was observed that SCC initiates even at a stress
of approximately 600MPa (0.6σy) in the case of a fluence of
40dpa, and only 400MPa (0.4σy) at a fluence of 70dpa. At
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70dpa, SCC occurs at 340°C but does not occur at 320°C, see
Figure 8.
Effect of test temperature on SCC initiation tests
The effect of test temperature on SCC initiation time and
SCC initiation threshold stress was analyzed based on the result
from the constant load SCC tests to show it in Figure 7.
Significantly difference in the “SCC initiation threshold
stress” between 290, 320, and 340°C was not observed. But the
effect of temperature on “SCC initiation time” was observed.
Therefore, I show the result that activation energy with a
representative SCC test results (FTTs) in Figure 9.
As a result, the activation energy was 26.3-34.7kcal/mol
(an average of 30.6kcal/mol).
Effect of material difference on SCC initiation
Effects of differences in material on SCC initiation
threshold stress were discussed in the following.
SCC threshold stress for BFBs and FTTs show significant
difference from Figure 8.
To evaluate the validity of lifetime evaluation for actual
plants BFBs using data obtained by testing FTTs, it is important
to investigate the effect of minor differences in materials on
IASCC characteristics in detail.
As a result, a difference was not recognized to tensile
strength, hardness with the bulk, RIS, quantity of irradiation
generation gas (helium and hydrogen).
However, irradiation hardenings of outer surface layer of
the C-ring type specimens were observed with the increase of
fluence to show it in Figure 10. In other words, the tendency
that the hardening of the outer surface layer before the
irradiation seemed to just increase was recognized in FTTs in
comparison with BFBs.
Therefore, we show the result that relationship between the
SCC initiation threshold stress and Vickers hardness of the
outer surface layer (about 2µm) for C-ring type specimens in
Figure 11. As a result, we understood that we could correlate
between SCC initiation threshold stresses and outer surface
layer hardness of the C-ring type specimens.
Therefore, for the difference of the SCC initiation
threshold stress of BFBs and FTTs, we confirmed what the
hardening on the outer surface layer of the C-ring type
specimens influenced.
IASCC EVALUATION METHOD
Life time evaluation concept
In the baffle structure, the core baffle is fixed to the core
baffle plate by many baffle former bolts, and the baffle plate is
installed to the core barrel by some barrel former bolts, i.e. the
functions of the baffle structure are maintained by many baffle
former bolts (Figure 1).
Therefore, even if a crack is generated in one of the baffle
former bolts, it does not cause the loss of the baffle structure
functions immediately.
However, when the number of
cracked bolts is increased, the function to retain the baffle plate
during seismic event is adversely affected. In other words, the
increase in the number of cracked bolts deteriorates the plant
safety and functions.
Thus, for the plant maintenance, the baffle structural
integrity has to be evaluated taking into account the number
and the locations of the bolts which are found to have the
potential of crack initiation in the evaluation.
Figure 13 illustrates the essential features of procedure
to specify the appropriate inspection period for the retention of
baffle structure functions.
For aging plant integrity
evaluation, the inspection period is determined by comparing
the number of bolts required for the retention of the baffle
structure functions with the number of the bolts with cracking
which is predicted by the evaluation method in this paper.
Considering SCC test results, IASCC initiation of BFB can
be described by material degradation index under neutron
irradiation and stress values in BFB. The unit of dpa
(displacement per atom) is adopted for the indicator of the
material degradation by neutron irradiation. The threshold
stress value of IASCC initiation decreases according to the
neutron irradiation increasing, therefore, for the life time
evaluation for aging PWR component, it is needed to consider
the BFB behaviors appropriately which vary from hour to hour.
