E-Journal of Advanced Maintenance Vol.1 (2009) 1-29
Japan Society of Maintenology

Stress Corrosion Cracking in Low Carbon
Stainless Steel Components in BWRs
Shunichi SUZUKI 1,*, Kenrou TAKAMORI1, Katsuhiko KUMAGAI1,
Akihiro SAKASHITA2, Norimichi YAMASHITA2, Chikashi SHITARA3 andYuichi OKAMURA4
1

Materials Engineering Center, Tokyo Electric Power Company (TEPCO),
4-1 Egasaki-cho, Tsurumi-ku, Yokohama 230-8510, Japan
2
Nuclear Asset Management Dept., TEPCO
1-1-3 Uchisaiwai-cho, Chiyoda-ku, Tokyo 101-0001, Japan
3
Nuclear Quality & Safety Management. Dept.,TEPCO
1-1-3 Uchisaiwai-cho, Chiyoda-ku, Tokyo 101-0001, Japan
4
Fukushima-dai2 Nuclear Power Station, TEPCO
12 Kobamasaku, Namikura, Naraha-machi, Futaba-gun, Fukushima 979-0695, Japan(Received; January 9th, 2009)

Abstract. In recent years, numbers of SCC have occurred in core shrouds and primary loop recirculation piping
made of low carbon stainless steels that had been recognized to be an SCC-resistant material. These incidents
resulted in long-term shutdown of Japanese boiling water reactors and have drawn social as well as technicalinterests.
This paper will provide an introductory review on (1) background of SCC observed in low carbon stainless steel
components in BWRs, (2) characteristics of SCC in core shrouds and PLR piping, (3) structural integrity
evaluation, (4) SCC mitigation techniques, (5) SCC mechanism and (6) research topics that should be covered
regarding these issues.

KEYWORDS: stress corrosion cracking, low carbon stainless steel, L-grade, BWR, core shroud, recirculation
piping, PLR, BWR

1.

Introduction

In order to ensure safe and stable operation of light water nuclear reactors that supply major partof electric power in Japan, voluntary as well as statutory inspections are being performed periodically.
As aging of nuclear reactors progresses, it is becoming commonly recognized that implementing latest
technical findings and data to maintenance activities are essential. Based on such consensus, Japanese
nuclear regulator (Nuclear and Industry Safety Agency, The Ministry of Economy, Trade andIndustry) has issued an evaluation guideline for aging of nuclear power plants. It has deemed that
continued safe and stable operation of aging plants can be attained as long as the proper maintenanceschedule is followed and revised based on reasonable inspection results using recently-developed
techniques and methodologies.
Quantitative understanding of structural material aging as well as lifetime evaluations areimportant parts of management of aging plants and life extension assessments. Typical aging
degradations and failure modes include fatigue, irradiation embrittlement, irradiation-assisted stresscorrosion cracking, aging, corrosion, abrasion, etc. Although there still remain arguments on whether
stress corrosion cracking (SCC) that usually occurs after certain incubation period should be attributed
to aging, it is clearly recognized as the most important phenomenon to be managed to attain safe andstable operation of power plants over a prolonged period.
*

Shunichi SUZUKI, E-mail: suzuki.shunichi@tepco.co.jp
ISSN – 1883 – 9894/09 © 2009 – JSM and the authors. All rights reserved.

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S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

In recent years, numbers of SCC incidents have occurred in core shrouds and primary loop
recirculation piping (PLR piping) made of low carbon stainless steels that had been recognized to be
an SCC-resistant material. These incidents resulted in long-term shutdown of Japanese boiling water
reactors (BWRs) and have drawn social as well as technical interests [1].
This paper will provide an introductory review on (1) background of SCC observed in SUS316L
components in BWRs, (2) characteristics of SCC in core shrouds and PLR piping, (3) structural
integrity evaluation, (4) SCC mitigation techniques, (5) SCC mechanism and (6) research topics thatshould be covered regarding these issues.
2.

Modern History of SCC in Low Carbon Stainless Steels

2.1.

Background of Low Carbon Stainless Steel Implementation in BWRs

From the late 1970’s, SCC was found in heat-affected zone of SUS304 stainless steel pipe jointsof Japanese as well as foreign BWRs [2]. SCC in SUS304 occurs when three primary factors, hightensile stress, corrosive environments and sensitization of the material present simultaneously. From
extensive research activities, it was concluded that the root cause of this phenomenon can be
attributed to the sensitization; decrease in corrosion resistance at the grain boundary due to the
Cr-depletion caused by the chromium carbide precipitation during welding process. For this reason,
the applicability of low carbon stainless steels such as SUS304L and SUS316L were evaluated toprevent sensitization. In 1977, SUS316(LC) had been developed by collaborative research project byJapanese BWR owners group. The regulator side (Nuclear Power Engineering Test Center and theJapan Power Plant Institute) had conducted a verification program and concluded that low carbon
stainless steels are significantly superior material in SCC resistance and increase the SCC incubation
period by 7 to 50 times compared to SUS304 [3].
Based on these evaluations, SUS316(LC) and SUS316L were adopted in PLR piping and in coreshrouds in Unit 2 of TEPCO Fukushima dai-2 nuclear power plant (operation started in February1984) and later units. Simultaneously, a deaeration operation (reducing oxygen level of reactor water)
during reactor start up was applied as an environmental mitigation technique and it was thenconcluded that sufficient treatment methodology was established to prevent SCC from occurring. In
the 1990’s, however, other incidents were observed in core shrouds of foreign BWR made of
SUS304L and of Fukushima dai-2 Unit 3 made of SUS316L
2.2.

Early Findings on SCC in Low Carbon Stainless Steels

Considering the SCC in low carbon stainless steel equipment of foreign BWRs, proactive
research programs on SCC in low carbon stainless steel were conducted domestically from themid-1990s. Several characteristics of the SCC in low carbon stainless steels as follows have been
revealed through these research programs by the early 2000’s.
2.2.1. Findings on SCC initiation
1. Constant load testing
The results of creviced uni-axial constant load SCC test (with a load of 3.5Sm (Sm is the
maximum allowable stress for the material in the plant design), corrosion potential of 250~280 mVSHE,
620 ºC×24 hours of heat treatment for sensitization), where SCC did not initiate in SUS316L up to7000 hours, showed that the material was high-resistant to SCC [4]. It should be noted, however, thatthe surfaces of these test specimens were polished before testing.
2.

CBB (Creviced Bent Beam) testing of solution annealed materials [5]When the surface layer is hardened by heavy grinding and the Vickers hardness exceeds 300HV,SUS316L becomes more susceptible to TGSCC (Trans-Granular Stress Corrosion Cracking) in hightemperature water (Fig. 1).
3. Initiation as TGSCC from the machined/ground surface and propagation as ISGSCC (Inter
Granular Stress Corrosion Cracking) [6].

2

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

CBB tests on SUS316L weld joints with surface machined layer revealed TGSCC initiations andsubsequent propagation as IGSCC (Fig. 2). While TGSCC initiation and IGSCC propagation wasobserved in surface-milled specimens, SCC wasn’t found in specimens with no heavy surface
machining.

Fig. 1. Relationship between Vickers hardness and the average stress corrosion crack depth in CBB tests

Welded metal

SCC
Heavily ground
layer

Welded metal

Heavily groundlayer

Test Period
Welded metal
Test Period:
2000HrTest Period:
2000HrTest Period

Fig. 2. CBB test results from samples taken from SUS316L welded joints in the Na2SO4 injectedenvironment (Dissolved oxygen: 8 ppm, Conductivity: 1.4 μS/cm) at 288ºC,

2.2.2.

Findings on SCC growth
The following are findings on the growth rate of SCC of low carbon stainless steels in normal
reactor water chemistry and in less aggressive environments such as hydrogen water chemistry.
1. Effects of Corrosion Potential [7]Figure 3 shows the dependency of SCC growth rate in stainless steels on corrosion potential. All
steel types indicated that crack growth slowed with decreasing corrosion potential. Crack growth rateof SUS316L and SUS316NG is lower than that of SUS304 by one order of magnitude. The thresholdcorrosion potential for SCC growth is -100 mVSHE for low carbon stainless steels, which is higher thanthat of SUS304 by about 100 mV. The fracture surface showed inter granular crack morphology thatlooks like “rock candy”.

3

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

2.

Effect of Stress Intensity Factor
SCC growth rate is dependent on stress intensity factor (K, hereafter), and drops as the applied K
decreases. This dependency of SCC growth rates of SUS316L and SUS304 on K in normal waterchemistry and hydrogen water chemistry are reflected in JSME code, S NA1-2002 [8] (Fig. 4).
2.2.3. Findings on the Effects of Heat Treatment [9]Even when solution heat treated low carbon stainless steels undergo sensitization heat treatment(e.g. 700 ºC×10 hours), precipitation on grain boundaries or Cr-depletion zone are not formed.
However, when the same heat treatment was applied after 10 % cold work, Laves layer precipitation(Fe-Cr-Mo) and W-shaped Cr-depletion layers are observed at grain boundaries (Fig. 5).
10-5

273～288℃
K=30MPa√m
316L
316NG
304
-7
10
304

Crack growth rate (mm/s)

Crack growth rate (mm/s)

10-6Inlet Conductivity

10-8* : 0.1 μS/cm** : 0.07 μS/cm*** : <0.1 μS/cm
Else: <0.2 μS/cm: SUS304: SUS316L: SUS316NG: SUS304L

10-6

JSME Code:Sensitized stainless steels

10-7

10-8

*
***

JSME Code:Low carbon stainless steels

**

10-9
-800 -600 -400 -2000ECP (mVSHE)

200

10-9

400

Fig. 3. Effect of corrosion potential on SCC
growth rate of austenitic stainless steels

1

10

Stress intensity factor K (MPa√m)

Fig. 4. SCC growth rate disposition curves foraustenitic stainless steels in JSME code

Grain
Boundary

Grain
Boundary

Fe-Cr-Mo
Precipitate

Measurement
path

Cr Precipitate

Grain
Boundary

Measurement
path

Cr wt. %Cr wt. %Cr wt. %Measurement
path

100

Distance from thegrain boundary (nm)Distance from thegrain boundary (nm)Distance from thegrain boundary (nm)(a) 700 ºC×10 hours(b) 10 % Cold Work +700 ºC×10 hours(c) 700 ºC×100 hours

Fig. 5. Effect of heat treatment and cold work on formation of Cr-precipitate and Cr-depletion zone at
grain boundaries of low carbon stainless steel

3.

SCC in Core Shrouds and Flaw Assessment

Before the implementation of Japanese fitness-for-service standards in July 2001, the first SCC
in low-carbon stainless steel component in Japanese BWR was observed in the lower ring weld (H6a)of the core shroud of Fukushima dai-2 Unit 3.

4

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

A part of the crack was taken out by EDM(Electrical Discharge Machining) from theshroud material. A detailed examination on thecrack was conducted, which revealed theexistence of the hardened layer within a range of
300 μm from the surface.
The crackmorphology in this hardened layer was TGSCC,
and that of the subsequent propagation was
IGSCC (Fig. 6).
Thorough inspections on low-carbon
stainless steel components of all the Japanese
BWRs were conducted starting from August
2002 and, as shown in Fig. 7, a number of SCC
had been identified [10] by ultrasonic testing(UT) and visual testing (VT). VT on coreshrouds were carried out using cameras thathave enough resolution to identify a wire with athickness of 1 mil (1/1000 inch = about 0.025mm) under water

Fig. 6. SCC observed on the edge surface ofshroud ring (SUS316L), adjacent to H6 weld of
the core shroud of Fukushima dai-2 Unit 3

Power
Station(*) Material- UnitReactor Pressure Vessel Core Shroud

1F-4

Hoop weld
seams
H1
H2
ShroudShell

V14, V16

Shroud
Ring

304L

Weld
Seam
H4
H3

2F-2

316L
H4

H3
H4

Flow of coolant

TGSCC

IGSCC

2F-3

316L

H6a

H3
H4
H3

H6b
H7a
H7b

2F-4

Core
shroudsupport(*)Abbreviations for Power Stations
1F: Fukushima dai-1
2F: Fukushima dai-2
KK: Kashiwazaki-Kariwa

316L

KK-1

316L

KK-2

316L

KK-3

316L

KK-5

316L

H4

Face

Location

Inner

Outer

Shell
Ring
Shell
Weld

Inner

Shell

Inner

Ring
Inner

Inner

Shell
Shell/Weld
Shell

H3

Outer
Inner

Ring

H4

Outer

Shell

H1
H6a
V16
H7
H6a
H7a
H4
V14

Outer
Inner
Outer
Inner
Inner

Ring
Shell
Ring
Ring
Shell

Max. Surface
Size
Width x
Height(mm)1
180 x 80
230
530 x 25
3
40 x 1717200 x 30
780 x 90
1
8
10
210
1
13
1
16 x 30530
1
31 x 191
11 x 159145
132 x 223
10 x 20
1
90
Many
–
1
20
Many
–
–
Many
–
1
18 x 151
33 x 16

Numberof
Crack

Fig. 7. Schematic showing weld seams in a core shroud and list of observed SCC as of 2002

As shown in Fig. 8, multiple SCC occurred discretely all over shroud ring edges (referred to as“circumferential crack”, hereafter). Due to heavy machining on the surfaces by milling during its
manufacturing, a hardened layer of over 300HV was formed as shown in Fig. 9 [11].
In the shroud’s shell, on the other hand, a smaller number of SCC occurred locally along
circumferential welds (referred to as “partial crack”, hereafter). The surface crack orientation of these
SCC did not necessarily follow the weld line. They seemed to be affected heavily by the multi-axial
welding residual stress profile. Some cracks showed a radial propagation on the surface as shown in
Fig. 10. A hardened layer of about 50 μm was observed locally at cracked area and was assumed to beformed due to grinder finishing during manufacturing processes. It is assumed that the material
hardening in the surface layer enhanced the SCC susceptibility [11].

5

9mm

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

32mm

WeldedMetal

H7a weld
CracksCrack
H7b weld
49mm

Schematic of the boat-shaped sample(boat sample) taken by EDM from theactual shroud

Outer
Face

Inner
Face

Fig. 8. Schematics showing multiple cracks occurred discretely over the entire circumference of the edgeof shroud rings.

Distance from the fusion boundary: 1.2mm: 4.5mm: 12mm

400
300

100 mmVickers Hardness (HV)

500

200
100

H4 Weld

0
0.0

0.1

0.2

0.3

0.4

180 mm0.5

Distance from the inner surface (mm)Fig. 10. Example of partial crack observed in the
shroud shell, near H4 weld of Fukushima dai-1
Unit 1.

Fig. 9. Vickers hardness profiles in thickness
direction measured on the boat-sample takenfrom the shroud ring (H7a) of
Kashiwazaki-Kariwa Unit 3.

3.1.

Fitness-for-Service Evaluations Based on the Code and Standards [10]3.1.1. Modeling of partial cracks
As shown in Fig. 10, SCC in the shroud shell often showed a two dimensional expansion on thesurface, including in axial direction (a direction perpendicular to the weld lines). However, sincebending loads become dominant in case of earthquakes, cracks projected to a cross-section
perpendicular to the shroud axis have been considered. The projected crack was then modeled as athrough-wall crack with the same surface length, and only the circumferential propagation was
evaluated, as a more conservative and simplified analysis (Fig. 11).
1. Modeling as throughwall cracks

3. Integrating all the cracks
in the same plane

Observedcracks

2. Predicted crack
growth in next five
years

Assumedcrack
lengthFig. 11. Modeling of partial cracks in fracture assessments

6

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

3.1.2. Modeling of circumferential cracks
Discrete partial cracks that expand
over the entire circumference of shroud weremodeled as a circumferential crack with auniform depth as shown in Fig. 12. Averagecrack depth as measured by UT was used inthis modeling.

3.2.

ObservedcracksAssumed crack

Fig. 12. Modeling of discrete partial cracks expanding
over the entire circumference as a circumferential
crack in fracture assessments

SCC Growth Evaluations

3.2.1. Background on SCC growth evaluations
SCC growth analyses are performed referring to the SCC growth rate disposition curve in JSME
code. These diagrams are established so that they envelops the maximum value of experimental dataobtained from laboratory SCC experiments that simulate actual environmental conditions (Fig. 4).
3.2.2. Crack growth rate for partial cracks
Since the size of partial cracks when they were found was sufficiently smaller than the
maximum allowable limit, as will be discussed later, crack growth evaluation of partial cracks was
conducted focusing mainly on simplifying the procedure by taking a highly conservative assumptionthat the crack growth rate is equal to the maximum rate in the JSME disposition curve for low-carbon
stainless steel (= 2.1×10-7 mm/s) independent of applied K.
In cases in which neutron fluence exceeds 3×1024 n/m2 like weld lines H3 or H4 that are locatednear the core fuel, the highest rate in JSME disposition curve for sensitized stainless steel (= 9.2×10-7mm/s) was used regardless of the applied K, to take the irradiation effect on SCC growth rate into
account.
When multiple cracks were found in the same weld seam, it was assumed that each crack willgrow in both circumferential directions at the highest rate. Fracture assessment was then performed byassuming a single crack with a length that is equal to the sum of the predicted length of each crack
after 5 years (Fig. 11).
3.2.3. Crack growth rate for circumferential crack
Detailed analyses were performed for circumferential crack growth in the wall-thickness
direction considering K vs. crack growth rate dependence, since there is a smaller ligament comparedto the cases of partial crack that grows in circumferential direction. Welding residual stress is adominant factor in SCC growth and is calculated by thermal elasto-plastic finite element (FE)analyses. Prior to an FE analysis, a full-scale mock-up was fabricated using material and welding
procedures equivalent to those were used on the actual component. The change in temperature profile
with time was measured with thermocouples. Heat input conditions in FE analyses were determinedso that the calculated temperature profile will correspond to the actual measurement.
The K-value was calculated using a formula in API-579 [12] since the K formula for acircumferential crack on the outer surface of a cylinder was more appropriate to evaluate cracks in the
shroud ring than the formula specified in JSME code as an example assuming a thumbnail crack in aplate.
As examples, welding residual stress distributions for H6a and H7a weld of a shroud of 1100
MW-plant are shown in Fig. 13. The change in K with crack depth and the change in crack depthwith time are shown in Figures 14 and 15, respectively. In each figure, (a) shows the result at 12 mmfrom the H6a fusion line, where cracks had been found in the shroud ring of Kashiwazaki-Kariwa
Unit 1; (b) shows the result at 10 mm from H7a, where cracks were found in Kashiwazaki-Kariwa
Unit 3.
Since H6a and H7a welds are located at the bottom part of the reactor core, the effect of neutron
irradiation is supposed to be insignificant, and the JSME disposition curve for low-carbon stainlesssteels was referred to in SCC growth evaluations.
In case of the H6a example, there is a large area in the center of the plate where residual stress7

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

400

Welding residual stress (MPa)

Welding residual stress (MPa)becomes compressive and, therefore, K becomes 0 at some point during the crack growth, as shown in
Fig. 14 (a). For this reason, the evaluation shows that crack will be arrested at about 30 mm from thesurface, as shown in Fig. 15.
300
200
100
0
-100
-200
-300
-400

0

10

20

30

40

50

400
300
200
100
0
-100
-200
-300
-400

0

10

Distance from the outer surface (mm)

20

30

40

50

60

70

Distance from the inner surface (mm)(a) H6a (12mm from the fusion boundary)(b) H7a (10mm from the fusion boundary)Stress intensity factor (MPa√m)

Stress intensity factor (MPa√m)

Fig. 13. Welding residual stress (axial component σz) profiles in plate thickness direction of core shroud
rings of a 1100 MW class BWR (by thermal elasto-plastic FEM analyses).