Threshold stress diagram of IASCC initiation
According to the result of constant load SCC test, the
threshold stress for IASCC initiation of the BFBs and FTTs
were observed slightly different. We think that the surface
hardening layer influences it. However, the mechanism of such
difference has not yet been clarified enough. Therefore, for
conservative and practical evaluation, based on the neutron
irradiation index (dpa) and the stress in the BFB at the
evaluation time point, three regions according to the level of
crack initiation possibility are allocated to the diagram in
Figure 14. The each region is defined as follows:
(1) Case that BFB stress and fluence are included in Region 1
The BFB is considered to have no possibility of crack
initiation at the evaluation time point
(2) Case that BFB stress and fluence are included in Region 2
The BFB is considered to have a possibility of crack
initiation at the evaluation time point
(3) Case that BFB stress and fluence are included in Region 3
The BFB is considered to have a high possibility of crack
initiation at the evaluation time point
Stress evaluation
The stress in the BFB for the evaluation of the crack
initiation possibility is calculated as the cumulative changes
during the time from the start of operation to the evaluation
time point. The reason is that, the stress at the specific
evaluation time point is affected by the history of changes in
configuration and mechanical property of material according to
the increase in the neutron irradiation. The cause of the stress is
considered to be the following factors:
(1) The load factors before plant operation (the tightening
stress, etc.)
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(2) The thermal load during plant operation (the stresses
due to the structural deformation caused by the temperature
distribution, etc.)
(3) The load factors due to aging (the stresses due to the
structural deformation, e.g. swelling, caused by the increase in
the neutron irradiation level, the stresses due to the changes in
mechanical property caused by irradiation creep/neutron
irradiation, etc.)
For the irradiation creep estimation, JNES had been
conducted in-pile creep tests in Halden reactor since 2005 to
2006, and examined the irradiation creep rate formula (Figure
12).
When the neutron irradiation is proportional to operation
time, the stress change in the BFB according to the time history
is similar to that according to the neutron irradiation.
Evaluation procedure
To evaluate the crack initiation time of the evaluating BFB,
the following procedure is applicable:
(1) Conduct an evaluation to predict the stress history and
the neutron irradiation history of the BFB, then, based on the
results, estimate the stress history as a function of the fluence.
(2) Compare the stress history obtained in above step with
the threshold stress diagram of IASCC initiation (Figure 14),
then, specify to which region evaluating BFB corresponds.
(3) Adopt above two step evaluation to all the managing
BFBs, these are indispensable BFBs to keep baffles structure
function, then, the structural integrity of whole baffles structure
should be judged from the number of damage BFB.
SUMMARY AND CONCLUSIONS
We examined the IASCC initiation behavior in the BFB,
by the constant load SCC tests in the hot lab facilities. Based on
the SCC test, the SCC initiation time was very short, around
several hundred of hours, compared to plant operation time.
Therefore, we focused on the relation between the material
degradation index due to neutron irradiation, dpa , and the BFB
stress. We present the evaluation method to secure structural
integrity of the baffle structure, including the threshold stress
diagram (fluence (dpa) vs. stress) and irradiation creep rate
formula.
ACKNOWLEDGMENTS
The JNES IASCC project “Evaluation Technology for
Irradiation Assisted Stress Corrosion Cracking” is supported by
the Ministry of Economy, Trade and Industry (METI).
REFERENCES
[1] JAPEIC Report: Development of technology for long life
plant (PLEX), “ Irradiated stainless steel SCC tests
(PWR)”, (1997/3) [in Japanese]．
[2] I. Suzuki, M. Koyama, H. Kanasaki, H. Mimaki, M.
Akiyama, T. Okubo, Y. Mishima, T. R. Mager, “Stress
Corrosion Cracking of Irradiated Stainless Steels in
Simulated PWR Primary Water”, International Conference
on Nuclear Engineering, ASME, 5 (1996), p.205-213
[3] R. P. Shogan, T. R. Mager, “Susceptibility of Type 316
Stainless Steel to Irradiation Assisted Stress Corrosion
Cracking in a PWR Environment”, Proceedings of the 10th
International Conference Environmental Degradation of
Materials in Nuclear Power Systems- Water Reactors,
NACE, (2001).