60
50
40
30
20
10
0

0

10

20

30

40

50

60

70

80
70
60
50
40
30
20
10
0

0

10

Distance from the outer surface (mm)

20

30

40

50

60

70

Distance from the inner surface (mm)(a) H6a (12mm from the fusion boundary)(b) H7a (10mm from the fusion boundary)Fig. 14. Changes in the stress intensity factor with crack growth for circumferential cracks in H6a andH7a welds of a 1100 MW class BWR (K-formulas specified in API579 were used)

70

60

Crack depth （mm）

Crack depth （mm）

70

50
40
30
20

50
40
30
20
10

10
0

60

0

5

10
15
20
Service period (year)2530(a) H6a (12mm from the fusion boundary)

0

0

5

10

15
20
25
Service period (year)3035(b) H7a (10mm from the fusion boundary)Fig. 15. Results of SCC growth evaluations for circumferential cracks in H6a and H7a welds(SCC growth rate disposition curve for low carbon stainless steels in JSME code was used)

3.3.

Fracture Assessments [13]3.3.1. Background on shroud fracture assessments
Primary functions of core shrouds are to support the core fuel and to channel the flow of coolingwater within the reactor pressure vessel. Since a shroud is not a part of the pressure boundary and issubjected to relatively low stress during normal operation, it can be regarded that the function and

840

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

structural integrity of shrouds are maintained even with some cracks present in them as long as
unstable fracture does not occur under the maximum load (= operational load + seismic load) assumed
in the shroud design.

Partial crack length (m)

12
10
8
6
4

Crack size as of 2002
Predicted size after 5
years of service period
Max. allowable size
1F-4: Unit 4 of Fukushima dai-1
2F-X: Unit X of Fukushima dai-2
KK-X: Unit X of Kashiwazaki-Kariwa

20

H4 H3 H4
1F-4
2F-3

H3 H4
2F-4

H4
KK-1

(a) Partial crack

Depth of Circumferential crack (mm)3.3.2. Limit load evaluation methods
The plastic collapse is assumed to be the primary fracture mode for components made of highly
ductile austenitic stainless steels such as SUS316L and SUS304.
Evaluations are, therefore, performed based on JSME Maintenance Code Attachment E-8;
“Limit Load Evaluation Methods” [8]. The most severe loading condition “Operation Conditions I, II+ S2” are used, referring to the earthquake response evaluations in the application documents forapproval of plant construction. The differential pressure between the inside and outside of the shroud
is also considered in accordance to these construction permit application values. Flow stress σf is set at
2.7 Sm and the safety factor is set at 1.39 based on the maintenance code.
Considering the maximum load at the time of an earthquake, the maximum allowable crack
length Llim was determined so that the stress on the net cross sections of cylinders corresponds to theflow stress.
Figure 16 shows the results of fracture assessments by the limit load evaluation conducted as of2002 when the cracks were discovered. Figure 16(a) displays the results for partial cracking; thedashed lines indicate the maximum allowable crack length evaluated based on the maintenance codes.
As shown in this figure, the structural integrity of shrouds is still maintained even if a half of theentire circumference (about 14~16 m) is separated by SCC. It is also clear from the figure that there
still remains enough safety margin even after adding further crack growth assumed in the next 5 yearperiod.
Results of circumferential crack evaluations shown in Fig. 16(b) indicate that predicted crack
length 5 years later does not exceed allowable limits. Considering that the result is drawn based on aconservative modeling of discrete cracks by a circumferential crack, it can be concluded that the
integrity of the shroud can be maintained at least during a 5 year-period after the evaluation.
60
50
40
30
20
10
0

H6a H7
KK-2

H6a H7aKK-3(b) Circumferential crack

Fig. 16. Results of fracture assessments of core shrouds by limit load analyses.

Since the crack growth rate at H6 weld decreases rapidly as the crack deepens, as shown in Fig.
15, the plant operation leaving the crack present was continued under such conditions that theperiodical inspections and structural integrity evaluations based on the code and standard would be
conducted.
The latest results of crack sizing as of July 2008 are summarized in Fig. 17 [14-15]. As can beseen in these figures, average crack depths as measured by UT are less than predicted values. Fromthese results, it was confirmed that the current crack growth evaluation procedures provide
conservative predictions of crack extension.

9

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs(*) Assuming 80 % of operation ratio

Wall thickness = 50.8mm

50Max. allowable depth

40

Predicted depth in 5 years(* ) from thecrack identification = 22.2mm

3020Predicted crack depth = 19.1 mm

10
0

5

Average depth at the latest
inspection in 2007 = 13.0 mm
10
15
20

(*) Assuming 80 % of operation ratio

60

Crack depth （mm）

Crack depth （mm）

60

25

30

Time (Effective Full Power Year)

Wall thickness = 50.8mm

50Max. allowable depth

40

Predicted depth in 5 years(* ) from thecrack identification = 22.0mm

3020Predicted crack depth = 18.7 mm

10Average depth at the latest
inspection in 2007 = 8.3 mm
10
5
15
20

0

25

Time (Effective Full Power Year)30(a) Kashiwazaki-Kariwa Unit 2(b) Kashiwazaki-Kariwa Unit 3
Fig. 17. Change in the average depth of cracks, that are left in H6a weld of core shrouds, during therecent plant operation (Kashiwazaki-Kariwa Units 2 and 3), in comparison with SCC growth predictions
using current evaluation code.

3.3.3. Fracture mechanics analysis
The fracture mechanics evaluations were applied to weld joints near the core fuel (H3 and H4)
with neutron fluence over 3×1024 n/m2. The K value was calculated for the largest crack found in aweld joint after adding predicted crack extension in the next 5 years to the detected crack size.
Fracture assessment was based on the comparison between the evaluated K and the fracture toughness
value KIC.
The crack modeling as a through-wall crack as shown in Fig. 11 was also used here. The stressintensity factor was calculated using the K-formula for a through-wall crack along the circumferenceof a cylinder [16].

K = G m σ πaHere,
σ = Pm + Pb: applied stress
Gm = 0.2227a(R·t)-0.5 + 1: modification coefficients for curvature
a: half of crack length on the surface
6

Partial crack length (m)The safety factor (SF) of 1.39 wasconsidered and the maximum allowable crack
size was determined as a value that satisfies K =KIC/SF. The fracture toughness KIC was assumed
to be 165 MPa for locations with the neutron
fluence of 3×1024~1×1025 n/m2 and to be 43.2
MPa for locations with 1×1025 n/m2.
Figure 18 shows the results of fracture
mechanics evaluations. Fluence of over 1×1025
n/m2 was assumed only in Fukushima dai-1 Unit
4. The allowable crack lengths for
Kashiwazaki-Kariwa Unit 1 is smaller because
the assumed seismic load in its design was
higher than other units. Despite the smaller limit
for crack length, the shroud’s integrity of all theunits was still secured for at least the next 5
years

5

Max. crack size as of 2002
Predicted size after 5 yearsof service
Max. allowable size

4
3
2
1
0

1F-4

2F-3

2F-4 KK-1 KK-2

1F-4: Unit 4 of Fukushima dai-1
2F-X: Unit X of Fukushima dai-2
KK-X: Unit X of Kashiwazaki-Kariwa

Fig. 18. Results of fracture assessments of H4
welds of core shrouds by fracture mechanics
evaluations taking irradiation effects into account.

10

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

4.

SCC in Recirculation Piping and the Flaw Assessment

Figure 19 shows SCC observed in PLR piping of TEPCO BWRs [17]. A larger number of SCC
are found in the main pipe (600A: 600 mm in diameter) compare to smaller pipes with diameters of
300 or 400 mm (300A or 400A, respectively).
So far, SCC were observed only in weld seams with the groove root angle of 30 degrees or larger.
Narrow-groove is commonly used in recently-fabricated seams in order to reduce the welding residual
stresses and no SCC was found in such seams. The material used for PLR piping is basically low
carbon stainless steel SUS316LC (SUS316NG) that is nitrogen-strengthened 316L.
Number of Cracked／Inspected Seams

Core shroud

Riser(Φ 300mm)

Header(Φ 400mm)

Kashiwasaki Fukushima-Kariwa-dai2

Reactorpressure
vesselDiameter of Pipe
(mm)
Unit 1
Unit 2
Unit 3
Unit 4
Unit 1
Unit 2
Unit 3
Unit 4
Unit 5
Total

600

400

300

Other

1 /6
4 /43
8 /28
8 /32
6 /33
4 /29
2 /26
5 /27
2 /29
40 /253

0/0
0 /16
0 /12
1 /12
6 /12
0 /12
0 /12
0/8
0 /12
7 /96

0 /10
0 /22
0 /20
0 /20
1 /20
0 /20
0 /20
0 /20
1 /20
2 /172

0/7
2 /33
2 /15
2 /33
1 /17
0 /16
0 /16
1 /19
5 /27
13 /183：Re-circulation loop
：Flow of Coolant
Main Pipe(Φ 600mm)

Re-circulation
pump

Fig. 19. Schematic showing recirculation (PLR) piping and list of observed SCC as of 2002

4.1.

Accuracy in Crack Sizing by UT

Before 2002, the accuracy in crack sizing for SCC in SUS304 by conventional UT (using lateralwave and tip echo) had been evaluated by a Japanese national project UTS; “Confirmation Testing for
Ultrasonic Inspection Flaw Detectability and Sizing Accuracy” [18]. Based on the evaluations in the
program, consideration of assumed error in SCC sizing of 4.4 mm (= as double standard deviation)
was required in UT inspections. As shown in Fig. 20(a), however, SCC sizing results [19] showedwider scatter than ±4.4 mm for actual cracks in PLR pipes. One of the main reasons for this larger
error is that some cracks have crossed the fusion boundary and have propagated into the weld metal,
generating additional echoes reflected from fusion boundaries and kinks in the crack wake that might
be judged as tip echoes. Since SCC in SUS304 tends to grow along the Cr-depletion zone in the HAZ
rather than propagating into the weld metal, it can be said that this is a specific problem for SCC in
low-carbon stainless steels.
However, it was then demonstrated that the comparable level of accuracy can be attained byenhanced UT techniques such as Tip-Echo method using longitudinal wave, Phased Array and soforth, operated by qualified inspectors (Fig. 20(b)) [19].
Based on the consensus that the skill of the inspector has significant effect on the accuracy of UT
inspections, the Performance Demonstration System, whose necessity has already been proven in theUS, was established recently to provide an objective scale to measure the skill of inspectors.

11

20

15

Measurement by advanced UT (mm)Measurement by conventional UT (mm)S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs
Plant A (Progress PT)
Plant B (Progress PT)
Plant C (Sectioning)
TEPCO plants (Sectioning)

10
Average - 2σ
(UTS program)

5.0

0.0
0.0

5.0

10

15

20

20
Average - 2σ
(UTS program)

15
L + S mode TE (A)L mode TE(B)S mode TE(B)
L + S mode TE (C)L + S mode TE (E)
PA_1+S mode TE(A)
PA_1(A)
PA_2(A)
PA_1(B)
PA_2(B)
PA(D)

10

5.0

0.0
0.0

5.0

10

15

Actual crack depth (mm)Actual crack depth (mm)(a) by Conventional UT(b) by Improved UT

20

Fig. 20. Accuracy in the sizing of crack depth (a) by conventional UT technique using lateral wave and (b)by improved UT techniques using phased array, longitudinal wave etc.

4.2.

SCC Growth Evaluations

4.2.1. Effect of material hardening on SCC growth rate
Recent research has revealed that theincrease in material yield stress due to work
300HV260-300HV
hardening etc. can accelerate the SCC growth240-260HV
220-240HV[20-22]. SCC in a PLR pipe joint propagates in
200-220HV
180-200HV-180HVthe localized hardened area formed adjacent to
the weld root due to thermal deformations
Weld
during welding, as can be seen in Fig. 21. Since
metal
the effect of material hardening on SCC growthrate is not considered in the current JSME
disposition curve, there was a concern that
evaluations based on the disposition curve canunderestimate the possible crack growth.
In order to evaluate the effect of work
hardening in PLR pipe joints, SCC growth rate
5mm
Crack
measurements were conducted using CT
Fig. 21. Vickers hardness profile adjacent to thespecimens taken from the weld metal andbutt weld-joint of 600A pipe showing hardening
hardened area in the base material of a full-size
adjacent to the root of the weld and in a thin layer
mock-up of a PLR joint, as shown in Fig. 22.
on the inner surface.
Results are shown in Fig. 23 with the JSME
disposition lines. Although the SCC growth
rates of the weld metal shown with solid
Outer surface
symbols were enveloped by the disposition
curve for low carbon stainless steels, the growth
rates of hardened areas shown with open
Inner surface
symbols exceeded the curve and it was(a)
Weldedmetal(b) Hardened area in HAZ
confirmed that consideration of the hardening
Fig. 22. Schematic showing relative locations andeffect is necessary for PLR pipe joints.
orientations of CT specimens taken from the
mock-up of a 600A pipe joint.

12S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

Crack growth rate (mm/s)

10-5316NG Hardened area (CT)316NG Welded metal (CT)316NG Welded metal (WOL)

10-6
JSME Code:Sensitized stainless steels

10-7
JSME Code:Low carbon stainless steels

10-8

10-9

1

10
Stress intensity factor K (MPa√m)

100

Fig. 23. Crack growth rates obtained from CT samples taken from the weld metal and hardened basemetal of the mock-up of PLR joints (600A).

4.2.2. Characteristics of SCC in actual PLR piping [23]Detailed examinations were performed on cross sections of 77 cracks observed in PLR pipe
joints of Japanese BWRs. Figure 24 shows the relationship between the crack depth measured on the
cross section and the period of the plant operation. The correlation between the crack depth and theoperation period was unclear and very few cracks have propagated more than 10 mm in depth.
Figure 25 shows typical cross sections for SCC in PLR piping. As can be seen in this figure, all
the cracks propagated toward the weld metal. The majority of the cracks continue to grow slightly into
the weld metal and there are few examples where the crack grew along the fusion boundary as cancommonly be seen in case of pipes made of SUS304.
The difference Δd between the crack depth d and critical depth dc (at which each crack reaches
the fusion boundary) was measured on cross sections as shown in Fig. 26 (Δd = d - dc). The histogramfor Δd (Fig. 27) clearly shows a peak at Δd = 0, which indicates that many cracks are located near thefusion boundary. Although thorough inspections were conducted within a short time period of 2002 to
2003, and the time when each crack initiated as well as how long they have been propagating varies
by cracks, many cracks were identified as their tips just happened to be at the vicinity of the fusion
boundary. Therefore, it is reasonable to consider that many of the cracks were arrested or, at least, thecrack growth rate remarkably decreased adjacent to the fusion boundaries.

Crack depth d (mm)

35
600A
550A
450A
300A

30
25

500A
400A
250A
150A

20
15
10
5
0

0

2

4

6

8

10

12

14

16

18

20

Service period (Effective Full Power Year)

Fig. 24. Relationship between depth of cracks (observed in PLR joints of Japanese BWRs) and operation
period of each plant.

13

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

1mm(a) 400A seam(b) 600A seam(b) 600A seamFig. 25. Typical crack morphologies of SCC in PLR weld joints where clear tendency for cracks to growtoward the weld metal was observed.

φ

d

dc

φ

L
(1) Within the base metal

φ

d

d =dc

(2) Arrest at the boundarydcφ(3) Growth into the welded metalddc(4) Growth along the weld interface

Fig. 26. Classification of SCC morphologies in PLR pipe joints and definition of parameters evaluated.

30
25

Frequency

20
15
10
5
0
-6.0 -4.0 -2.00
2.0 4.0
Distance between the crack tip andthe weld interface ∆d (mm)Fig. 27. Frequency of the distance between crack tip and fusion boundary that shows clear peak at nearΔd = 0, which indicates that the tip of many cracks was just on the fusion boundary.

4.2.3. Crack growth evaluation considering the actual crack behavior
Estimated SCC growth behavior in PLR piping based on the examinations on actual components
discussed above is schematically shown in Fig. 28. A stress corrosion crack is supposed to: 1) initiate
after a certain period of incubation time, 2) grow at a high growth rate in the hardened areas of thebase material, 3) arrest or significantly slow down adjacent to the fusion boundary and 4) grow within

14

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

Lower growth ratedc

Crack depth d

the weld metal at a lower rate than that of the hardened base metal.
Although the time periods 1) and 3) have a large impact on the lifetime of actual components,
quantitative evaluations are difficult at this moment since both processes include probabilistic
uncertainties.
In order to take into account the difference in crack growth rates between periods 2) and 4), theJSME disposition curve for SUS304 in Fig. 23 that envelope the mock-up data was used while thecrack is growing within the hardened base metal (depth d < dc). The disposition curve for low-carbon
stainless steels was used once the crack was evaluated to be deeper than dc and to be propagating into
the weld metal, since the data from the weld metal of the PLR mock-up were enveloped by this curve.
In Fig. 29, the critical depth dc, measured on cross sections of actual PLR joints, is shown as afunction of L, the distance between the crack initiation site and the fusion boundary. As shown in thisfigure, a relatively good correlation was observed between dc and L. As L is measured by UT
inspections, the critical depth dc at which a detected crack reaches the welded materials can be
estimated from the relationship shown in Fig. 29.

dc④Weldedmetal②Hardened
area

0

Time

③ Arrest at
the boundary① Incubation
periodFig. 28. Crack behavior of SCC in PLR piping suggested by multiple flaw evaluations of actual
components.