[4] K. Fukuya, K. Fuji, M. Nakano, N. Nakajima, M.Kodama,
“Stress Corrosion Cracking on Cold-Worked 316 Stainless
Steels Irradiated to High Fluence”, Proceedings of the 10th
International Conference Environmental Degradation of
Materials in Nuclear Power Systems- Water Reactors,
NACE, (2001).
[5] J. Conermann, R. Shogan, K. Fujimoto, T. Yonezawa, Y.
Yamaguchi, “Irradiation Effects in a Highly Irradiated Cold
Worked Stainless Steel Removed from a Commercial
PWR”, Proceedings of 12th International Conference
Environmental Degradation of Materials in Nuclear Power
Systems- Water Reactors, TMS, (2005).
[6] K. Fujimoto, T. Yonezawa, E. Wachi, Y. Yamaguchi, M.
Nakano, R. P. Shogan, J. P. Massoud, T. R. Mager, “Effect
of the Accelerated Irradiation and Hydrogen/helium Gas on
IASCC Characteristics for Highly Irradiated Austenitic
Stainless Steels”, Proceedings of 12th International
Conference on Environmental Degradation of Materials in
Nuclear Power Systems- Water Reactors, NACE, (2005).
[7] K. Fukuya, K. Fuji, H. Nishioka, Y. Kitunai, “Evaluation of
Microstructures and Microchemistry in Cold-work 316
Stainless Steels under PWR Irradiation”, J. Nuclear Science
and Technology, Vol.43, No.2, pp.159-173, (2006).
[8] P. Freyer, T. R. Mager, M. A. Burke, “Hot Cell Crack
Initiation Testing of Various Heats of Highly Irradiated 316
Stainless Steel Components Obtained from three
Commercial PWRs”, Proceedings of 13th International
Conference on Environmental Degradation of Materials in
Nuclear Power Systems Whistler, British Columbia August
19 - 23, 2007.
[9] H. Nishioka, K. Fukuya, K. Fujii, T. Torimaru, “IASCC
Properties and Mechanical Behavior of Stainless Steels
Irradiated up to 73dpa”, Proceedings of 13th International
Conference on Environmental Degradation of Materials in
Nuclear Power Systems Whistler, British Columbia August
19 - 23, 2007.
[10] K. Takakura, K. Nakata, M. Ando, K. Fujimoto, E. Wachi,
“Lifetime Evaluation for IASCC Initiation of Cold Worked
316 Stainless Steel’s BFB in PWR Primary Water”
Proceedings of 13th International Symposium on
Environmental Degradation of Materials in Nuclear Power
Systems-Water Reactors, August 19-23, 2007, Whistler,
Canada.
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Former Plate
Core Barrel
Baffle Former Bolt
Former Plate
Baffle Plate
Core Barrel
Baffle Plate
Former Plate
Baffle Former Bolt
Core Barrel
Barrel
Former Bolt
Former Plate
Figure 1. Core Internals of PWR Plant
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Table 1.
Material
C
0.05
0.06
0.05
BFB
FTT1
FTT2
Si
0.55
0.63
0.63
Chemical composition of BFB&FTT materials
Chemical Composition (wt%)
Mn
P
S
Ni
Cr
1.55
0.021 0.025 12.45 17.71
1.68
0.020 0.005 12.25 16.25
1.70
0.019 0.005 12.32 16.59
Mo
2.26
2.36
2.31
Fe
Bal.
Bal.
Bal.
Table 2. Final cold work condition and grain size number
Material
Final Cold Work
Grain Size No.