Depth when the crack reaches
the interface dc (mm)16dc = dc + 2σ
= 1.0 L + 5.7

14

dc = 1.0 L + 3.0

12
+ 2σ

10
8

dc = dc - 2σ

6
4
dc

2
0
0.0

L
2.0

4.0

6.0

8.0

10.0

Distance between where the crackdeveloped and the fusion boundary L (mm)Fig. 29. Relationship between dc (crack depth when a crack reaches the fusion boundary) and L (distance
between the initiation site of the crack and fusion boundary on the inner surface).

4.2.4. Tentative Analysis and Future Problems
Considering the possible error in the measurement of L by UT, the 2σ-offset line of the best-fit
line of Fig. 29 is used in the tentative evaluations of the JSME code. Here, σ is the standard deviation
assuming the residual follows the normal distribution.
As an example, estimated results for initial cracks with 2 mm depth and 20 mm length located at
4.3 mm from the fusion boundary in 600A (main pipe) and 400A (header) pipes are shown in Fig. 30.
In both cases, it was assumed that the crack reaches the interface and growth rate changes at a depthof dc =10mm. As references, evaluation results, obtained using the disposition curve for SUS304 forthe entire period, are shown with dashed lines in Fig. 30.

15

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

In the 600A pipe, crack growth rate decreases and the crack slows down rapidly in both cases
due to the residual stress profile. In the case of 400A pipe, on the other hand, there is a difference
depending on the assumption of the crack growth rate and a significant difference between thesecalculated results and actual plant data.
25

30
25
Sensitized stainless→ Low carbon stainless

20
15

Crack depth d (mm)Crack depth d (mm)35

Sensitizedstainless

10
5
0
0

2

4
6
8
10
12
14
16
18
Service period (Effective Full Power Year)

20

20

Sensitized stainless→ Low carbon stainlessSensitizedstainless

15
10
Initial Crack Size ：Depth 2.0mm
Length 20mm

5
0

0

2

4
6
8
10
12
14
16
18
Service period (Effective Full Power Year)(a) Main Pipe (Φ600 mm)

20(b) Header Pipe (Φ400 mm)

Fig. 30. Examples of SCC growth evaluation using the provisional methodology that uses two different
disposition curves, one for hardened base metal and the other for weld metal.

4.3.

Examples of the Structural Integrity Assessment Based on Maintenance Codes

Based on the findings described above, structural integrity assessments were performed for 600A
and 400A piping. The initial crack size was determined based on the maximum crack size observed in
the actual PLR pipes with each diameter (Table 1). The extra margin of 4.4 mm (based on the double
standard deviation of sizing error in the UTS project) was added to the initial size, since the PD
system for enhanced UT techniques had not been established when the cracks were observed.
According to EB-4420 of JSME code, assessments based on elastic-plastic fracture mechanics
were carried out. Detailed procedures of the evaluation are described in Appendix E-9. They areapplicable to cracks with their depth less than 75 % of the wall thickness and their surface length lessthan 1/6 of the circumference. In case of PLR piping, the maximum allowable size is not prescribedby the fracture assessment but by this applicability range of the evaluation procedure.
Table 1 lists the results for the time for a crack to reach the maximum allowable size that wasevaluated based on the procedure described in the previous section. The JSME disposition curve forsensitized stainless steel was used for crack growth evaluation within the hardened area.
Table 1

Example of Crack Growth Evaluations and Fracture Assessment of PLR Piping

Depth (mm)Initial Crack Status
Length (mm)Time (year)
Crack Status when
Depth (mm)Crack Depth =75 % of Wall Thickness
Length (mm)Time (year)
Crack Status when
Depth (mm)Crack Length =1/6 of Circumference
Length (mm)Evaluation Period (year)
Acting Bending
Operation
Stress (MPa)
StatusFracture
Allowable Bending
A, B
Assessment
Stress (MPa)at the end of
Acting Bending
Evaluation
Operation
Stress (MPa)PeriodStatusAllowable Bending
C, D
Stress (MPa)

16

600A
13.6 (=9.9+4.4)
39.0
> 40
–
–
9.8
14.2
289
5

400A
11.2 (=6.8+4.4)
330
5.2
19.5
223
4.4
16.5
190
4.4

4.0

1.7

63.0

41.0

20.4

9.4

169

125

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

For both diameters, the limitation for crack length (<1/6 of the circumference) will first be met
after 9.8 and 4.4 years of continued operation in 600A and 400A pipes, respectively. The limitation
for crack depth (<75 % of wall thickness) will not be met at least the next 40 years for 600A pipe.
Fracture assessments were performed for the assumed crack size if the plant operation continued
for a certain “evaluation period”. The evaluation period for 400A pipe was set as 4.4 years from theremnant life before the crack length limitation will be met, whereas 5 years was conservatively used
for 600A pipes with remnant life of 9.8 years. Three different criteria for fracture assessment are
specified in Appendix E-9, and allowable bending stress was evaluated here as an example. The
applied loads considered include internal pressure, thermal loads, own weight, seismic load andwelding residual stresses. Welding residual stresses were evaluated using thermal elasto-plastic finite
element analysis and other loads were determined using the maximum value assumed in the design of
PLR piping with the same diameter among the Japanese BWRs.
As shown in the lower portion of Table 1, acting bending stress at the end of the evaluation
period after crack growth was evaluated to be lower than the limit values.
5.

Common and Specific Characteristics for SCC in Core Shroud and Recirculation Piping

5.1.

Common Characteristics for both Core Shrouds and PLR piping

5.1.1. Existence of the hardened layer on the surface• Hardened layer over 300HV with a thickness of a few hundred μm was observed on theas-machined surface at the edge of the shroud ring. On the shell of shrouds, hardening was foundonly at the thin layer on the surface with heavy grinder finishing after welding.
• A thin hardened layer was observed on the inner surfaces of recirculation piping. This surface
layer was assumed to be formed during machining to align the bottom faces of the groove.
Hardening was also found adjacent to the weld root due to plastic deformation during welding.
• The hardness of the inner surface near the weld root of PLR piping was lower due to annealingby the weld heat input.
5.1.2. Dominant effect of principal stress on the crack orientation• Surface crack orientation of SCC in low carbon stainless steel was strongly dependent on theprincipal stress direction, whereas the profile of the sensitized area has a dominant effect on SCC
orientation in SUS304.
• Unlike the SCC in SUS304, some cracks in low carbon stainless steels were not parallel to theweld joint due to the effect of the hoop stress in both core shrouds and PLR piping.
• Cracks initiated by residual stress due to grinding propagated in radial or random directions. Thisis because residual stress by grinding is significantly anisotropic and grinding direction is usually
arbitrary.
5.1.3. SCC mode in crack initiation and propagation• Initial propagation within the surface hardened layer occurred as TGSCC.
• Subsequent propagation occurred as IGSCC.
5.1.4. Non-sensitized material
No significant Cr depletion zone was observed in HAZ of core shrouds and PLR piping.
5.1.5. SCC propagation into the weld metal
In both core shrouds and PLR piping, the ferrite content of the dilution zone where crack
propagated into the weld metal was below 5 %, which may have increased the SCC susceptibility of
the weld metal.
5.2.

Specific aspects for SCC in core shrouds – Effect of Neutron Irradiation•

Noticeable material hardening was observed around H4 and H3 weld joints located at the center

17

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs•of core shrouds as neutron fluence increases.
Neutron irradiation promoted the formation of the Cr depletion zone at grain boundaries that wasobserved in SUS304L core shroud with fluence of 1.3×1025 n/m2 (>1MeV).

6.

SCC Mitigation Techniques

There are some SCC mitigation techniques available for core shrouds to reduce surface residual
stress such as laser peening [24], water jet peening [25] (Fig. 31) and surface polishing. Applying
these techniques after removing cracks by machining can also be effective.
For PLR piping, on the other hand, Corrosion Resistance Cladding (CRC), removal of surface
hardened layer, crack removal, and Induction Heating Stress Improvement (IHSI) [26] are applicable.
Heat Sink Welding (HSW) and Narrow Groove/Fine Line Welding (FLW) are also effective to reduce
welding residual stress.
As for environmental factors, Hydrogen Water Chemistry (HWC) and Noble Metal Chemical
Addition (NMCA) have been applied in multiple BWRs. Those techniques reduce the corrosion
potential by lowering the concentration of dissolved oxygen and hydrogen peroxide. The relationshipbetween the amount of hydrogen injection and the corrosion potential is shown in Fig. 32 [27].
1200
レーザパルス

Lens

Target material

Compressive stress

Residual stress (MPa)

PlasmaWater△ Before LP treatment○ After LP treatment

900
600
300
0
300
600

After LP treatment

900

Schematic of Laser Peening

1200
0

200

400

600

800

1000

Depth from the treated surface (mm)ECP in RPV drain

Threshold potential for SCC
development in Type 304 stainless steel

Radiation dose rate of
the main steam

Radiation dose rate

Corrosion potential ECP (mVSHE)

Fig. 31. Schematic of laser peening and its effect in reducing surface residual stress.

Dissolved hydrogen in the feed water (ppm)Fig. 32. Effect of dissolved hydrogen in the feed water on the corrosion potential and the radiation dose
rate of main steam.

A titanium oxide (TiO2) injection technique is now being developed [28]. Figures 33 and 34
show the principle of the technique that utilizes Cherenkov light emitted by radiation(gamma-ray)/water interaction within the reactor core. This is a revolutionary technique to prevent
corrosion by increasing the anodic current with photo-electric currents.
It is important to select a proper combination of these techniques for each situation.

18

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

Photo-electrode (anode) reaction

2H2O + O2

light
Light - Cherenkov(in reactor core)

Anode reaction

Cathode reaction(Corrosion reaction)

2H2O

2M2+

4OH-

4e-

- Thermal excitation(+ promotion by Hydrogen)4e-O2 + 4H+

4e-

Normal reaction
Corrosion mitigation by TiO2

Fig. 33. Principle of SCC mitigation technique by TiO2 injection.

Corrosion potential ECP (mVSHE)

0.1

Upper sample with TiO2
Lower sample with NO TiO2
GND

0
-0.1

Exposure oflower sample(with NO TiO2)-0.2
Exposure of upper sample(with TiO2) to the light

-0.3
-0.4

0

1

2

3

4

5

6

7

8

Time (hour)Fig. 34. Effect of TiO2 injection in lowering corrosion potential.

7.

Clarification of the SCC Mechanism

In addition to flaw assessments from the engineering viewpoint discussed above, the mechanistic
understanding of SCC on non-sensitized stainless steels is also desirable to ensure the integrity ofplant components. A research committee on SCC Mechanism (SCCM committee) had been organizedby the Japan Society of Corrosion Engineering. About 60 members are researchers and engineers infields such as corrosion, welding and metallurgy, from universities, public research institutions, BWR
and PWR plant fabricators, metal companies, chemical plant manufacturers and so forth. Research
program applications from diversified points of view were submitted by committee members andseveral programs were selected and funded by the board members of the committee. Accepted
research programs focused on root causes for 1) the higher susceptibility of hardened surface layers to
SCC initiation, 2) cracks to propagate as IGSCC, and 3) cracking in welded material. Latest results
are summarized below.
7.1.

Increase in SCC Susceptibility of Surface Hardened Layers

7.1.1. Formation of surface hardened layers [23]Counter boring prior to welding (machining on inner surface to align grooves) resulted in theextremely thin layer with hardness over 300HV (Fig. 21). Moreover, as can be seen in the example ofa main pipe with a diameter of 600 mm in the same figure, the base metal adjacent to the weld root
had noticeably been hardened due to heat contraction during welding, and SCC propagated within thishardened area.
7.1.2.

Very fine grain layer formed on the surface (Nippon Nuclear Fuel Development, Co., Japan19

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

Atomic Energy Agency, Tohoku University and National Institute for Materials Science)
Fine grain structure layer was found within a depth of several μm from the surface with typicalgrain size of tens to a few hundred nm [29-31] (Fig. 35, 36).
The oxidation observed in surface fine grain layers suggests that the decrease in corrosion
resistance promoted SCC initiation. Similar fine grain layer was also observed on surfaces with heavymachining following weld heat input, which suggests that recrystallization due to the combination of
surface machining and heat input is the root cause of the formation of the fine grain layer.
Although a microscopic magnetic layer (martensite phase) was also identified on the surface, it
was considered that the layer was not directly related to SCC initiation since the amount of themartensite phase is quite small.
The corrosion resistance of slip bands created by heavy cold work was also found to be low [32](Fig. 37). In oxygenated high temperature water, generally, a small amount of impurity ionsconcentrate at the crack tip. Thus corrosivity of the crack tip water chemistry becomes significantly
higher than the bulk water, which accelerates the corrosion along slip bands. Considering thesefindings, SCC is more likely to develop in a trans-granular manner since the number of slip lines isrelatively high within the surface layer where the effect of machining is significant.

Pipe inner surface
Fine grain layer- Grain size:
40 nm～- Layer thickness:
4 μm

8mm
Crack

Weld
metal

Pipe inner surface

Pipe Innersurface

Fine
grain
layer

Normal grain
layer- Grain size:
Few μm～Region ofinterest

Oxide
layer

Oxide
layer
1μm

Fig. 35. Very fine grain structure layer observed on the inner surface of the PLR pipe joint.
Magnetic reaction on the surface

1mm

TEM analysis (316NG)

MFM analysis (316L)

Formation of Fine grain layers by
Machining (Lathe) + Heat input

Formation of Martensite phase by
Machining (Milling) + Heat input

Fig. 36. Fine grain structure formed by combining surface machining and heat input simulating welding.

20

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

Measured path
Slip lines

Averaged surface profile

40nm←Shroud surface

Slip lines

Slip lines

Boat sample from the core shroud(Fukushima-dai2 Unit 2, H6a weld seam)

Machined test sample (Milling)

Fig. 37. AFM observation that showed a tendency of preferential corrosion along grain boundaries.

7.1.3. Tensile residual stress formation in surface hardened layers (Osaka University)
Residual stress at the surface can become higher when welding residual stress is added to theoriginal surface stress due to surface machining. Simulation work was conducted on the effect ofinteraction between surface machining and welding upon the surface residual stress and it wasevaluated that approximately 600 MPa of surface residual stress can be generated. (Fig. 38)

Fig. 38. Results of numerical simulations of the interaction effect between welding and machining on
residual stresses.

7.1.4. Threshold stress for TGSCC in surface hardened layers (Sumitomo Metals, Hyogo University
and Japan Atomic Energy Agency)
The surface residual stress measurement was conducted in the synchrotron radiation facility
Spring-8 in Hyogo, Japan. Four-point bend SCC initiation tests were also performed in a simulated
BWR environment. These tests revealed that as it goes deeper inside of the material, the residual
stress drops rapidly at a depth of a few to tens of μm and that the threshold surface stress for the
TGSCC initiation was 600 MPa [31,33] (Fig. 39). In solution-annealed materials with no surface
machining, no SCC was found even at a stress level above the yield stress. These tendencies that
surface hardening is necessary for SCC to occur agree well with SCC occurrence observed in actual
components.
From discussions above, it can be concluded that current preventive treatments such as surface
polishing, IHSI and narrow groove welding for PLR piping as well as surface polishing and peening
for core shrouds are effective from mechanistic viewpoints.

21

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

Surface-hardened plate

1200

Stress (MPa)

1000

1400

Applied stress
0MPa
150MPa310MPa460MPaStress on the surface，MPa

1400

800
600

Annealed plate

400
200
0
0

5

10 15 20 25 30 35 40 45 50: SCC: SCC/No SCC: No SCC

1200
1000
800

σth = 600 MPa

600
400
200
0

Observed depth (µm)

288℃，8ppmDO
SO42-: 200 ppb
1000h

Surfacehardened

Coldrolled

Annealed

Residual stress profile in
CBB test results
thickness direction
Fig. 39. Comparison between surface residual stress due to machining and tentative threshold stress for
SCC initiation obtained from CBB testing.

7.2.

Root Cause of IGSCC Growth in Non-sensitized Stainless SteelGrain boundary

Slip along a grain boundary(Strain rate: 1.25×10-4 /s at 288ºC)

Amount of slip along thegrain boundary (μm)

7.2.1. Effect of grain boundary deformation on IGSCC growth (Osaka University)
It is important to evaluate the resistance of the grain boundary to slip deformations duringloading when evaluating IGSCC growth. Tensile tests on SUS316 with cold work ratios of 0 %, 20 %,30 %, 40 %, and 60% were conducted at 288 ºC in vacuum at a strain rate of 1.25×10-4 /s. Differencesin mechanical properties between intra-grain and grain boundary were evaluated as a function of thecold work ratio.
Results showed that the slip at grain boundaries was small below 20 % cold work, but thematerial was more susceptible to deformations when cold work ratio exceeded 30 %. Vickers
hardness values for 20 and 30%-cold work materials are roughly 230 and 300, respectively. As shownin Fig. 40, the change in crack growth rate versus grain boundary slip/hardness and that versus grainboundary slip susceptibility show similar profiles [33]. This suggests that IGSCC growth rate is anincreasing function of the susceptibility to grain boundary slip.
Although it remains unclear whether this grain boundary slip triggers SCC or it occurs as a resultof SCC, it was at least confirmed that strain concentration occurred as applied load increased.
Considering this observation, higher strain energy due to strain accumulation caused oxide film
rupture and caused IGSCC growth.
1.4
1.2
1
0.8

Cold work reduction ratio: 0% (As received)
: 20%
: 30%
: 40%
: 60%

Hardness (HV)
168
236
319
376
409

0.06

0.1

0.6
0.4
0.2
0
0

0.02

0.08

StrainFig. 40. Relationship between the amount of grain boundary slip and cold work ratio.

22

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

7.2.2. Detailed observation on the inside of stress corrosion cracks in actual components
The inside and the tip of stress corrosion cracks taken from actual components were examinedby Transmission Electron Microscope (TEM) by PNNL (Pacific Northwest National Lab.) as a
TEPCO program. Cr-rich oxide was found at the crack tip and Ni enrichment as well as Cr and Fe
depletion was observed in front of crack tips as shown in Fig. 41.
It was supposed from the above observations that the environment at the crack tip was highly
corrosive with low corrosion potential and pH, which is quite similar condition that is assumed in
SCC in sensitized SUS304. The segregation of Ni, Cr and Fe in front of the crack tip can also imply
that the crack was growing at lower rate or even arrested.
From the findings described above, it seemed that similar phenomena were taking place at the
tip of SCC in both SUS304 and low carbon stainless steels.
When SCC testing on materials identical to actual components is conducted in an accurately
simulated BWR environment, oxide film of a nature identical to that of the actual component is
supposed to be observed on crack faces. In typical laboratory testing, however, segregation of alloy
elements ahead of the crack tip is less likely to occur, probably due to effects of the deformation
taking place adjacent to the crack tip because of the intentionally sustained crack growth. This implies
that the crack growth rate results from laboratory tests conservatively evaluate SCC growth behavior
in actual plants.