BFB
20%
4.8
FTT
12%
9.2
1.27
φ7.65
Table 3. Constant load SCC test condition
Items
Condition
H3BO3
1,200 ppm as B
LiOH
2.0 ppm as Li
Dissolved Oxygen (DO)
< 5 ppb
Dissolved Hydrogen (DH)
30 cc/kg H2O･STP
Temparture
320, 340°C
Duration Time
Max. 2,000hr
7～10
Figure 2.
Constant load SCC test specimen (C-ring type: BFB&FTT)
Figure 3.
Constant load SCC test equipment and specimen holder
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BFB
FTT
Figure 4. Photograph of C-ring BFB and FTT specimen after SCC test
BFB
FTT
Figure 5. SEM images of C-ring BFB and FTT specimen after SCC test
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1400
1400
Test Temparature：290℃
1200
1000
1000
Stress, σ (MPa)
Stress, σ (MPa)
Test Temparature：290℃
1200
800
600
400
FTT Failures
800
600
400
20dpa FTT Failures
200
200
40dpa FTT Failures
70dpa FTT Failures
0
0.1
1
10
100
Time to Failure, t （ｈｒ）
1000
0
10000
0-
1400
20
40
60
Fluence, φ（dpa）
80
100
1400
Test Temparature：340℃
Test Temparature：320℃
1200
1200
1000
1000
BFB Failures
FTT Non-Failures
10dpa
10dpa
20dpa
20dpa
10dpa
20dpa
20dpa
30dpa
40dpa
40dpa
70dpa
70dpa
800
600
400
200
0
Stress, σ (MPa)
Stress, σ (MPa)
FTT Failures
BFB Failures
BFB Non-Failures
BFB Failures
BFB Non-Failures
FTT Failures
FTT Failures
Non-FTT Failures
Non-FTT Failures
FTT Failures
Non-FTT Failures
FTT Failures
Non-FTT Failures
0.1
800
600
400
200
0
1
10
100
Time to Failure, t （ｈｒ）
1000
10000
-0
1400
Test Temparature：340℃
1200
1200
1000
1000
Stress, σ (MPa)
Stress, σ (MPa)
40
60
Fluence, φ（dpa）
80
100
1400
Test Temparature：320℃
800
600
20dpa
20dpa
20dpa
40dpa
40dpa
70dpa
70dpa
400
200
0
0.1
BFB Failures
BFB Non-Failures
FTT Failures
FTT Failures
Non-FTT Failures
FTT Failures
Non-FTT Failures
1
10
100
Time to Failure, t （ｈｒ）
Test Temparature：290, 320, 340℃
600
400
1000
0
10000
0-
BFB Non-Failures (340℃)
BFB Non-Failures (320℃)
FTT Non-Failures (340℃)
FTT Non-Failures (320℃)
FTT Non-Failures (290℃）
INSS Non-Failures 〔9 〕
WH Non-Failures 〔8〕
20
40
60
Fluence, φ（dpa）
80
100
Figure 7. Constant load SCC test results at
290/320/340°C as stress (σ) versus fluence
BFB Failures (340℃)
BFB Failures (320℃)
FTT Failures (340℃)
FTT Failures (320℃)
FTT Failures (290℃)
INSS Failures 〔9〕
WH Failures 〔8 〕
-2
-3
Threshold stress of BFBs
1000
BFB Failures
FTT Failures
200
1400
1200
BFB Non-Failures
FTT Non-Failures
800
Figure 6. Constant load SCC test results at
290/320/340°C as stress (σ) versus failure time
800
ln(1/t)
Stress, σ (MPa)
20
600
400
○75dpa,800MPa:75dpa,
30.7kcal/mol
800MPa
▲40dpa,800MPa:
4.7kcal/mol
40dpa,
800MPa
■40dpa, 600MPa: 26.3kcal/mol
40dpa, 600MPa
Total: 30.6kcal/mol
-4
-5
-6
y = -15384x + 21.199
200
Threshold Stress of FTTs
-7
1.6E-03
0
-0
Figure 8.
20
40
60
Fluence, φ（dpa）
80
100
Figure 9.