Scan1

Scan2

tip
scan2
scan1

Fig. 41. TEM observation of the inside and the tip of a crack taken from PLR pipe joint of
Kashiwazaki-Kariwa Unit 1.

7.3.

Summary of SCC Mechanisms

Through the research programs of the SCCM committee, new insights into the mechanism of
SCC initiation and propagation were obtained. These new findings are summarized as follows (Fig.
42).
(1) It was experimentally confirmed that surface machining and welding enhances SCC
susceptibility. These can be another accelerating factors of SCC initiation and propagation
besides sensitization.
(2) Recently performed SCC growth tests seem to be considering SCC promoting factors or crack
conditions properly or conservatively, therefore, SCC growth rate disposition curves in JSME
code are supposed to be adequate to be applied.
(3) The efficiency of current preventive treatments such as stress reduction, surface polishing andenvironmental mitigation has also been validated microscopically.
(4) A more quantitative understanding of parameters, such as strain, hardness, threshold stress,
grain boundary slip and degradation in corrosion resistance, were identified to be important inthis program. Further investigation for other possible factors affecting SCC of low carbon
stainless steels mechanistically should be focused in future studies.

23

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs
Fine grain layer(lower corrosion
resistance)

Surfacemachining

Formation of finegrain (hardened)surface layer

Residual stress

Surface(Coolant watercontact)
Oxide

Residual stresson the surface
Welding

Slip
line
すべり線
(Lower corrosion
resistance)
Crack growth due tocorrosion reactions

Hardening
due to plastic
deformation
Welding residual
stress(>600MPa (tentative
threshold for TGSCC))

Higher susceptibility to slips
along the grain boundary
due to higher cold work ratio

Fig. 42. Summary of the latest findings from research programs focused upon mechanistic understandingof SCC in low carbon stainless steels.

8.

The Necessity for Rationalization of the SCC Growth Evaluation Methodology

Due to technical problems remaining in evaluations of SCC such as (1) larger crack sizing error
by UT and (2) shortage of data to consider the hardening effect on crack growth rate quantitatively,
current evaluation methodology according to the JSME code case requires highly conservative
assumptions in each step of the evaluation. As a result, current inspection intervals seems to be
unreasonably short considering the actual trend of SCC in core shrouds and PLR piping described in
Sections 3 and 4. Establishing more reasonable evaluation methodology based on objective andtechnical background is, therefore, an urgent task for Japanese BWR owners group.
In order to accumulate data to reduce uncertainties related to the item (2) above, detailed
examinations on actual PLR pipe welds removed for replacement were conducted. The summary andfuture works are as follows.
8.1.

Summary of SCC Evaluations on Recirculation Piping

Conservative assumptions required in the current SCC evaluation for PLR piping are:(1) Possible error in crack sizing by UT: As measured crack depth + 4.4 mm is used as initial
crack depth in flow assessments.
(2) Crack modeling: A partial crack is modeled conservatively as a through-wall crack with thesame surface length. Discrete cracks over entire circumference are modeled conservatively
as a circumferential crack.
(3) Calculation of K: Formula for a crack in a flat plate is used for cracks in a cylinder.
(4) SCC growth rate: Disposition line is drawn so that it envelops all the data ever obtained fromlaboratory tests.
(5) Crack initiation: A crack is assumed to be present from the beginning of the plant operationneglecting the incubation period for SCC initiation.
Although current crack growth evaluation according to the maintenance code predicts that SCC
would grow rapidly and continuously, the majority of the actual cracks are no deeper than 10 mm.
Possible reasons for this discrepancy are: (1) many cracks reached and arrested at the weld interface attheir depth of around 10 mm, and/or (2) stress intensity factor can decrease as crack grows dependingon the residual stress profile, which may result in crack deceleration (negative dK/da effect).
Further rationalization can be attained by taking the incubation period before the SCC initiation or
possible crack arrest at the interface into consideration in the crack growth evaluation procedure (Fig.
28).
Applying hydrogen water chemistry as the mitigation technique is beneficial for its effect not only
on crack growth rate reduction but also on the extension of the incubation period before SCC
initiation. Establishing more reasonable methodology for lifetime evaluation taking such mitigation
effects into account properly is desirable (Fig. 43).

24

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs
NWC

Crack size

(Normal Water
Chemistry)

Crack behavior
under NWC

ＳＣＣ deceleration by
applying HWC

Crack behavior
under HWC

HWC

(Hydrogen Water
Chemistry)

0

Extension of incubation period
by applying HWC

Incubation period
under NWC

ServiceperiodFig. 43. Consideration of beneficial effects of HWC in SCC growth evaluation.

8.2.

Remaining Tasks for Lifetime Evaluation with Higher Accuracy

Stress intensity
factor K

Residual stress8.2.1. SCC growth behavior under constant displacement conditions
The change in the stress intensity factor
Configuration of thewith crack growth differs depending on thegroovewelding residual stress profile. Since most of thecrack growth rate data are obtained from testingunder increasing K condition as crack grows,
Crack depth aevaluations based on such data can be tooconservative when cracks are growing underdecreasing K condition. And such condition＋＋
０
isn’t uncommon in actual components－
depending on residual stress profile (Fig. 44).
FRI model has been developed by Tohoku
Increasing K (Negative dK/da)
University to evaluate SCC growth rate basedHigher crack driving force
on mechanistic ground [34]. The distinctive
for a same K value (?)
feature of this model is that it considers effects
Decreasing K (Negative dK/da)
Lower crack driving force
of crack tip strain rate on SCC growth rate
for a same K value (?)
quantitatively. From this model, it is predicted
Crack depth athat crack growth rate drops significantly evenat relatively high K when K decreases as theFig. 44. Effect of residual stress profile on thecrack grows. This prediction is consistent withchange in K as crack growth and possible effect ofexperiment results under constant displacement
differentiation of K on SCC growth rate.
conditions as well as the actual crack behavior
observed in core shrouds (Fig. 45).
70

60

60

50

50

40
30
20

Crack Depth (mm)70

Maximum depth (by UT)
Average depth (by UT)

40
30
20

10

10

0
70 60 50 40 30 20 10 0
K （MPa√m）

0

Average：15.5mm
Average: 8.6mm
Plant A
Plant C
Plant B
Plant D
(Since 1987) (Since 1990) (Since 1993) (Since 1993)

Fig. 45. Change in K as a circumferential crack grows at H7 weld of core shrouds in comparison with theaverage crack depth of SCC observed in H7 welds of 4 Japanese BWRs.

25

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

8.2.2. Effect of material hardening on SCC growth rate [35]Since degree of material hardening adjacent to a crack differs depending on various factors suchas pipe diameter, weld condition, crack location, crack path and so forth, applying only one growthrate disposition curve determined with conservative assumptions to all the cracks regardless of theabove factors sometimes provides unreasonable prediction of crack behavior.
1. Hardness measurement on actual PLR pipe welds
Quantitative evaluation of hardening effect on SCC growth rate requires (1) the degree ofmaterial hardening in the specific weld joint and (2) the quantitative relationship between the degreeof hardening and SCC growth rate.
To obtain an insight into the item (1), detailed examinations were performed on 38 PLR pipe
weld joints removed from actual BWR plants for replacement. The primary objective was tosystematically relate hardness profiles and pipe configurations/weld conditions.
Vickers hardness (HV), yield stress and plastic strain by EBSD (Electron Back-Scatter
Diffraction pattern, by GE Global Research as a TEPCO program) measurement were used asmeasures of the degree of hardening. Small tensile test bars were taken from locations with different
HV and tensile tests were performed in air at room temperature.
As can be seen in Fig. 46, yield stress and tensile strength become higher and elongation
becomes lower as it goes closer to the root of the weld.
Hardness measurement was also performed at the test section of each specimen. It was
confirmed that hardness and yield stress show clear proportional relationship as shown in Fig. 47.
2. Relationship between Plastic Strain and Hardness
Figure 48 shows plastic strain profile near the inner-surface of PLR piping measured by EBSD.
Plastic strain is higher near the weld fusion lines and differences between manufacturers are
somewhat unclear.
Differences in plastic strain profile by pipe diameters (welded by the same manufacturer) are
shown in Fig. 49. This figure indicates that the difference in hardening degree by pipe diameter was
significant and that plastic strain increase as pipe diameter increases. From this result, it is understoodthat the present assessments procedure that uses crack growth rate disposition curve for significantly
hardened material regardless the pipe diameter can be excessively conservative for pipes with smaller
diameters.

T16-6

T16-5

T16-4

T16-2

Weld metal
Base metal

T16-3

Weld
Weld metal
metal
Base
Base metal
metal

T16-3

T16-4

T16-5

T16-6

σ0.2 : 645 MPa
800 σB : 740 MPa

σ0.2: 595 MPa
σB : 720 MPa

σ0.2 : 548 MPa
σB : 696 MPa

σ0.2: 467 MPaσB : 663 MPa

σ0.2(0.2% Proof Stress): 384 MPa
σB (Tensile Strength): 629 MPa

Stress （MPa)1000 T16-2

600
400
200
0
0

2

4 0

2

4 0

2

4 0

Displacement （mm)

2

4

0

2

4

6

Fig. 46. Difference in tensile properties between locations adjacent to the weld root of a PLR pipe joint
removed from Kashiwazaki-Kariwa Unit 1 for replacement.

26

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

0.2% Proof Stress σ0.2 (MPa)

800
700
600
500
400
σ0.2 = 2.93 HV - 147

300
200
100
0

0

50

100

150

200

250

300

Vickers Hardness HV

Fig. 47. Relationship between Vickers hardness and 0.2 % proof stress measured on a PLR pipe joints
removed from Japanese BWRs for replacement.

30

No.1

25

Plastic strain (%)

No.3
No.2
No.7
No.6

Scan C

Scan A

20
15
10
5

No.5
Scan C

Fabricator A, F316
Fabricator B, 316TP

Fabricator B, F316
Fabricator A, Cast

Scan A

0
-25

-20

-15

-10

-5

0

5

10

15

20

25

Distance from the fusion boundary (mm)Fig. 48. Difference in plastic strain profiles near PLR pipe joints (600A) between fabricators.

30

No.1
Φ600 mm

Φ600 mm

No.2

No.4

Φ600 mm
Φ400 mm
Φ300 mmScan A

Scan C

Plastic strain (%)

No.3

25

Φ600 mm
Φ400 mm
Φ300 mm201510No.8 Φ400 mm No.9 Φ300 mm
5Scan C

Scan A

0
-25

-20

-15

-10

-5

0

5

10

15

20

25

Distance from the fusion boundary (mm)Fig. 49. Difference in plastic strain profiles near PLR pipe joints (Fabricator A) between pipe diameters.

8.2.3. Application of the latest findings to the maintenance code
As discussed above, pipe diameter and crack paths within the hardened area of base metal are nottaken into account in the present provisional maintenance code and SCC growth rate for sensitizedstainless steel is used in all cases. An extensive SCC growth rate measurement was conducted by
JNES (Japan Nuclear Energy Safety Organization) using specimens taken from hardened base metal

27

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

and weld metal of full size mock-up of PLR piping [36]. According to their data, the SCC growthrate for sensitized stainless steel is nearly twice as that obtained from a 600A mock-up specimen.
Since the degree of hardening of 400A and 300A pipes is even lower, there remains a possibility thatthe current SCC growth evaluations for medium-diameter pipes become too conservative.
Tentative SCC growth analyses were performed using the SCC disposition curve based on JNES
data. The K calculation formula for circumferential crack in a cylinder was incorporated in the revised
JSME code as of 2004 and was used in these tentative evaluations. Results are summarized in Fig. 50.
Large discrepancies in crack growth behavior were observed, especially in the case of 300A pipeshown in Fig. 50(b) depending on the crack growth disposition curve and K formula. As clearly beseen in this figure, current evaluations assuming a crack in flat plates are noticeably conservative.
35

20

25

②

③

Crack depth (mm)Crack depth (mm)Pipe diameter: 600mm

Pipe diameter: 600mm①
dc=10mm

30

④

20
15
10
5
0

Semi-elliptical Semi-ellipticalin flat plates in cylinders

K formulaSensitized stainless
HAZ of low carbon stainless*

0

50

100

150

①
②

③
④

200

250

Time （year）

dc=10mm

15

②

①
③

④

10
5
0

Semi-elliptical Semi-ellipticalin flat plates in cylinders

K formulaSensitized stainless
HAZ of low carbon stainless*

0

5

10

15

①
②

③
④

20

25

Time （year）(a) 600A Pipe(b) 300A Pipe
Fig. 50. Effect of K-formula and the SCC growth rate disposition curve on SCC growth prediction.

9.

Future Topics that Should be Covered

The following is a summary of future tasks to be investigated and data necessary to beaccumulated in order to develop more reasonable structural integrity assessment procedure.
(1) SCC growth rate data for cracks within welded metals and HAZ(2) SCC growth rate data for irradiated material(3) Threshold stress on the surface with machined layer(4) Necessary conditions for SCC initiation in weld metals(5) Further mechanistic understanding on SCC initiation and propagation (materials, stress andenvironment).
For items (1) to (4), consideration of the effect of water chemistry is also necessary.
Item (1) is especially important to rationalize the structural integrity assessment procedure. Some data
will be obtained from research projects conducted by Japanese utilities within the short term, andmore extensive testing will be performed through national projects assuming the application of thelatest data and findings to the code.
It is expected that the national projects on IASCC will provide valuable data related to the item
(2).
Items (3) and (4) will be evaluated through international joint research projects between Japaneseutilities and foreign research institutes.
More contributions from universities, national research institutes and specialists in non-nuclear
industries are expected, especially in research programs related to item (5). In order to determine amore reasonable interval of inspections of SCC defects, establishing a quantitative procedure for
evaluation of the SCC incubation period, considering its probabilistic nature, is desirable.

Increased effort will be made to promote the industry-government-academia researchcollaboration to obtain high quality and state-of-the-art data and to implement them to themaintenance code and standard promptly after open disclosures.

28

S. Suzuki et al./ Stress Corrosion Cracking in Low
Carbon Stainless Steel Components in BWRs

References[1] Okamura et al., Trans. of SmiRT 17(2003)[2] Cheng, C. F., Ellingson, W. A., and Park, J. Y, Corrosion/76, (1976)[3] Nuclear Power Engineering Test Center, Japan Power Plant Institute; “Present Nuclear Power
Facility Reliability Verification Testing” (1984)[4] Japanese BWR owners group research data
[5] M.Tsubota et al., 7th international symposium on Environmental Degradation of Materials in
Nuclear Power Systems - Water Reactors, Vol.1(1995)[6] TEPCO research data
[7] Ito et al., Proc. 48th Japan Conf. Materials and Environments, JSCE (2001), pp.103-106[8] JSME, Codes for Nuclear Power Generation Facilities –Rules on Fitness-for-Service for Nuclear
Power Plants -, JSME S NA1-2002[9] Katayama, Proc. 46th Japan Conf. Materials and Environments, JSCE (1999), pp.163-166[10] Suzuki et al., Maintenology Vol.3, No2 (2004), pp.59-64[11] Suzuki et al., J. of High Pressure Institute of Japan (JHPI), Vol.42, No4 (2004), pp.13-22[12] API 579, American Petroleum Institute, “Recommended Practice 579, Fitness for Service”(2000)[13] Advisory Committee for Natural Resources and Energy, Nuclear and Industrial Safety
Subcommittee (4th) Materials[14] Tokyo Electric Power Company, Press Release, Jul. 10, 2008[15] Tokyo Electric Power Company, Press Release, Jul. 31, 2008[16] D.P.Rooke and D.J.Cartwright, "Stress Intensity Factors”, Sept.1974[17] Suzuki et al., Maintenology Vol.3, No.2 (2004), pp.65-70[18] Japan Power Engineering and Inspection Corporation, Inspection Techniques for Nuclear Power
Facilities/Verification Testing for Flaw Detection and Sizing Accuracy of Ultrasonic Testing,2002 Report[19] Advisory Committee on Energy and Natural Resources Nuclear Safety and Peace Preservation
Subcommittee Nuclear Power Facility Structure Health Evaluation Committee (8th), Japan
Power Engineering and Inspection Corporation Confirmation Testing Related to Improving
Recirculation Pipe Sizing Accuracy for Ultrasonic Testing[20] Andresen et al., Corrosion, Paper 02511(2002)[21] Andresen et al., Proc. 11th international symposium on Environmental Degradation of Materials
in Nuclear Power Systems - Water Reactors (2003), pp.816-833[22] Jenssen et al., Proc. 7th international symposium on Environmental Degradation of Materials in
Nuclear Power Systems - Water Reactors (1995), pp.553-562[23] Kumagai et al., Proc. ASME PVP2004, Vol.479 (2004), pp.217-223[24] Sano et al., Toshiba Review, 53-10(1998), pp.49-52[25] Saito et al., J. of Water Jet Technology Society of Japan, 20-1 (2003), pp.4-12[26] Ishikawajima-Harima engineering review, 18-1 (1978), pp.29-37[27] Atomic Energy Society of Japan, Reactor Core Chemistry Handbook[28] Takamori et al., “DEVELOPMENT OF BWR COMPONENTS SCC MITIGATION METHODBY THE TIO2 TREATING TECHNIQUE”, Proc. 13th international symposium on
Environmental Degradation of Materials in Nuclear Power Systems - Water Reactors (2007)[29] Kitsunai and Kodama, Proc. 52nd Japan Conf. Materials and Environments, JSCE (2005), B-212[30] Ueno et al., Proc. 52nd Japan Conf. Materials and Environments, JSCE (2005), B-210[31] Takeda et al., Proc. 52nd Japan Conf. Materials and Environments, JSCE (2005), B-205[32] Nemoto et al., Proc. 52nd Japan Conf. Materials and Environments, JSCE (2005), B-211[33] Nishimoto et al., Proc. 52nd Japan Conf. Materials and Environments, JSCE (2005), B-207[34] Shoji et al., 7th international symposium on Environmental Degradation of Materials in Nuclear
Power Systems - Water Reactors (1995), pp.881-889[35] Suzuki et al., Japan Society of Maintenology 2nd Academic Lecture Presentations, 2005[36] Ando et al., “CGR Behavior of Low carbon Stainless Steel of Hardened Heat Affected Zone in
PLR Piping weld Joints”, Proc. 13th international symposium on Environmental Degradation ofMaterials in Nuclear Power Systems - Water Reactors (2007)29