Total constant load SCC test results as stress
versus fluence
8
1.7E-03
1/T
1.8E-03
Test temperature (activation energy)
effect of constant load SCC tests
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600
600
BFB(1)
FTT(1)
B3-13（BFB 11.4dpa）
B4-11（FTT 10.5dpa）
B3-14（BFB 10.4dpa）
20dpa
500
B3-31（BFB 19.8dpa）
B4-28（FTT 19dpa）
Vikers hardness (HV0.1)
Vikers hardness (HV0.1)
B4-27（FTT 19.5dpa）
500
B3-32（BFB 19.8dpa）
400
300
B4-12（FTT 9.5dpa）
20dpa
10dpa
200
FTT 0dpa
400
300
10dpa
200
0dpa
100
0
50
100
150
Depth profile from outer surface of C-ring type specimens (µm)
Figure 10.
100
200
0
50
100
150
Depth profile from outer surface of C-ring type specimens (µm)
200
Hardness profile from outer surface of C-ring type specimen for BFB&FTT
SCC initiation threshold stress（MPa）
1000
900
800
700
600
y = -3.3088x + 2036.6
R2 = 0.7204
500
400
300
200
BFB
100
FTT
0
300
350
400
450
500
Vickers hardness of the outer surface layer （HV0.1）
Figure 11.
Relationship between SCC initiation threshold stress
and surface hardness of C-ring type specimen
Creep Rate, mm/mm/dpa
0.001
Irradiation creep rate formula
340℃
ε＝A・σ・[ 1-exp(-A’・F)] + B0・σ(1+B1・σ3)・F
290℃
ε ： Creep strain
B0 ： 0.98x10-6 /(dpa・MPa)
B1 : 0.33x10-8 MPa -3
σ ： Stress (MPa)
A ： 0.73x10-6 /MPa
F : fluence (dpa)
A’ : 4.3 /dpa
Temperature : 290C-340C
Formula
Bσ（1+B''σ^3）
0.0005
0
0
200
400
600
800
Stress, MPa
Comparison of Halden experimental data and this formula
Figure 12. Formula of irradiation creep rate
9
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：Achievements in the Project
Stress Evaluation of Bolt
(Detailed Analysis)
－Tightening Stress
－Thermal Deformation
－Irradiation creep
－Swelling
－Irradiation Hardening
Evaluation of Threshold Stress due to IASCC
(The SCC Test applying constant load with various
temperatures, irradiation levels and stresses)
Evaluation of Plant Operating Time and Irradiation Level
(considering the difference in operating characteristics)
Prediction of Timing of When BFB Is Damaged
Defining the allowable number of
defect bolts for the retention of the
baffle structure functions.
Inspection：Determine whether the number of defect bolts can be within the
allowable number of defect bolts(i.e. the continuing use is possible or not) by the
next time of inspection
Repair Work Is Necessary.
Repair Work：Replace part or all of the defect bolts.
Figure 13. Essential Features of Procedure to Specify Appropriate Check Timing
for Retention of Baffle Structure Functions
バッフルフォーマボルト割れ無し
1400
試験温度：290、320、340℃
Test
Temperatures：290 deg.C., 320 deg.C.,
and 340 deg.C.
Stress σ（MPa）
1200
No Crack in Baffle Former Bolt test specimen
バッフルフォーマボルト割れ有り
Crack Generation in Baffle Former Bolt test specimen
No Crackシンブルチューブ割れ無し
in Thimble Tube test specimen
Crack Generation
in Thimble Tube test specimen
シンブルチューブ割れ有り
1000
Region 3
800
600
Region 2
400
200
0
Region 1
Curve A：Threshold line of SCC Initiation in
Thimble Tube test specimen
-0
20
40
60
Neutron Irradiation Level（dpa）
80
100
Figure 14. Threshold Stress Diagram of IASCC Initiation
10
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